XX/2000 MassMin 2000 MassMin 2000 Proceedings ABN 59 836 002 494 BRISBANE, Australia - 29 October to 2 November 200
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XX/2000 MassMin 2000
MassMin
2000
Proceedings
ABN 59 836 002 494
BRISBANE, Australia - 29 October to 2 November 2000
MassMin 2000 29 October - 2 November 2000 Brisbane, Queensland
The Australasian Institute of Mining and Metallurgy Publication Series No 7/2000
Published by THE AUSTRALASIAN INSTITUTE OF MINING AND METALLURGY Level 3, 15 - 31 Pelham Street, Carlton Victoria 3053 Australia
© The Australasian Institute of Mining and Metallurgy 2000
All papers (abstracts and full papers) published in this volume were refereed prior to publication.
The Institute is not responsible as a body for the facts and opinions advanced in any of its publications.
ISBN 1 875776 79 6
Desktop published by: Donna Edwards, Penelope Griffiths, Angie Spry and Miriam Tankey for The Australasian Institute of Mining and Metallurgy
Printed by: New Generation Print and Copy 12 Barkly Street Brunswick East Vic 3057
Organising Committee Dr Gideon Chitombo (Conference Chairman) Professor Ted Brown (International Advisory Committee Chairman) Libby Hill (Conference Secretary)
Advisory Members Ken Owen (Anglogold South America) Alan Guest (De Beers) Joe Luxford (Luxford Mine Management) Jock Cunningham (CSIRO) The Late Dr David Dekker (CSIRO) Professor Ernesto Villaescusa (WASM) Dr Robert Trueman (JKMRC) Italo Onederra (JKMRC) David Goeldner (JKMRC) Naomi Mason (JKMRC) John Collins (The AusIMM Southern Queensland Branch) The AusIMM Southern Queensland Branch Executive Committee
The AusIMM Miriam Way (Events Manager) Alison McKenzie (Events Secretary) Penelope Griffiths (Publications Manager) Angie Spry (Publications Assistant)
Reviewers We would like to thank the following mining professionals for their contribution towards enhancing the high-quality of papers included in this volume. Bryan Beighton
Brian Hall
Don Runge
Ian Bell
Glen Heslop
Frank Russell
Martyn Bloss
Michael Hood
Mike Sandy
Barry Brady
Bill Hustrulid
Gordon Smith
Rick Brake
Jarek Jakubec
Martin Smith
E T Brown
Robin Kear
Alex Spathis
Geoff Bull
Dave Landriault
Sam Spearing
Richard Butcher
Peter Lilly
Dick Stacey
Emmanuel Chanda
Andrew Logan
Dan Stewart
Ian Clark
Loren Lorig
Mike Struthers
Kit Clifford
Joe Luxford
John Summers
Jock Cunningham
Hans Mulhaus
Alan Thompson
Stephen Duffield
Simon Nickson
Bob Trueman
Gavin Ferguson
Dave Ortlepp
Mike Turner
Russell Frith
Dan O’Toole
Andre van As
Peter Gash
Ken Owen
Mark van Leuven
Stewart Gillies
Chris Page
Dave Wilson
Ian Gipps
Yves Potvin
David Winborne
Tony Grice
Kevin Rosengren
Chris Windsor
Alan Guest
Kevin Ross
Brian White
Foreword MassMin 2000 follows on from two successful mass mining conferences staged in Denver Colorado in 1981 and Johannesburg, South Africa in 1992. This series of conferences has since been termed ‘MassMin’. Arguably they have become the ‘pinnacle’ international forum for discussing and sharing both technical and operational issues and experiences associated with application of methods such as; caving (block, panel, sublevel) and large-scale stoping (sublevel and longhole), including their derivatives (underground benching, front cave, inclined footwall, etc). The 102 papers in these MassMin 2000 proceedings effectively reflect the significant interest, developments, advances and experiences gained in underground mass mining since the MassMin 92 event. These proceedings thus provide the ‘state-of-the-art’ in underground mass mining. Papers have been written on some of the worlds largest current and future underground mining operations and projects, eg CODELCO Divisions (Chile); De Beers and Palabora (South Africa); PT Freeport Indonesia (Indonesia); Philex (Philippines); Tongkuangyu (China); Henderson and Bingham (USA); Brunswick and the Hudson Bay 777 project (Canada); LKAB Kiruna (Sweden); Shabanie and Trojan Mines (Zimbabwe) and indeed Northparkes, Olympic Dam, BHP Cannington, Perseverance Leinster Nickel Operation (Australia). Unique mass mining experiences in Russia, Zambia and Ghana are also discussed. Papers on novel enabling technologies (eg equipment, explosive products, mine planning and production management systems) designed to make underground mass mining more efficient, productive and safe are included. The MassMin series of conferences add real value to our industry. Technical contributions in the form of papers and financial support in the form of sponsorship is indeed a true investment towards our collective mining future and know-how. As such all authors (and their respective companies) and sponsors (major, minor and exhibitors) of this MassMin 2000 conference are gratefully acknowledged. Finally, the Director and Administration of the Julius Kruttschnitt Mineral Research Centre are gratefully acknowledged for allowing me the time and allocating me with the administrative support in the form of Libby Hill to organise and chair this event. Contributions from members of the CSIRO Mining and Exploration Division, Brisbane are also acknowledged. Finally the contributions of The Australasian Institute of Mining and Metallurgy (the Events Department – Miriam Way and Alison Mackenzie and the Publications Department – Penelope Griffiths and Angie Spry) towards the organisation and promotion of the MassMin 2000 were significant and are gratefully acknowledged.
Dr Gideon Chitombo MassMin 2000 Conference Chairman
Proceedings Sponsors
Block Caving Study - Stage 1 ICS The International November 1997 - November 2000
De Beers Consolidated Mines Limited CODELCO Chile Itasca Consulting Group Inc. JKMRC Newcrest Mining Limited Noranda Inc. Northparkes Mines PT Freeport Indonesia Rio Tinto Limited TVX Gold Inc.
The SPONSORS MassMin 2000 Proudly Sponsored by:
Major Sponsors
Sponsors
Arnall Mining Products Division
Welcoming Cocktail Party
Official Conference Dinner
Official Conference Satchels
Technical Session
Technical Session
Conference Pads and Pens
Technical Session
Technical Session
Caving Workshop Sponsors
NEWCREST MINING
The Exhibitors Australia’s Mining Monthly Dyno Nobel Asia Pacific ORICA Explosives Brandfield Concept Engineering Gemcom Australia Pty Ltd Runge Pty Ltd Caterpillar Elphinstone Pty Ltd ISS Pacific Pty Ltd Sandvik Tamrock CMTE JKMRC WASM University of Queensland – Foots Institute
Official Publications National
International
Australia’s Mining Monthly
MINING Magazine
Contents Mass Mining Design Methodology Orebody Delineation and Reserve Estimation — Seeking Common Sense from Extensive Investment in Fact
J Marek
3
Estimation of Resources and Conversion to Reserves — Protocols for the Assessment, Reduction and Management of Risk
J Beniscelli, P Carrasco, P P A Dowd, G Ferguson and E Tulcanaza
9
Characterising the Mining Environment for Underground Mass Mining
E T Brown and K J Rosengren
17
Method Selection for Large-Scale Underground Mining
W Hustrulid
29
The Use of Evaluation Surfaces to Assist in the Determination of Mine Design Criteria
R M Kear
57
Analysis and Management of Mining Risk
J Summers
63
Simulation Modelling of Mining Systems
B E Hall
83
Open Pit to Underground — Transition and Interaction
T R Stacey and P J Terbrugge
97
Quantifying Geotechnical Risk in the Mine Planning Process
T P Horsley and T P Medhurst
105
Dilution Control in Southern African Mines
R J Butcher
113
Reflections of a Mine Scheduler
J Luxford
119
Optimising Mobile Equipment Resources in Massive Mining
T Puhakka
129
Intelligent Mine Technology Program and its Implementation
J Pukkila and P Särkkä
135
Use of OPTIMINETM Simulation Tool for Mobile Fleet Selection
T Puhakka and V Kainulainen
145
Planning and Scheduling
Equipment
Drilling and Blasting Designing and Delivering Explosives Systems and Solutions for the Underground Massive Mining Industry: The Next Five Years with Dyno Nobel
W R Adamson
151
Digital Blasting — An Opportunity to Revolutionise Mass Underground Mining
D Kay
155
Enabling and Potential Technologies Hydraulic Fracturing as a Cave Inducement Technique at Northparkes Mines
A van As and R G Jeffrey
165
Tele-Operation at Freeport to Reduce Wet Muck Hazards
G Hubert, S Dirdjosuwondo, R Plaisance and L Thomas
173
In Situ Stress Measurements Using Oriented Core
E Villaescusa, M Seto and G Baird
181
Terratec Universal Raiseborer — Raisebore Machine Technology
C Paterson
187
The Potential Use of Foam Technology in Underground Backfilling and Surface Tailings Disposal
A J S Spearing, D Millette and F Gay
193
The Design, Testing and Application of Ground Support Membranes for Use in Underground Mines
A J S Spearing and J Champa
199
Remote Monitoring of Rock Mass Deformation During Mining
M T Gladwin, R L Gwyther and M Mee
209
The Palabora Underground Mine Project
K Calder, P Townsend and F Russell
219
Cave Mining at Premier Diamond Mine
P J Bartlett and A Croll
227
Panel Caving Experiences and Macrotrench (Macrozanja) — An Alternative Exploitation Method at the El Teniente Mine, Codelco, Chile
G Diaz and P Tobar
235
Evolution in Panel Caving Undercutting and Drawbell Excavation, El Teniente Mine
J Jofre, P Yáñez and G Ferguson
249
The Pre-Undercut Caving Method at the El Teniente Mine, Codelco Chile
E Rojas, R Molina, A Bonani and H Constanzo
261
Esmeralda Mine Exploitation Project
M Barraza and P Crorkan
267
Block and Panel Caving
‘An Underground Air Blast’ — Codelco Chile - Division Salvador
R De Nicola Escobar and M Fishwick Tapia
279
Freeport Indonesia’s Deep Ore Zone Mine
J Barber, L Thomas and T Casten
289
Excavation Design and Ground Support of the Gyratory Crusher Installation at the DOZ Mine, PT Freeport Indonesia
T Casten, R Golden, A Mulyadi and J Barber
295
Commissioning of Two 750 kW Centrifugal Fans at PT Freeport Indonesia’s Deep Ore Zone Mine
F Calizaya, T Casten and K Karmawan
301
A Case History of the Crusher Level Development at Henderson
M F Callahan, K W Keskimaki and W D Rech
307
The New Henderson Mine Truck Haulage System — The Last Step to a Totally Trackless Mine
W D Tyler, K W Keskimaki and D S Stewart
317
Application of Block Caving System in the Tongkuangyu Copper Mine
Zhou Aimin and Song Yongxue
325
Block Cave Design and Geotechnical Considerations Design of the Second Block Cave at Northparkes E26 Mine
S Duffield
335
Modelling and Design of Block Caving at Bingham Canyon
C J Carter and F M Russell
347
Considerations for Design of Production Level Drawpoint Layouts for a Deep Block Cave
A R Leach, K Naidoo and P Bartlett
357
Noranda’s Approach to Evaluating a Competent Deposit for Caving
S Nickson, A Coulson and J Hussey
367
The Past Focuses Support for the Future
A D Wilson
385
Rock Mechanics as Applied in Philex Block Cave Operations
R S Dolipas
395
Block Cave Undercutting — Aims, Strategies, Methods and Management
R J Butcher
405
The MRMR Rock Mass Rating Classification System in Mining Practice
J Jakubec and D H Laubscher
413
The Role of Mass Concrete in Soft Rock Block Cave Mines
R J Butcher
423
Meeting Geotechnical Challenges — A Key to Success for Block Caving Mines
Dianmin Chen
429
Block Caving — Controllable Risks and Fatal Flaws
T G Heslop
437
Draw Control in Block Caving Forecast of Ore Recovery in the Bulk Caving System
Y V Kuzmin
457
An Application of Linear Programming for Block Cave Draw Control
A R Guest, G J van Hout, A von Johannides and L F Scheepers
461
PC-BC: A Block Cave Design and Draw Control System
T Diering
469
Draw Control at Premier Mine
P J Bartlett and K Nesbitt
485
An Update on Cave Development and Draw Control at the Henderson Mine
W Rech, K W Keskimaki and D S Stewart
495
The Evolution of Sublevel Caving at Trojan Mine, Bindura, Zimbabwe
J G Taylor and R Chinamatira
509
Sublevel Cave Drop Down Strategy at Perseverance Mine, Leinster Nickel Operations
P L Wood, P A Jenkins and I W O Jones
517
Footwall Stability at the LKAB’s Kiruna Sublevel Caving Operation, Sweden
E Henry and C Dahnér-Lindqvist
527
Sublevel Caving — Today’s Dependable Low-Cost ‘Ore Factory’
G Bull and C H Page
537
Gravity Flow of Broken Rock — What is Known and Unknown
A Rustan
557
Drawpoint Design in Caving and Stoping Mines
F O Otuonye
569
Sublevel Caving
Sublevel Caving Design
Sublevel and Longhole Stoping A Review of Sublevel Stoping
E Villaescusa
577
Mount Isa Mines — 1100 Orebody, 35 Years On
D Grant and S DeKruijff
591
George Fisher Mine — Feasibility and Construction
L B Neindorf and G S B Karunatillake
601
The Ventilation and Refrigeration Design for Australia’s Deepest and Hottest Underground Operation — the Enterprise Mine
R Brake and B Fulker
611
Bulk Low-Grade Mining at Mount Charlotte Mine
P A Mikula and M F Lee
623
The Mine Management System at Olympic Dam Mine
S Youds
637
Materials Handling at Olympic Dam Mine — Olympic Dam Aiming for ‘Gold’
P Bowman
645
The Automation of Western Mining Corporation’s Olympic Dam Underground Rail Haulage System
R Doubleday and D Mee
649
Open Stope Design at Normandy Golden Grove Operations
T M Calvert, J B Simpson and M P Sandy
653
Open Stope Mining in Canada
Y Potvin and M Hudyma
661
Open Stope Mining Strategies at Brunswick Mine
B Simser and P Andrieux
675
Evolution of Vertical Crater Retreat Mining at Mindola Mine, Zambia
E K Chanda and C Katonga
685
An Experimental Study on Large-Diameter Longhole Mining of High Stope at Anqing Copper Mine
Wang Renfa, Xia Qian and Jiang Zhiming
697
Impact of Stope Geometry on Backfill Systems for Bulk Mining
M Dorricott and A Grice
705
Cannington Paste Fill System — Achieving Demand Capacity
M Bloss and M Revell
713
Engineering the New Olympic Dam Backfill System
G Baldwin and A G Grice
721
Backfill Research at the Western Australian School of Mines
C Wang and E Villaescusa
735
Backfill Systems
Design and Geotechnical Consideration Geotechnical Aspects of Open Stope Design at BHP Cannington
G C Streeton
747
Rock Mechanics Design and Practice for Squeezing Ground and High Stress Conditions at Perseverance Mine
M A Struthers, M H Turner, K McNabb and P A Jenkins
755
Stope Design Based on Realistic Joint Networks
M Grenon and J Hadjigeorgiou
765
Control of Induced Seismicity at El Teniente Mine, Codelco-Chile
E Rojas, P Cavieres, R Dunlop and S Gaete
775
Support Appropriate for Dynamic Loading and Large Static Loading in Block Cave Mining Openings
T R Stacey and W D Ortlepp
783
Seismicity at Big Bell Mine
M Turner and J Player
791
Study of the Geodynamic Regime of the Region Under Large-Scale Mining in the Apatite-Nepheline Deposits in the Kola Peninsula, Russia
A A Kozyrev, V I Panin and V A Maltsev
799
Experience in Block-Pillar Mining Under Rock Burst Conditions
A A Kozyrev, V A Maltsev, V V Rybin and V S Svinin
805
Seismicity
Mass Mining Initiatives and Opportunities Massive Mining Techniques at 2300 m Depth
N M Schwab
813
A New Future at HBMS — The 777 Project
P R Jones
819
Hudson Bay Mining and Smelting — ‘A Turnaround Success’
H R Rood and D W Nisbet
825
The Obuasi Gold Mine — A Mine Under Change
M P Kelly and S C Goel
831
The e-Mine
D Penswick and K Gilliland
839
Study on Mining Method in Deep Position of Jinchuan Nickel Mine, China
M Cai, L Qiao, C Li and S Wang
843
Modelling Assisted In Situ Leach Mining for Non-Ferrous Minerals
J Liu, H Muhlhaus, B H Brady and S Hancock
847
The Potential for Undersea Mining
D Dekker and T McConachy
853
Vertical Pit Mining — A Novel Alternative to Open Pit or Underground Methods for Mining of Appropriate Massive Shallow Orebodies
M S Redford and P J Terbrugge
859
Future Trends for Underground Thick Seam Coal Mining in Australia
B K Hebblewhite
869
Research and Development The New Development of Continuous Mass Mining Technology Research in China
Wu Aixiang, Hu Hua, Gu Desheng, Li Jianxiong and Yu Youling
877
Grade Estimation for Short-Term Planning in Block Caving Mines
E Rubio, M Scoble and W S Dunbar
881
Numerical Simulations of Bulk Handling in Screw Conveyors by Three-Dimensional DEM
Y Shimizu and P Cundall
887
Simulating Block Cave Secondary Breakage — An Application of Information and Operations Management Tools in Mass Mining Systems
S Dessureault, M Scoble and E Rubio
893
Examination of Bit Wear in Rock Drilling Through Wavelet Analysis
F O Otuonye
897
Selection of Hoisting Wire Ropes Used in Indian Mines — An Economical Approach
A K Basu and M K Singh
903
Alternative Access, Mining and Hoisting for Underground Deposits
K Biegaj
911
The Influence of Surface Geometry on the Load Transfer Mechanisms of Grouted Bolts — A Laboratory Study
N I Aziz, B Indraratna and A Dey
917
The Load-Bearing Process of Fully Coupled Rock Bolts
Chunlin Li
933
MassMin 2000
Mass Mining Design Methodology Orebody Delineation and Reserve Estimation — Seeking Common Sense from Extensive Investment in Fact
J Marek
3
Estimation of Resources and Conversion to Reserves — Protocols for the Assessment, Reduction and Management of Risk
J Beniscelli, P Carrasco, P P A Dowd, G Ferguson and E Tulcanaza
9
Characterising the Mining Environment for Underground Mass Mining
E T Brown and K J Rosengren
17
Method Selection for Large-Scale Underground Mining
W Hustrulid
29
The Use of Evaluation Surfaces to Assist in the Determination of Mine Design Criteria
R M Kear
57
Analysis and Management of Mining Risk
J Summers
63
Orebody Delineation and Reserve Estimation — Seeking Common Sense from Extensive Investment in Fact J Marek1 INTRODUCTION The organisers of this function have suggested that this presentation should accomplish the following broad goals:
• set the technical tone for the conference to follow, implying some discussion estimation; and
of
the
‘state-of-the-art’ of
reserve
• raise some current and common problems or deficiencies in our art in order to be thought provoking. As an engineer it is always easier to raise problems than to solve them with state-of-the-art. Consequently, I have had to limit the extensive list of problems that could have been discussed within this presentation. In regard to the state-of-the-art, this presentation will not be an esoteric discussion of the latest statistical estimation processes, although I intend to challenge those with skills in that area to address some of the thought provoking issues later on. The subtitle of the paper was established after reading an excerpt from ‘Life on the Mississippi’ by the American humorist Mark Twain in which Twain observed that the Mississippi River had become shorter during his working period on the River due to natural cut-offs and man made channels. Projecting into the future, he calculated that by now, Cairo, Missouri and New Orleans, Louisiana would become a single city with one mayor (the two cities are still separated by some 500 air miles). As a moral to this story Mark Twain observed that one can obtain: ‘wholesale returns of conjecture out of such a trifling investment of fact’. From time-to-time, the art form of reserve and particularly resource estimation has been guilty of the same sin. Some of us have been guilty of developing extensive block model and statistical science from trifling investment of reliable support data. The causes of this malady are varied. However, it appears that it has become easier to sit at the keyboard fitting variograms or rotating and zooming the orebody than digging through tedious data issues. Consequently, the discussion of the state-of-the-art will focus primarily on the state of the support data with which we must work. There have been advancements in the assay and data collection process, however as always; the reliability of any given data set must be confirmed prior to its use. Within this paper, computer based modelling will refer to block model procedures since the majority of mass mined deposits would geometrically lend themselves to the block model methodology.
STATE OF THE DATA ART The assembly of a block model incorporates a number of different data types and sources. Those sources include topographic information, geologic knowledge, rock density estimates, and of course, the grade estimate. As the industry considers the development or expansion into areas that are more metallurgically complex, and limited in grade, the integration of 1.
Independent Mining Consultants, Inc, 2700 Executive Drive #140, Tucson Arizona 85706, USA.
MassMin 2000
all data disciplines into a single structure for mine planning becomes more important. The model is usually the framework or basis from which mine planning, cost estimation, budgeting, and overall feasibility is established. The integration of all of the data sources requires a broad knowledge of the mining process, as well as the traditional modelling issues of geologic information, and geostatistical methods. Successful modellers must have a broad knowledge and skill base and the ability to work with many different engineering disciplines. Geostatistics and computer geekery are no longer sufficient (and probably never were). The reliance on the hard data points for the assembly of a model is obvious. If the database is not robust, the entire modelling effort is moot, no matter how elegant the method. The final reserve is no better than the data used to assemble it. The infamous JORC Code definition for indicated resource states that (It (the resource) is based on exploration, sampling, and testing information gathered through appropriate techniques from locations such as outcrops, trenches, pits, workings, and drill holes. The locations are too widely or inappropriately spaced to confirm geologic and/or grade continuity but are spaced closely enough for continuity to be assumed). The definition goes immediately to the appropriateness of the data and the spacing of the data. There is (correctly) no discussion of the kriging method, variogram, or estimation variance.
The database is clean? The state-of-the-art in physical data collection and sample preparation procedures has been relatively unchanged for decades. That said many practitioners are not knowledgeable in sample preparation and handling techniques. This is probably a function of the fact that there is little university training or emphasis in the basics of sample preparation. There are some excellent short courses available on the subject, but most of us have learned the subject from the school of experience, or not at all. In the last year, my firm inherited a client assembled block model that appeared correct relative to the database that was used in the assembly of the model. We observed there were multiple drilling programs based on the drill hole series names. The model builder did not know the drill method used for each program, and did not initially recollect that two different assay methods had been used over time. When compared against each other on a nearest neighbour basis, one drill program constituting about 25 per cent of the drill hole data, was decidedly high biased relative to the remaining drill program data. Although the model assembly was state-of-the-art, the database was in a questionable state. In my opinion, this type of situation occurs in as many as 20 to 25 per cent of the models my firm has been asked to review. Consequently, the state-of-the-art may from time-to-time be in a poor state of affairs. Of the models we have had to reject as unacceptable, most have been over support data issues. Rarely has the rejection been over a model methods issue. Many practitioners inherit data sets that predate their involvement on the property or in older districts, predate their birth. Records of sample methods and assay procedures may have been lost to antiquity. In these cases, comparisons against
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3
J MAREK
more recent data should be completed and an archeological dig through the archives is warranted. Proper validation of the data against historic production records is also a viable option for supporting historic data sets. The solution with new projects is obviously to put best practices for sample preparation and data handling into effect from the first moment a bit touches the ground on a project. Drill methods, hole diameters, sample preparation procedure, and assay method documentation should be stored with the certificates of assay and core logs. Today that probably means hanging a ‘readme’ file on the CD. Of equal importance is the quality control program that is an integral part of the data collection process. All well maintained programs should run periodic check assays along with a standards assay program. It has been my experience that most projects do collect and perform these assays. The procedures should not only check pulps, but should also sample coarse rejects, and the other half of core from time to time. However, I find that the analysis of this data is either non-existent, or the analysis consists of a geologist giving an ‘eyeball’ scan to the check assay set to see if they are in the ball park with the original assays. A primary recommendation for all data collectors is to complete a periodic statistical analysis of the quality control data. Use the information as an on going diagnostic to help assure that the data collection process is going well. The procedure is necessarily reactive in nature rather than proactive because the original information is already in a database when the check assay results arrive. However, identifying a problem after three months of drilling and making a change that requires minor re-assay is significantly better than waiting until the model assembly stage to find out that a large portion of the database is biased or unreliable.
New frontiers bring new data challenges A significant positive trend in the data collection state-of-the-art is the incorporation of a broad range of data types into the model assembly process. Although still of major importance, grade is not the only thing being modelled these days. Process metallurgical response, acid rock drainage issues and geotechnical parameters are now the new frontier in modelling. Modelling these parameters however require additional data that may not have a sample preparation protocol established in the same manner that we would expect to use on a metal like gold, for example. Many of the same principals apply such as, unbiased representation of particle size within the sample, and sufficient sample size. Many of the protocols must be modified based on the mineral environment and perceived process option under consideration. A positive trend in the field of metallurgical process modelling is the application of ‘metallurgical mapping techniques’. I use this term to describe the use of a number of bench scale tests with good spatial distribution to measure the variation of process response across the deposit. Rock type, alteration, mineralogy, weathering, and grade can all be compared with the multiple process tests to begin to understand the metallurgical zoning (or lack there of) of the deposit. This new frontier poses some challenging problems in the area of data collection. Measuring flotation response, for example, is a function of the test and process selected. Changing grind size alone will have a significant impact on the test results. Collector and frother selection will and should change with time over the course of a mine life, but trying to determine where to apply the changes when sampling a green field deposit is problematic. A suite of consistent tests followed by one or two iterations of modified procedures on the same set of samples may be the only way to answer the questions to feasibility level confidence.
4
Those of us who are not flotation specialists have little idea of how to measure surface chemistry response on a local scale in a deposit. We know mineralogy is important, but after that, most ore reserve estimators are as challenged trying to understand float response as the process engineer is challenged trying to grapple with Krige’s relationship. Additional assay methods have become more popular in recent years for the categorisation of mineralogy and process response. In copper base metals, the sequential assay process has been a major benefit to categorising and quantifying the process response of a number of deposits. For those not familiar with the method, a number of lixiavants are used in sequence that differentially dissolve the copper minerals. A typical procedure could be: total copper, sulphuric acid soluble (green minerals), followed by cyanide soluble assays on the reject of the acid soluble (chalcocite, some bornite). By subtraction, the Total copper – (acid + cyanide) is a reasonable approximation of primary chalcopyrite. The method provides a quantitative and analytical procedure for measuring the mixture of mineral species in any given sample. If head and tail samples of the metallurgical tests use the same sequential procedures, a good understanding of the process deportment of the mineralogy can be developed. In the case of copper leaching, a recovery function can be developed based on both the acid soluble, and cyanide soluble species. This particular method may require adjustment by deposit. In Chile, one often finds the use of acetic and citric acid soluble copper assays as the first step in the acid soluble sequence. Although unusual when compared with practices in North America, the exotic copper oxides in Chile can be well analysed with these techniques. Be aware of weaknesses in all assay methods. For example, cyanide dissolves enargite, so caution must be applied to the use of cyanide soluble techniques so that arsenic bearing minerals are not reported as soluble sulphide copper. The sequential assay method is promising in that mineral percentage can be measured by something more repeatable than a geologists ‘eyeball’. Gradational changes such as secondary enrichment in high rainfall environments can be represented more correctly than the traditional procedure of drawing a line between primary and secondary mineralisation. When modelling process response, we must still do battle with the process engineer who demands the ‘representative sample’ particularly in complex orebodies where there is more than one type of mineralisation . Although bulk sampling and pilot plant scale work has its very useful place, the traditional approach of using one or a very few bulk samples to define metallurgical response is akin to defining the grade distribution of an entire orebody by digging a 1000 ton hole in the middle of the outcrop. The obvious answer is that both metallurgical mapping programs and pilot scale testing should be considered since both provide specific types of information to describe the deposit spatially. As a suggestion based on experience, avoid the preparation of ‘ore type blend samples’ for metallurgical testing until the process ore type categories are understood and the individual responses tested. I have found that blends created early in the project missed the blend ratio of the final plan by a wide margin consequently causing avoidable heartburn to our colleagues at the mill during start-up. In complex environments, some component of the blend will almost certainly be metallurgically problematic. From a blended sample it is difficult to identify what the problem was and where it came from. If the problematic area is defined by small-scale mapping tests, and if the process zones are in the model, the proper blend ratios are an outfall of preliminary mine plans. Blends can be prepared with varying amounts of the ‘problem’ ore type to determine the plant response and reagent adjustments required.
Brisbane, Qld, 29 October - 2 November 2000
MassMin 2000
OREBODY DELINEATION AND RESERVE ESTIMATION
Experience has shown that the process data format should be carefully thought out when starting the modelling process. The simple task of calculating composite values can be problematic if basic algebra is not applied. The most straightforward example is the calculation of an average recovery from a number of samples. As an alternative, calculate the recovered grade of each sample, and the average recovered grade for all the samples. This algebraically correct result will be different than the simple mean of the individual sample recoveries. When applying weight factors as in composite length, or tonnage weights in resource tabulations, the use of simple recovery is algebraically wrong and will produce the wrong answer. I have found it necessary with all process modelling to do the simple units cancellation that my first year physics professor beat in to me.
Don’t just describe it, assay it! With the popularity of sequential assay and other metallurgical response procedures, the under sampling of one method versus another becomes a serious problem to the reserve estimator. Dual metal deposits are cursed with the same problem. For example, a gold deposit with a silver by-product credit may have silver assays on only ten to 20 per cent of the intervals where gold is assayed. A second example is a copper deposit where total copper is assayed in every interval, but acid soluble copper was only ran if total copper was above 0.50 per cent. In the first case it would be impossible to estimate silver with any precision relative to gold, and the second case, a built in bias exists in the estimation of acid soluble based on the total copper assay. An even larger problem is the case where rock units that are perceived to be barren are not assayed at all. In these cases, the model builder has the choice of placing a value of 0.0 or a code for ‘no assay’ in the missing intervals. The first case is generally an overall low bias. The second case can result in assayed high grades being extrapolated into unassayed areas. Under sampling, as in the first example of silver, usually results in model blocks with no assigned value for silver. When calculating the average silver grade associated with the gold bearing rocks, those blocks are typically and necessarily defaulted to a 0.0 grade. It is understood that assays cost money, but the savings of $US 7.00 to $US 20.00 per drill interval when the cost to drill that same interval was $US 200.00 may not be an efficient use of funds. When one considers that even 0.50 gm/tonne silver over one million tonnes could contribute 12 000 ounces of recoverable silver or about $US 60 000 to the project, one could have purchased 6000 silver assays and not have lost money. Another concern is not running assays because of the perceived or falsely observed condition that one mineral or the other does not exist. Sequential copper assays where cyanide soluble was not run because the geologist did not observe the presence of chalcocite is an example we have faced on a recent project. Chalcocite in low concentrations is hard to log, its usually black and it appears to be good at hiding. There was no choice but to assign 0.0 to all intervals where chalcocite was not assayed. Re-assay of some intervals later proved the assignment of 0.0 to be substantially conservative. In other words there was indeed leachable chalcocite in the unassayed intervals. A recent trend that has been brought over from exploration techniques is the use of multi-element analysis and accessory element assays on regularly spaced intervals throughout the deposit. We recently had the opportunity to work on an African copper project where in addition to total copper, acid soluble, and cyanide soluble, we had total sulphur, and CO2 assays on most drill intervals. The presence of sulphur correlated precisely with the metallurgical mapping response to flotation. No sulphur, no flotation. In the absence of sulphur, the cyanide soluble copper was native copper that also correlated well with gravity recovery
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responses. CO2 results were a direct correlation to gang acid consumption for the leach process component of ore. Altogether, the metallurgical mapping proved the correlation with the accessory assays. The accessory assays were completed at significantly lower cost than additional metallurgical bench scale tests throughout the deposit. When all done, the CO2, sulphur, and sequential copper results categorised the deposit sufficiently thoroughly so that bulk samples could later be collected from specific process category zones within the deposit. The final block model had block-by-block estimates of recoverable copper by leaching, and flotation. The best process for each block could be determined economically and the process allocation stream reliably estimated from the block model. We have observed recent applications where multi-element information combined with the assay data could be used to categorise the deposit into zones of acid generation versus acid neutralisation. I have also observed another company that is developing a model of pyrrhotite location in their deposit using the iron to sulphur ratio from the multi-element data combined with the geologic logging.
Geologic logging The state of the geologic logging art itself is relatively unchanged over the last 100 years. Logging data is now keyed directly into lap top computers at the logging table rather than requiring extensive data transcription, but the process is still a function of human judgement regarding rock types, minerals observed, alteration, and a myriad of other descriptive parameters. The input of this information into the model process is usually difficult because the geologist doing the logging is often trying to describe relative intensity of a gradational parameter. The modelling process historically tends to categorise that information into the binary categories of Presence or Absence, on or off. This may be an acceptance of the difficulty of the gradational judgement, or perhaps a simple way of coping with descriptions like ‘minor cpy, cc dom, min ser, porph’. The human judgement factor is difficult to calibrate between any two core loggers, and just as difficult to keep consistent within one core logger between Monday at 8:00 am and Friday at 5:00 pm. The logging geologists are the only people who have actually looked at each core and assay interval. The pattern recognition developed during that work is valuable but often difficult to quantify. One of the biggest headaches to the core logger is that logging must necessarily be completed weeks in advance of available assay. The logging geologist knowledge of grade is limited to what can be visually seen. So he is forced to record everything that might be important in advance of any confirming knowledge. Determining if a logged parameter may have impact on mineralisation must be determined after the fact by experience guided trial and error. In summary, it should be noted that the field personnel who log core and collect samples tend to be junior individuals, often recently out of school. Their experience in judgement calls are consequently limited compared with a gray head who may have touched 100 000 metres of core in his career. When logged parameters become critical to reserve and resource estimation, the reliability of the log should be questioned. If relogged, by all six geologists on the property, do you get six similar answers or four similar answers with two dissentions. If the later is the case, is the estimate of ‘measured’ confidence? In typical model applications there is no provision to calculate the unexplained variance or nugget effect of core logging data. Perhaps there should be such a measure. In the mean time, I ask myself if I can see the property in core that is being used in the model interpretation. I can’t tell biotite from chalcocite, and I find that I am not alone. The difference is that I admit it to myself and everyone else rather than convince myself that I can differentiate the two. If nothing else, it has kept
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me from having to log core. I figure that if I can reliably identify the parameter or contact of importance, trained geologic professionals should be able to so also. When one starts drilling a deposit it is difficult to know what to log and whether it will be important. However, the final measure is if the logged parameter has anything to do with the grade of the metal to be produced, the recovery of the metal to be produced, or the physical property of the rock containing the metal.
STATE OF THE INTERPRETIVE ART The judgement issues raised during logging are also valid in the geologic interpretation process. The physical effort of ‘connecting the dots’ on section requires continuous consistent judgement to develop a realistic and useful interpretation. The state-of-the-art for this process is now the interactive, solids/wire frame modelling tools available in many software packages. We have found that the tools are generally robust and in many cases convenient, but based on observation, I am not convinced that the resulting interpretation is any better or more accurate than the old fashioned process of drawing lines on sections and plans. The use of screen based interpretation is now becoming such accepted practice that my questioning of the process is often treated as blasphemy or at least heresy. There are some challenges associated with screen based data interpretation that should be kept in mind. The first is scale. A 1:2000 metric map that is 1 m on a side presents a 2 km by 2 km area. That map will now be collapsed to around 20 to 30 cm on a side on the typical work station screen. At that scale it is apparently difficult to observe the ‘red’ dots (data) in the interpreted ‘blue’ area. Statistical checks on the accuracy of the boundary versus the contained data can be embarrassing. Computer generated contours or indicator estimated boundaries typically have lower error rates albeit without the input of geologic sense. The small-scale and the ability to rotate and look at the deposit from any angle also seems to minimise interest in interpreted detail. The shaded solid looks so good, that it can’t possibly be wrong. I have sat next to practitioners of this art form that have literally given me a bad case of spatial disorientation by performing computer generated aerobatics around the deposit. I often ask the ‘pilot’ to take their hand off of the mouse and ask them which way I am looking at the deposit. A finger jabs at the direction indicator window on the screen. ‘Yes, but simply tell me if I am looking North or South and am I looking up or down?’ I am continually amazed how many times I hear the answer ‘I don’t know!’ One can’t help but respond by saying ‘So why are we doing this?’. I strongly suggest stopping the screen work from time to time and plotting map scale levels or sections of the resulting work. Paper is less transient than the CRT screen and a little time at the light table (remember the light table?) often finds errors in interpretation or spatial patterns in the data that would be difficult (impossible?) to observe on the moving CRT. The process of looking at paper forces the brain to slow down and rethink the interpretation in another media. Interpretation is a skill requiring human pattern recognition. The computer is a wonderful tool for doing the boring repetitive work and presenting the data so the human can focus on recognising and interpreting the important patterns. However the brain still requires time to assimilate the data and the speed that one can make a computer assisted poor interpretation seems to be a priority compared with a well considered and reliable interpretation by any means. Use every cost-effective tool at your disposal to get it right. The pencil is a remarkable tool that has been with us for thousands of years. Don’t be embarrassed to admit that you have used one on the project.
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I am forced to draw the conclusion that the words: ‘I checked it on the screen’ should go down as another famous lie along with ‘the cheque is in the mail’. The tools available to the industry to support interpretive work are outstanding compared with those of only a few years before and it is indeed possible to develop well-supported reliable computer assisted interpretations. However it is also possible to fall into Mr Twain’s trap and produce: ‘Wholesale returns from trifling investment’. In this area I fear, the state-of-the-art is more art than solid state. Once an interpretation is complete, the impact of interpreted boundaries on metal grade distribution can be checked by developing nearest neighbour pairs of data from opposite sides of interpreted domain boundaries. The spacing between the data pairs can be sorted by distance so that one obtains pairs that are 10 m, 20 m, … 100 m apart from both sides of a potential estimation boundary. Conventional hypothesis tests can be applied to this paired data to determine if the boundary is indeed a separation between two populations. This tool has led to a number of contentious discussions with interpreting geologists. However the simple comparisons of arithmetic means of closely spaced data from both sides of a computer modelled fence can be fairly compelling evidence when making the judgement of population boundaries.
COMMON CHALLENGES The main problem in discussing challenges facing the modeller and mine planner is focusing on the few that can be presented in this forum and time frame. As the reader will observe from the previous paragraphs on data and interpretation, the author may have presented a somewhat negative outlook on the state-of-the-art. If one assumes that the database has survived the data verification process, there are indeed two challenges that face the industry that this author believes are worthy of discussion. The first is common to all-modelling practices, surface or underground, and the second is unique to the mass mining underground environment.
The traditional modelling process is solving the wrong problem The block model procedures that have been commonly applied over the last 25 years have been driven by the paradigm breaking work of Messrs Matheron and Krige. Countless practical, accurate, and functional models and mine plans have been developed with these tools. However, in the spirit of raising thought provoking issues, I pose the question: Have we been solving the right problem? The kriging method, its derivatives, and the continuously popular inverse distance techniques are all procedures where the block grade is estimated as a weighted average of surrounding drill hole assay composite data. The selected estimator and the input parameters are used to set the relative composite weights. In all but the most esoteric cases, these methods are sensitive to how far you search, and how many samples or composites you use. In a well-drilled environment, the selected estimator is of secondary importance. The kriging estimator in particular is what statisticians call a ‘BLUE’ or Best Linear Unbiased Estimator. Indeed, it has been my experience that carefully applied kriging results in a sound estimate of the global deposit mean. The problem is however that mine planners do not particularly care about the deposit mean. They are significantly more interested in the tonnage and mean grade above a cut-off. In reality, the mine planner, both surface and underground treats a block model as a one iteration simulation of the deposit. The model was designed to provide reasonably reliable estimates of mean grade over large defined
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volumes. The mine planner calculates a cut-off grade and uses that parameter to define a volume within which the mean grade above that defined block count cut-off is tabulated. The intent of the grade estimation routines, and the mine planning application are two different goals. The tonnage and grade above cut-off from a ‘conventional’ model cannot be correct over a wide range of cut-off grades due to the volume-variance relationships that we all understand as Kriges’ Relationship. The classic result is that a kriged or low power inverse distance grade estimate often tends to overestimate tonnage and underestimate grade. Although not strictly true, this is a common response at a number of locations. It is understood that there have been attempts to address these issues with unstable techniques such as the log normal shortcut or MIK (multiple indicator kriging) with affine transform. But their weaknesses must have become well known because we seldom encounter those methods any more. The smoothing of grade and overestimation of tonnage response has been receiving the high tech title of ‘smearing’ among model builders in the US. Successful model builders have all developed practical methods of getting around this problem so that their models provide reasonable estimates of tonnage and grade over a useful range of mine planning cut-off grades. Grade outlines, one or two stage indicators, limiting search radii versus grade, and limiting the maximum number of composites to estimate a block are all techniques that have been used to increase the block variance so that a conventional block model (one grade per block) provides reasonable results over an acceptable range of cutoffs. Mining engineers however are also becoming more clever and the art form of maximising project NPV versus cut-off grade is becoming more and more popular. This type of analysis assumes that the model provides correct estimates of tonnage and grade for all cut-off grades being analysed. I have been involved with projects where the time sequenced cut-off grades have changed by factors of three to four over the projected mine life. These are broad ranges of cut-off over which one assumes the model will provide accurate results. The challenge to the clever estimator is to develop an improved algorithm for grade estimation that produces a grade in each block so that when a cut-off grade is applied, the model prediction of tonnage and grade above cut-off is relatively reliable. Conditional simulation has been the great hope for the solution to this problem. It may finally be the case that the computer hardware is fast enough so that a reasonable set of ‘practical’ case histories can be developed that will finally allow conditional simulation to become part of the main stream of mine planning efforts. There are some North American base metals producers who are experimenting with inverse distance procedures where the power weight factor is increased until recent mine history is matched by the model prediction. The block variance is increased substantially and the power weights approach the values of nearest neighbour assignment. Given large enough volumes such as six months to a year, this procedure can provide sound estimates for mine planning. However, model local estimates compared with production data were often distinctly in error. The weakness falls back to one of the basic premises of geostatistics in that any one sample is not particularly reliable in the face of any reasonable nugget effect. Conventional linear kriging made fewer local errors, but had the smearing issues discussed previously. Although an interesting and educational experiment, this approach is probably not the best available solution. The paradox of sound global estimation versus local reliability reminds me of the infamous Heisenburg Uncertainty Principal in particle physics. One can know either the location or momentum of an electron, but not both. The state of the electron must be
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expressed as a probability. Perhaps the mine plan estimate of tonnage and grade should be expressed as a probability and the mine plan should consider the probability density function of the tonnage and grade estimate. In summary, I believe we have been solving the wrong problem for some time. The industry should be challenged to develop estimation methods that are truly applicable to mine planning. These new tools and their output should be understood by a wide range of users and the final results should provide management with a more reliable and understandable answer to guide decision making.
Where is the ore now? A problem specific to underground bulk mining is the location and grade of the ore once production is commenced. The subsidence issues of block caving in particular have caused block cave planners fits since the production method was developed. In order to illustrate the problem, the following paradox is offered. Again quoting the JORC Code, measured category resources are based on data where: ‘The locations are spaced closely enough to confirm geological and/or grade continuity’. It is possible to develop measured resources and consequently proven category reserves in a block cave feasibility study prior to commitment to production. However, once the undercut is complete and the ore begins to move, the locations of the original samples themselves are in question so that the local confidence in grades overlying the draw points do not qualify as measured or proven category any longer. The locations of the data that were originally spaced close enough to confirm continuity are now themselves in question so that the estimate may no longer strictly qualify as ‘measured’. The uncertainties and in some cases outright amazing feats of rock displacement that occur over active block caves are well known. The only general rule was written by Sir Isaac Newton in that we expect the displacement vector of any given block of rock to have a net downward component. There have been and continue to be simulation predictions and empirical data that provide guidance about the geometry and dilution effects of the draw down cone versus a myriad of geotechnical, and production rate parameters. It is understood that there now are block cave scheduling tools that utilise the geotechnical and production data to predict diluted head grade from the drawpoint over time based on the block model. However, at least some block cave operations I am aware of estimate the vertical draw from each column of blocks and tabulate the predicted grade with simple vertical displacement calculations. Cross dilution may or may not be applied to the blocks prior to the vertical column grade calculations. Although not high tech, the procedure is simple, repeatable, and above all understandable to all involved. When this process is plotted on block model cross sections, the result is an arch geometry of blocks already removed from the model. As simple as it sounds, it would be convenient to have that same cross section plotted with the vertical translation showing the remaining block laying on top of the drawpoint rather than hanging in air. A level map of the material immediately over all of the drawpoints should show the model prediction of available grade at any point in time. The challenge is simply to develop a block translation routine that can be operated simply, reliably, and reversibly. Reversibility means to provide the ability to return to the undisturbed model and repeat the procedure with different assumptions regarding cross dilution, and draw rate, and draw cone geometry. I suspect that at least one of you from a producing mine will corner me in the pub later and tell me that you have been doing this simple task for years. If so, please convince the software vendors to offer the algorithm in a simplified state so that we can all benefit from its use.
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The issue of production planning from a block cave ties immediately to production reporting and production validation. Since there is currently no way to survey the undisturbed volume reporting from the model, the model estimate of tonnage production is typically set equal to the production tonnage report and the corresponding volume tabulated from the model. Over the course of a long period like a year, it is possible to compare the model prediction of diluted ore versus mill head and tail sample actual results. Given a long enough period (a large enough tonnage) the impacts of local uncertainties average out and a reasonable check of the model to mill is possible. The clever will notice that the mine call of head grade has not been included in this model validation process. This is based on my personnel experience where the mine daily grade reports are based on grab samples from the drawpoint, or grab samples from rail cars or trucks. The size distribution of the material in the drawpoint render the mine production samples as highly speculative. Other than helping to determine the stop draw point when several days run have reported as waste, the daily muck bay grab sample is of limited reliability. When taking all of the issues associated with block cave production grade reporting in to account, one is forced to observe that:
• One does not know where the ore at the drawpoint came from.
• One does not know how much cross dilution has occurred prior to reaching the drawpoints.
Block cave miners either intrinsically understand these issues or they delude themselves into believing that one or several of the above bullets points are actually correct. It appears to this observer that better production grade reporting and control from the mine is not particularly important or more effort would have been historically placed on these issues. The move to belt haulage underground does provide some capability to apply belt scales for tonnage reporting. However, it is doubtful that statistically correct head grade samplers could be regularly employed without installing a sample crushing facility in the circuit to get sufficiently large tonnage to overcome the size distribution of the material In-place, and unbiased samples from the material immediately over the drawpoints would be a substantial benefit to block cave production planning. However drilling and sampling procedures that would work reliably in moving ground are certainly problematic at this time. Unfortunately, the geomechanics constraints, and the above described grade reporting issues mean that head grade manipulation from a block cave operations is limited. There is only minor ability to concentrate on high or low grade zones depending on the metal market and corporate strategic policy. Consequently, profitability has traditionally caused the block cave miner to focus on reduction of operating costs rather than maximising head grade or net profit. An interesting question to ponder is if there were an improved ability to predict future production by panel, would there be a better overall strategic policy for maximising share holder wealth?
• One does not know the in situ or bulked volume of the tonnage reported from the drawpoints.
CONCLUSION
• The tonnage estimate is typically based on bucket or hoist counts.
• The mine production estimate of head grade is highly questionable. The mill heads and tails are better only in that they are usually reconciled against the smelter. The only thing one is sure of is that the mill will report a lower head grade than the mine, and the best estimate of the contained metal is the final smelter settlement report that occurs some months after the production occurred. All of these issues considered, it makes one wonder if the remaining reserve at an advanced block cave operation still qualifies as ‘Indicated’ or ‘Probable’ and consequently as a reserve. Mill and smelter production history versus model guided budget prediction is perhaps the only confirmation information that gives one any confidence at all.
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By now the reader will have identified this article for the collection of old stories and unsolved issues that it is. For the most part, the main points that have been brought forward have intended to point out that better knowledge of the orebody is the key to better estimates for tactical and strategic production planning. Since the ‘perfect’ model is out of reach. The best hope is for the ‘close enough’ model. Consequently, the goal of the effort was summarised in the subtitle in that we seek Common Sense estimates that are developed from Extensive Investment in Fact rather than Wholesale Conjecture from Trifling Investment of Fact.
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Estimation of Resources and Conversion to Reserves — Protocols for the Assessment, Reduction and Management of Risk J Beniscelli1, P Carrasco2, P P A Dowd3, G Ferguson4 and E Tulcanaza5 INTRODUCTION International codes of practice offer broad definitions with respect to the classification of resources in terms of risk. Specific guidelines and protocols, to analyse sample data, carry out geostatistical appraisals, develop geological models and finally estimate and classify resources, are not addressed. Neither are detailed requirements for engineering mineable reserves from estimated resources. In addition, international codes of practice fail to distinguish between types of deposit, other than in broad terms, as for example, The AusIMM JORC Code (JORC, 1999) which makes reference to diamond and coal deposits. The objective of this paper is therefore to set in motion the process of specifying detailed protocols, acceptable to the international community, for the assessment, reduction and management of risk associated with the development and extraction of mineral deposits. Ordinarily, the main risks and issues encountered during the process of estimating resources and subsequent conversion to reserves, encompass:
1.
following internationally accepted practices for the collection, manipulation and analysis of geoscientific data; and
2.
verifying that the procedures actually performed meet the associated standards.
Increasing amounts of data will also reduce uncertainty, but only if the procedures used to collect and appraise such data follow the relevant procedures. It is thus the aim of a series of papers in preparation, Beniscelli et al (in prep), to put forward detailed guidelines, for:
• project sponsors – to provide a global view of the work and documentation needed to support requests for funding various phases of project development and budget approvals;
• geologists – to indicate a checklist of the requirements and standards for the exploration and estimation of resources; and
• engineers and economic analysists – the protocols for characterising, modelling, engineering and evaluating the extraction of reserves.
• failure to understand relevant geological controls of mineralisation;
• validation of grade interpolation; • optimism in geotechnical appraisals of dilution and stability and the consequent selection of inappropriate mining methods;
• metallurgical modelling of minerals and contaminants
It is therefore hoped to encourage the development of consistent standards for declaring reserves within the mining industry. Data from projects and operations will be utilised to illustrate the benefits of adopting a rigorous approach to estimating resources and engineering reserves, Beniscelli et al (in prep).
reporting to concentrates;
• appraising the hydraulics of heap leaching and the numerous factors that can affect recovery;
• difficulty of managing metal price cycles and changes in regulations within the constraints of long-term mining plans; and
• development of better evaluation techniques for mining ventures. Uncertainty is inherent in estimation and engineering work, and its identification, reduction and management should be a fundamental goal of all involved in the mining industry (Figure 1). Variability may be reduced by:
ESTIMATION OF RESOURCES Four main activities characterise the process of estimating mineral resources (Figure 1):
• • • •
geoscientific data collection, model derivation and validation, estimation and classification of resources, and exploration program.
Geoscientific data collection The objective of sampling is to provide sufficient, reliable, unbiased and precise geoscientific data to establish continuity of mineralisation and consequently, to support the estimation of resources. Normally, the main sampling techniques include core and non-core drilling, channel and grab or chip samples. Data are gathered to determine (Table 1) the:
1.
Mining Consultant, Geplam, Codelco-Chile, Huerfanos 1270, Santiago, Chile.
2.
Geological Consultant, Geplam, Codelco-Chile, Huerfanos 1270, Santiago, Chile.
• quantity and spatial distribution of mineralisation and
3.
Head, Department of Mining and Mineral Engineering, Head, School of Process, Environmental and Materials Engineering, University of Leeds, Leeds, LS2 9JT, United Kingdom.
• shape and volume of mineralised zones; • geotechnical characteristics of mineralised bodies and
4.
Mining Consultant, The Seltrust Group, Regal House, 70 London Road, Twickenham, Middlesex TW1 3QS, England, United Kingdom.
5.
Manager, Geplam, Codelco-Chile, Huerfanos 1270, Santiago, Chile.
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associated subproducts;
adjacent country rock; and
• geometallurgical properties of the mineralised zones and affiliated contaminants. The combined process of collecting and assaying samples is the most critical activity in estimating resources. Without high standard and unbiased assays, the ensuing estimation work will have little value.
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FIG 1 - Estimation of resources and conversion to reserves.
TABLE 1 Sampling program objectives. Geology
Geotechnics
Metallurgy
• Grades and density • Stability
• Work index/energy usage
• Geological units
• Mining methods
• Kinetics and reagents
• Orebody volume
• Layout and sequence • Acid usage
• Future exploration
• Dilution
• Recoveries Impurities/ contaminants
Model derivation and validation A major fatal flaw encountered within the process of estimating resources, is the failure to understand the relevant geological controls of mineralisation. An appropriate geological model of a deposit provides the foundation upon which all consequent estimation and engineering is based. It is therefore essential to develop sound geological models with respect to the established geological controls. Subsequently, geological units, which are used to control the interpolation of mineral grades and densities, may be developed. Evolving geological units, is a major focus and integral task of estimating resources.
Estimation and classification of resources The estimation of resources, following the derivation of models and suitable geological units (Figure 2) involves the:
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1.
interpolation of grades within geological units;
2.
calculation of tonnage - from densities previously assigned; and
3.
classification of the estimated tonnes and interpolated grades.
Grade interpolation In order to generate practical results from the interpolation of grade, in a form suitable for mine planning, it is important to utilise for the interpolation work: 1.
a strategy appropriate to the resource type; and
2.
controls defined by the geological units developed, together with pertinent search techniques.
Validation of the interpolation of grade is an integral part of the estimation process. Within the verification procedure, the treatment of high-grade samples is especially critical. Validation of grade interpolation should establish that search and interpolation criteria are free from conditional and global bias. A good verification check, is to compare plots of tonnage versus grade for both the estimated and actual data. Significant variation of estimated data from the actual data can be readily seen and should give some indication as to where differences may lie. The conditional bias suffered by sectional, polygonal and inverse squared methods of grade interpolation is considered reason enough to preclude the use of such techniques.
Tonnage estimation Calculation of tonnage within block models is usually based upon the densities of specific geological units. Thus, the density of the geological unit which passes through the centre of a block should be used to calculate the block tonnage.
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FIG 2 - Estimation of resources - typical methodology.
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Block model
GEOTECHNICAL ENGINEERING
Block models are vehicles for reporting interpolated grades, tonnes and other data. Construction of a block model is normally based upon the plans and sections generated from geological modelling work. Geological units may be assigned to the block model by coding drill hole core in lengths of say 1.5 m. Details of the configuration and information recorded for block models and individual blocks, should include: 1.
Total number of blocks - detailed in the vertical and two horizontal directions.
2.
Block size and orientation.
3.
X; Y; Z block co-ordinates.
4.
A code for the geological unit of the block.
5.
Estimated kriged grade.
6.
Kriging variance.
7.
Estimated kriged grade and tonnage above cut-off grade.
8.
Number of samples used in the estimation.
9.
A code indicating the proportion of the block below the surface. For example IT = 20 means that 20 per cent of the block is rock and 80 per cent is air.
Geotechnical engineering is a key function for the assessment, reduction and management of risk with respect to: 1.
appraisal of stability and dilution;
2.
selection of appropriate support systems, mining method layout and sequence;
3.
effect of mining on surface installations; and
4.
stability of surface infrastructure such as the foundations of shaft headframes, concentrators and tailings dams.
Consequently, geotechnical data are collected as part of the work of the geological team. Other data are gathered by geotechnical personnel from the laboratory, (properties of the rock mass), and field (stress and hydrological measurements). In addition, it is useful to report and consider previous ground conditions experienced, if there is a history of mining in the project location, in terms of the:
10. Block tonnage. 11. A resource category code: 1 = Measured; 2 = Indicated; 3 = Inferred; 4 = default value and 5 = not estimated. Indicators are used in kriging to estimate the proportion of each block above cut-off grade.
1.
mining history – methods, planning factors and success;
2.
design and stability of pit slopes, underground access and extraction development; and
3.
type and success of support systems.
In order to carry out sound assessments of geotechnical risk, engineering work encompasses the derivation of rock mass and stress models for use in a number of techniques, such as: 1.
rock mass classification systems;
2.
limiting equilibrium; and
3.
Classification issues Given an in-depth understanding of geological controls, various methods can be utilised to classify resources into three categories which reflect the: 1.
quantity and quality of data available;
2.
degree of correlation or continuity assigned; and
3.
confidence demonstrated by the assessment.
A prime objective of the system adopted to classify the degree of confidence in the estimate of resources must be to relate the definition and expectation of continuity to a practical drill hole spacing. Techniques in use, include:
• sampling density methods: decreasing size of drilling grid, or tonnes per metre drilled;
• precision of estimation; • resource reliability rating (block core recovery), with
Utilisation of specific methods should be supported by reference to their successful use in similar environments and for similar mining methods. Groundwater considerations with respect to the prediction of drainage, inflow into surface and underground mines and pore pressures in pit slopes demand specific engineering techniques. Finally, monitoring programs, as described by Franklin (1990) are formulated and implemented with the objectives of: 1.
indicating how monitored;
2.
initiating the program as early as possible in project development, in order to obtain a datum level prior to monitoring the response of the rock mass to mining; and
3.
aiding the estimation of costs.
precision of estimation;
• relative kriging variance; and • conditional simulation. Detailed descriptions of these methods are widely available. Contributions of particular interest have been made by Royle (1977), Diehl and David (1982), Ravenscroft (1992) and Annels (1996).
Exploration program
2D and 3D numerical modelling utilising the following codes: - boundary element, - finite element, - finite difference, and - discrete element.
potential
problem
1.
-
design and operation of systems, computer-based data acquisition,
Exploration is ordinarily an on-going activity in most mining operations, and should be carried out to:
3.
seismic events and blast vibrations,
• reconcile the estimation model with the reality of production
4.
rock stress changes and movements,
5.
support pressures, loads and strains,
6.
water pressures and flows,
7.
tailings dams, waste dumps and backfill.
or indicated category; and
• replace exhausted reserves.
12
be
monitoring objectives,
pit slope stability,
• infill zones where data are insufficient to support a measured
would
Such programs should consider:
2.
results;
areas
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ESTIMATION OF RESOURCES AND CONVERSION TO RESERVES
MINING PLAN
1.
Ordinarily, the greatest risks involved in designing and operating mines include:
mining methods and equipment for drilling, blasting, loading and hauling ore and waste;
2.
development and stripping requirements to achieve and maintain an optimum rate of production in underground and open pit mines; and
3.
general mine infrastructure and transportation of personnel and supplies.
• • • •
inadequate selection and design of the mining method; higher dilution than estimated; greater costs than appraised; and unstable conditions affecting mining costs and productivity.
Mine engineering tasks should therefore focus upon these areas. The essential elements of developing a mining plan involve:
• setting an optimum rate of production and defining an associated extraction strategy;
• selecting an appropriate mining method - open pit or underground;
• engineering
the method infrastructure and services;
and
the
supporting
mine
Engineering design – methods – infrastructure and services Method specific techniques are employed for the design and optimisation of both open pit and underground mines. Individual methodologies are involved in these processes and reference should be made elsewhere for detailed information and guidelines as proposed by Ferguson (in prep). Assessments have to be made of the services and utilities required to support the mine plan and associated operations:
• determining operational and manpower requirements; and • drawing up schedules of activities and resources.
1.
ore and rock handling facilities;
2.
explosives magazine;
3.
workshops and equipment maintenance facilities;
Production rate and extraction strategy
4.
compressed air – mobile and static plant;
Attaining an optimum rate of production has significant economic ramifications for all mineral deposits, given the:
5.
mine drainage and slimes handling facilities;
6.
administration, technical and laboratory facilities;
7.
environmental management – waste disposal and safety services;
8.
power supply and distribution;
9.
delivery and storage of fuel and lube; and
1.
adverse effects of fixed costs; and
2.
positive effect of high production rates with respect to economies of scale, both in the mine and plant.
Various rules of thumb exist for setting annual mine rates of production which attempt to balance the size and value of resources against the cash costs of removing them as efficiently as possible. A useful guide to follow is the relationship developed by Taylor (1986) derived from a substantial database of projects. The technique therefore has a historical basis and provides a good starting point in the optimisation process. In essence, the method balances costs against the ability of a deposit to sustain the selected rate of production. In determining an optimum output, cognisance must be taken of the: 1.
mining method;
2.
physical restrictions of deposits – with respect to access for mining equipment; and
3.
on the other hand, possible market constraints with respect to the supply, demand and price relationship.
Given the inter-related nature of mining adjacent sectors or pushbacks, it is necessary to derive a global mine extraction strategy and development plan. Mine extraction sequences have to consider practical, economic and geotechnical factors. In general, within practical and geotechnical constraints, potential plans should broadly follow a declining cut-off grade strategy, whether open pit, or underground. Given an optimised extraction program, development considerations should be taken into account and programs derived to meet the needs of the production plans.
Mining method selection Recommendations should be made with respect to the mining method and sequence necessary to secure an optimum run-of-mine output, taking into account the geological formation of the deposit, the distribution of the mineralisation, geotechnical and economic factors. Aspects that should be covered by the appraisal, utilising industrial methods of engineering with reference to benchmarking data of existing mines and operations in similar environments, include:
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10. industrial and distribution.
potable
water
supply,
storage
and
These should be integrated within the overall requirements of the mine and processing complexes.
Operational and manpower requirements An important area of the mining engineering work is to establish the equipment, operational and manpower requirements to effectively carry out the plan. Estimates should thus focus upon the main elements for specific activities: 1.
equipment,
2.
consumables,
3.
equipment spares and maintenance materials,
4.
power, fuel and lube, and water, and
5.
manpower.
Given the significance of labour costs in most projects and operations it is important to establish, as early as possible, the likely contribution of manpower to the project costs. Thus, a series of organograms should be prepared for each activity and activity group, detailing the levels of manpower required in relation to the mining methodology and associated equipment. In addition, an estimate should be made of the contractor, maintenance and services labour envisaged to support the mining plan, based upon an optimum production rate, mining methods and associated equipment. Accordingly, levels and class of supervision may be assigned. Manpower estimates should be based upon industry benchmarking data and should include allowances for mining engineering staff and other engineering and technical services personnel, such as survey and ventilation.
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13
J BENISCELLI et al
Schedules
2.
Schedules provide the basis for financial evaluation, as well as determining the magnitude of the investment required and on-going funding. Schedules therefore have to relate to the proposed extraction and development strategies, as detailed in the mining plan. Thus, detailed schedules of the various physical quantities and work units are prepared. Project scheduling software should be used to compile and co-ordinate the plans of the following activities and resources as suggested by Ferguson (2000a):
characterisation of data to develop geometallurgical units and predictive models.
Geometallurgical characterisation of deposits, (in the form of geometallurgical units), is needed for long-term planning and evaluation, whereas predictive numerical models, based upon testwork results, are required to support:
• the engineering of reserves from estimated resources; • production decisions; and • reconciliation and grade control.
1.
production,
2.
development,
3.
equipment,
4.
mine services,
Solids modelling
5.
provision of utilities,
6.
supplies, and
The methodology of defining geometallurgical units is similar to the associated generation of geological units, necessitating:
7.
manpower.
Thus, geometallurgical units are essentially long-term planning tools, whereas predictive models are utilised for both long-term planning and operational purposes.
1.
core log data – lithology, mineralogy, mineral hardness, alteration and presence of clay;
2.
a knowledge of the geological environment and structural zones;
3.
assay results from sampling – metal grades, contaminants;
4.
microscopy to determine grain size, texture and, if necessary, to confirm the identification of minerals; and
milling,
5.
geological plans and sections.
concentration, leaching, and
Typically, for the analysis of a copper porphyry, some seven geometallurgical overlays would be created, in addition to the seven overlays constructed for the geological model, including:
smelting.
1.
per cent Fe per sample,
2.
As ppm per sample,
3.
Sb ppm per sample,
4.
mineral grain size,
5.
mineral hardness,
METALLURGICAL ENGINEERING Common metallurgical technical risks that can jeopardise the viability of mining operations exist in all phases of the extraction process (Table 2):
• • • • •
crushing,
TABLE 2 Metallurgical technical risks. Activity
Risk
Crushing
• Mineralogy
6.
alteration and clay content, and
Milling
• Work Index • Run-of-mine granulometry - SAG mills
7.
structural zone.
Concentration
• • • •
Leaching
• Complexity of mineral and gangue mineralogy • Clays within leaching materials • Absolute temperature and daily variation in temperature • Different leaching characteristics of low, medium and high-grade minerals • Balance of capacities: mining - agglomeration leaching - and SX/EW processes • Insolubility of some oxide minerals and solubility of some sulphide minerals
Smelting and Refining
Contaminants Grain size Mineralogy Flotation characteristics
• Contaminants • Environmental pollution • Waste disposal
14
Geometallurgical units The creation of appropriate geometallurgical units, involves the analysis of: 1.
summary statistics of the original sample data - histograms and probability plots;
2.
summary statistics of composited data - histograms and probability plots; and
3.
variography – experimental and modelled variograms.
Assay data are queried by histograms and probability plots to ascertain:
The adoption and implementation of metallurgical protocols is thus essential with respect to determining the characteristics of host rocks, mineralisation, gangue and consequently, run-of-mine mixes. Engineering work thus needs to be based upon a well founded program of: 1.
Given the overlays, 3D solids models can be constructed for each parameter and related factors such as work index, for subsequent interpolation work and distribution within block models.
geoscientific data collection from field and laboratory testwork; and
1.
the distribution of contaminants and other factors; and
2.
establish the number contaminant sample set.
of
populations
within
each
Given the analysis of data, determination of appropriate geometallurgical units, and construction of 3D solids models, the interpolation of contaminants and other factors may be carried out. Subsequently, testwork is undertaken to develop process
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ESTIMATION OF RESOURCES AND CONVERSION TO RESERVES
models and associated designs for heap leaching oxide minerals and the concentration of sulphide materials. Collection of metallurgical data required for the design of potential leaching operations should concentrate upon an assessment of heap leaching characteristics: 1.
chemical characteristics – mineral grade average and range, mineral location in relation to grain size, content of chlorides, silica, clays and other impurities;
2.
physical characteristics of minerals - natural moisture content, specific gravity, bulk density, impregnation moisture content, dynamic moisture content, angle of repose and work index;
3.
iso-pH leach and acid dosage tests; and
4.
dissolution rate of metals.
However, work should also be undertaken on the minerals processing aspects of crushing and agglomeration of run-of-mine ore. In a similar fashion, collection of metallurgical data required for the design of concentration processes should focus upon an assessment of comminution and flotation characteristics: 1.
physical characteristics of minerals - natural moisture content, specific gravity, bulk density, angle of repose and work index;
2.
chemical characteristics – mineral grade average and range, mineral location vis a vis grain size, mineralogy, content of arsenic, iron, clays and other impurities; and
3.
re-milling and cleaning of concentrates.
Some form of metallurgical testwork is required at each project stage to appraise likely recoveries and possible problems from impurities. Tests should be carried out on representative samples and representative mixes of geometallurgical units associated with the proposed mining program. Methodologies are available to relate results from laboratory and pilot scale investigations to industrial size operations, for example, with respect to oxides projects: 1.
leach cycle,
2.
copper recovery,
3.
copper recovery scaling factor, and
4.
acid consumption.
analysis of grade, tonnage, dilution, metallurgical recovery and penalties, and the price of products. Although the single most sensitive factor in almost all mining projects is the ore reserve, it is either ignored or improperly used in sensitivity and risk analyses Dowd (1994). Unlike financial variables, the components of the ore reserve, (grades and tonnages of mining blocks), are a function of location within the orebody, and physical (mining), access to these components is therefore a function of time, which has a critical effect on forecast cash flows. It is essential that these fundamental concepts are included in risk analyses of mining projects. In general, economic evaluation work for mine budgets and projects should encompass: 1.
compilation of the project costs estimated by discipline engineers;
2.
derivation of revenues from the sale of metals and subproducts;
3.
determination of the cost of sales and penalties applied against the level of impurities in concentrates and cathodes;
4.
discounted after tax cash flow analysis; and
5.
analysis of technical risks.
An Activity Based Costing methodology as put forward Turney (1996), is recommended for the estimation of costs, which: 1.
allocates the cost of resources to activities and thence to cost objects;
2.
takes into account the performance factors of the resources used to undertake the activities and other material circumstances that influence the efficiency of the activity;
3.
assigns the costs of resources to activities; and
4.
consigns activity costs to cost objects.
Utilising this system, provides a clear audit trail of the assumptions and calculations used to arrive at the estimate of costs. The system, is very much in line with industrial engineering and zero based budgeting methods with respect to appraising the performance of activities, with a view to improving productivity and lowering costs. With regard to the evaluation of operations and projects, Smith (1991) recommends:
Metals and impurity schedules
• applying different rates of inflation to labour, materials,
Given the establishment of predictive models and suitable scaling factors, it is possible to schedule production, subproducts and the associated impurities. Key to reconciliation and improvement of the geological and metallurgical models, is the installation of suitable sampling points throughout the extractive process.
• developing and evaluating after tax cash flows.
INFRASTRUCTURE AND SERVICES At first sight it may appear that the provision of infrastructure and services is unlikely to encompass risks of sufficient level to affect the viability of projects. However, mining operations cannot operate without power and water, and free and unfettered access to the property. The engineering of these areas therefore demands special attention, and in particular the assumptions made at the early stages in the life of a project.
BUDGET AND EVALUATION Assumptions made in the derivation of budgets and project cash flows can pose significant risk. Perhaps the greatest enemy is optimism, for example with respect to a realistic sensitivity
MassMin 2000
utilities and metal prices; and Historically, metal prices invariably inflate at a lower rate than costs. It is therefore important to know when a project will become unprofitable and to assess what steps will be needed to maintain profitability. Evaluations that do not take into account tax bear no relationship to reality. Tax must be paid by all businesses and is usually a substantial charge. Thus, the timing of tax payments and the allowances against which tax may be offset, are major factors and cannot be ignored. Different companies utilise different criteria to rank and evaluate projects. The internal rate of return is probably the most common technique which can be related to other criteria. A minimum after tax internal rate of return of 15 per cent would probably trigger a project go-ahead. Obviously, if net present value is the criteria for ranking projects, the discount rate of return selected to carry out the evaluation is of prime importance as various workers have indicated (Brealey and Myers, 1987; Smith, 1991; Smith, 1994).
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15
J BENISCELLI et al
MINEABLE RESERVES It is clear that the essential criteria that ‘allow’ conversion from a resource classification to a reserve classification, include the completion and appraisal of:
short-term basis, over meaningful periods of production, to provide a dynamic validation of models and estimates. For example, as a new open pit bench is exposed and mapped, the geological model should be updated and the subsequent re-estimation reconciled with the borehole and mapping results.
• geotechnical characterisation of the ore resource to ensure the realistic appraisal of dilution, support needs and selection of appropriate mining methods, all of which have a major impact upon operating cost and securing an uninterrupted supply of mineral to the plant;
• geometallurgical characterisation of the ore resource for utilisation in the mining plan and to assure that the economic evaluation is founded upon realistic estimates of recovery, work index and contaminants, which affect estimates of revenue, power requirements and smelting penalties respectively;
• infrastructure requirements which can present fatal flaws with respect to the appraisal of isolated mining sites; and
• economic evaluations which consider: tax and nominal cash flows; and encompass risk and sensitivity analyses that take into account the location and access to ore reserves. Additional areas have to be examined, which include:
The broad extent of the technical requirements to develop a well engineered reserve from a resource, estimated using internationally accepted standards, with the fundamental objective of managing risk, has been demonstrated. Further, and more detailed clarification is needed for the assessment, reduction and management of risk, based upon industrial experience and practice. The most common fatal flaws, and hence greatest risks, associated with estimating resources and engineering reserves, include:
• failure to develop an understanding of the relevant geological controls;
• manipulating
classification criteria to demonstrate preconceived notions of the size and quality of resources;
• optimism in geotechnical appraisals of dilution and stability; and
• unsatisfactory
1.
ownership and permits,
2.
environmental management,
3.
implementation of good closure practices, and
4.
reconciliation and grade control.
metallurgical
models
and
engineering
assumptions.
Obviously, projects must meet the requirements of the mining regulations. In addition, the risks of poor environmental management can attract substantial costs, as can operating and constructing mining projects without thought of closure requirements. An environmental management plan should be put in place to monitor the sensitive variables indicated in the environmental impact statement. The plan should incorporate measures to minimise or avoid all risks of air, ground and water contamination by leakage or superficial run-off of solutions, aerosols, gases and solid particles. Specific measures should be described for leach solutions, acid storage, heap leaching, acid mist control, rainwater management, dust emission control, control of solid residues and the sewage system. Equally, there are strong financial arguments for implementing the requirements for mine closure as early as possible in the life of a project, especially prior to designing the site. The objectives of a mine closure plan are to:
• assure the controlling authorities of a successful closure; and • release the mine owners and operators from their obligations so that the site can be disposed of in an appropriate manner. It may also be necessary to provide some means of financial surety to indemnify the authorities against closure and rehabilitation costs. Cut-off grade is a factor that provides a foundation from which the estimation of resources may proceed, but which has both technical and economic implications that have to be clearly established. The declaration of reserves is the most appropriate place to consider cut-off grades, given the technical and economic parameters. Reconciliation of the estimation of resources and the engineering of reserves, against production results, should form an integral part of the declaration of reserves, as a means of controlling the quality of the estimation and engineering procedures used. Reconciliation is best performed on a
16
CONCLUSIONS
Finally, suitable economic and risk evaluation methodologies are required, capable of taking into account price fluctuations, the effect of short-term changes in mining regulations upon long-term mining plans, and the grade and tonnage components of ore reserves.
REFERENCES Annels, A E, 1996. Ore reserves: errors and classification, Trans Instn Min Metall, A (105) 8 p Brealey, R A and Myers, S C, 1987. Principles of Corporate Finance, 847 p (McGraw-Hill). Diehl, P and David, M, 1982. Classification of ore reserves/resources based upon geostatistical methods, CIM Bull, 75 (838):127-136. Dowd, P A, 1994. Risk assessment in reserve estimation and open-pit planning, Trans Instn Min Metall, A (103):A148-A154. Ferguson, G A, 2000. Mining Engineer’s Toolkit, Camborne School of Mines Association Journal, pp 10-13. Ferguson, G A, (in prep). Suggested Protocols for Estimating Resources and Engineering Reserves. http://194.168.97.8/~gavin.ferguson/seltrust/protocols.pdf. Franklin, J, 1990. Mine Monitoring Manual. Special Volume 42, 156 p (The Canadian Institute of Mining and Metallurgy). JORC, 1999. Australasian Code for Reporting of Mineral Resources and Ore Reserves (The JORC Code), The Joint Ore Reserves Committee of The Australasian Institute of Mining and Metallurgy, Australian Institute of Geoscientists and Minerals Council of Australia. Ravenscroft, P J, 1992. Recoverable reserve estimation by conditional simulation, in Case Histories and Methods in Manual Resource Evaluation, pp 289-298 (Geological Society Special Publication). Royle, A G, 1977. How to use geostatistics in ore classification, World Mining, 52-56. Taylor, H K, 1986. Rates of working of mines - a simple rule of thumb, Trans Instn Min Metall, A (95):A203-A204. Turney, P B B, 1996. Activity Based Costing – The Performance Breakthrough,. 382 p (Kogan Page Ltd). Smith, J, 1991. How companies value properties, CIM Bull, 84(953):50-52. Smith, L D, 1994. Discount rates and risk assessment in mineral project evaluations, Trans Instn Min Metall, A (103):A137-A147.
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Characterising the Mining Environment for Underground Mass Mining E T Brown1 and K J Rosengren2 ABSTRACT The purposes of defining the mining environment for underground mass mining and the uses of characterisation data are outlined. The requirements for data in the areas of geology, surface and groundwater hydrology, topography and environmental constraints and geotechnical studies are identified. Emphasis is then placed on the collection and interpretation of geotechnical data for establishing the in situ stress regime and for rock mass characterisation. Methods used for diamond drilling, core logging, borehole logging, data analysis and presentation, rock mass modelling and rock mass classification are outlined. Finally, recent and potential advances in the state-of-the-art are identified in areas such as geophysical methods, measurement while drilling, computerised data collection and analysis, rock mass modelling, in situ stress measurement data presentation and interpretation, and some applications of rock mass characterisation data.
INTRODUCTION For the purposes of this paper, underground mass mining will be taken to include caving methods of mining (block, panel and sublevel caving) and open stoping and its derivatives (including benching) on a scale which requires the use of sub-levels. The introduction of caving methods of mining from the end of the nineteenth century (Peele, 1941) preceded the development of modern mining geomechanics. However, it can be argued that the introduction of open stoping and related methods of mining massive metalliferous orebodies depended greatly on the development of modern mining geomechanics which took place from about 1960 (Hood and Brown, 1999). Central to these developments were methods of collecting and processing data of a range of types which allowed the rock mass and the environment in which mining takes place to be described qualitatively and quantitatively for use in mine design studies. In recent years there has been renewed interest internationally in the mining of massive, often lower grade, orebodies by caving methods. A feature of block and panel caving methods of mining is that, while they have low operating costs, they require high levels of capital investment in infrastructure and development before production can commence. A second feature of these methods is that the mine construction and development required are not readily or economically adaptable to other methods of mining if, for some reason, the chosen mining method proves to be unsuccessful. Accordingly, it is especially important that, when these methods of mining are being considered, the mining environment, especially the geotechnical environment, is understood sufficiently well to permit critical decisions to be made reliably in the pre-feasibility and feasibility study stages of a project. Not to do so invites disaster. This paper outlines the needs for data collection and analysis in defining the (mainly) geotechnical environment for underground mass mining projects and the methods used for collecting those data. In these areas, the paper builds on earlier more general accounts of geotechnical data collection and analyses given by Hoek and Brown (1980), Brown (1981) and
Priest (1993) and the more specific accounts of Bartlett (1992), Laubscher (1981, 1993), Mathews and Rosengren (1986) and Panek and Melvin (1987). The paper also draws on work carried out recently as part of the International Caving Study (eg Onederra, 1999; Harries, 2000; JKMRC and Itasca, 2000). The paper does not consider in detail the requirements for establishing the general geology of the mine site and its surrounds or for delineating the orebody and the distribution of grades within it. It will be assumed that these studies will have been carried out before, or are being carried out in parallel with, those considered here. However, as will be argued below, there is much to be gained by considering the requirements for geotechnical data collection when planning and executing diamond drilling campaigns for exploration purposes.
PURPOSES OF DEFINING THE MINING ENVIRONMENT As Laubscher (1993) has suggested, the mining environment as it is being identified here must be defined, or re-defined, in a number of circumstances:
• for new mining projects on greenfield sites; • for planning the mining of new mining blocks or orebodies in current operations at established mines;
• where a change from open pit to underground mass mining is being considered; and
• where difficulties encountered in operations require a review of the current mining method, planning parameters, layout or detailed mine design. Examples of each of these circumstances have been encountered in mass mining projects and operations in recent years and are addressed in papers to be presented at this conference. In each of these circumstances, data of the types being considered here may be required for the following major purposes:
• mining method selection including cavability studies and stable span evaluation;
• detailed design of mining excavations including their sizes, shapes and requirements for support and reinforcement;
• fragmentation studies which influence issues such as drawpoint spacing and design and equipment selection (including crushers);
• extraction or production level layout in caving mines and detailed design requirements;
including
support
and
reinforcement
• mine infrastructure location and design; • impacts of mining on the surface including the nature and extent of caving zones, interactions with water courses or storages, impacts on surface installations, and impacts on local communities;
1.
Senior Deputy Vice-Chancellor, The University of Queensland, Brisbane Qld 4072. E-mail [email protected]
• where large-scale open stoping is being considered, the need
2.
FAusIMM, Principal, Kevin Rosengren and Associates Pty Ltd, PO Box 361, Ashgrove Qld 4060.
• risk assessment, especially for major hazards such as mud
MassMin 2000
for and availability of fill; and rushes, major instabilities and associated air blasts.
Brisbane, Qld, 29 October - 2 November 2000
17
E T BROWN and K J ROSENGREN
DATA REQUIREMENTS The data required for these purposes can be considered as falling into several categories.
Geology It may be assumed that the regional geology will have been assessed during the exploration stage of a new project or will be well established and understood on continuing projects (eg Howell and Molloy, 1960). The local mine geology must be known and understood in some detail for the purposes identified above. Knowledge is required of issues such as:
• orebody shape, size and the distribution of grades; • the nature of the country rocks and of any weathered or transported overburden materials; and
• structural features such as faults, shear zones, dykes, sills and folding. Generally, the information available from exploration drilling is incomplete with the result that planning and production engineers may find themselves ‘mining blind’ (Hood et al, 1999). Poor detailed knowledge of the orebody geometry in underground metalliferous mines can result in dilution or incomplete recovery or both. Developing the ability to ‘see’ through the rock mass in order to gain a more detailed knowledge of the ore grades and boundaries, and of the rock structure and strength, would bring immense benefit. Geophysical techniques using seismic and electromagnetic methods, for example, are considered likely to provide an effective means of supplementing the information available for drilling (Hood et al, 1999). It is especially important that both major and minor faults and shear zones intersecting the orebody and the nearby country rock be identified. The potentially deleterious effects of faults intersecting, or in close proximity to, mining excavations have long been recognised and dealt with. The nature and magnitudes of these effects may vary with the orientation of the fault, the geomechanical properties of the adjoining rock, the nature of the fault material and fault surfaces (friable or broken material, clay or other filling, slickensiding), the size of the excavation and the presence of water. Some of the observed effects of faults on underground mining excavations include:
• off-setting of the orebody; • slip on the fault leading to a re-distribution of stresses around the excavation;
• fretting or chimneying of friable fault and surrounding material above the back or hanging wall;
• isolation of large blocks or wedges that become free to slide or fall into the excavation;
• general sloughing of destressed or unrestrained rock leading to dilution;
• inability to form a satisfactory anchorage for, or to complete the installation of, reinforcing elements such as rock bolts and cable bolts; and
• the provision of a conduit for water flows into the excavations. Sourineni et al (1999) recently carried out a study of fault-related sloughing in open stopes and gave several examples of major fault-induced failure. Heslop (2000) points to several effects of faults in block caving operations.
Surface and groundwater hydrology Surface and groundwater management is of little concern in some mass mining operations but is vitally important in others. It is necessary, therefore, that issues such as the location of surface
18
water courses and storages, rainwater drainage, groundwater hydrology (including the potential for recharge) are evaluated in the feasibility study stage. If in caving operations, the ingress of water into the caving zone can be prevented, the mining excavations will serve to drain the surrounding rock mass with the result that the mine will be dry. However, in other cases, including areas of extremely high rainfall and where there are adjacent water storages and tailings dumps, water control and mud rush problems can be of concern (eg Barber et al, 2000). Recently, Butcher (2000) has noted that mud rushes have plagued large-scale underground diamond mining operations since the latter part of the nineteenth century. He identifies the main causes of mud rushes as being:
• underground workings daylighting to surface; • poor water control/management/drainage systems design which, in a caving operation, will result in groundwater and rain water entering the muck pile above the current workings;
• poor practices for waste management including − dumping of tailings and slimes in old open cuts or open pits, and
− siting of tailings dams above caving operations; and • poor draw control practices, causing ingress of mud or water by drawing the waste capping above the ore block.
Topography and environmental constraints The topography in the area of the orebody will have a major influence on the locations and costs of surface infrastructure and underground accesses. The local topography will also have an influence on the hydrological issues just discussed, on the local in situ stresses (see below) and on the way in which any caving zone eventually propagates to surface (eg Brown and Ferguson, 1979). Obviously, the existence of communities and of utilities such as roads, power lines and pipe lines of various types in the zone likely to be affected by the mine must be established and taken into account. There are many examples of the positive impacts of new mining projects on local communities by providing jobs, improved services and custom for local businesses. The environmental impacts of mining have become of major concern in recent years. They are noted here for completeness and will be considered no further. A particular issue in some parts of the world, including Australia, is native land rights and the existence in the area influenced by mining of sacred and archaeological sites. Expert studies of all of these issues are usually required to inform the definition of the mining environment of any mass mining project. The Century Zinc Project in North West Queensland, Australia, provides an especially good example of the successful resolution of issues of this type (Williams, 1999).
Geotechnical studies Most of the key issues referred to in the section outlining the uses of the data required to define the mining environment, require geotechnical data for their resolution. Thus, the emphasis of the remainder of this paper will be on the collection and assessment of geotechnical data. As well as the general geological and hydrogeological data referred to above, the geotechnical data required for the purposes being considered here includes:
• Measurements or estimates of the regional and mine site in situ stresses. The stresses induced around mining excavations have major influences on excavation stability and, importantly in the current context, on cave propagation (Kendrick, 1970; Krstulovic, 1979; van As and Jeffrey, 2000).
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CHARACTERISING THE MINING ENVIRONMENT FOR UNDERGROUND MASS MINING
• Measurements of the physical and mechanical properties of the lithological units making up the orebody and the immediate country rocks. These include:
− − − − −
unit weights, uniaxial compressive and tensile strengths, shear strengths of discontinuities, shear strength parameters of intact rocks, stiffnesses and deformation moduli of discontinuities and intact samples,
− hardness, toughness, abrasivity and drillability indices. Standard methods of measuring these properties are given by Brown (1981).
• Discontinuity survey data obtained through core logging, downhole logging of boreholes or scanline mapping of exposed faces. The data required are the locations, orientations, nature and condition of all discontinuities encountered and, in exposures, their terminations. These data are vitally important in cavability, fragmentation and excavation stability assessment studies.
• Rock mass classification of all lithological units using one of the established methods such as those due to Barton, Lien and Lunde (1974), Bieniawski (1976) or Laubscher (1977, 1994). These values are used in cavability studies, empirical methods of stability assessment and in estimating rock mass strengths using methods such as those developed by Laubscher (1977, 1994) and Hoek and Brown (1980, 1997). In the authors’ experience this geotechnical data collection phase is not always carried out adequately in terms of the nature, quantity or quality of the data collected or the time at which it becomes available for use in feasibility and subsequent mine design studies. The remainder of this paper will concentrate on this geotechnical data collection component of studies aimed at defining the mining environment for potential underground mass mining projects. A small number of selected topics will be discussed in some detail in preference to attempting to provide a necessarily brief overview of all aspects of geotechnical data collection and their analysis and application.
IN SITU STRESSES It is well established that the behaviour of underground excavations is influenced by the pre-excavation state of stress (eg Hoek and Brown, 1980). The stresses and displacements induced in the rock surrounding an excavation will depend on the initial state of stress (which may, itself, have been influenced by other nearby openings), the geometry of the excavation and the constitutive (stress-strain) behaviour of the rock mass. These induced stresses and displacements will influence the stability of the excavation, the need for reinforcement of the rock mass or filling the excavation, and the initiation and propagation of caving. Except in particular geological environments, it is usually not possible to predict the in situ state of stress using the principles of mechanics. This is because both the magnitudes and orientations of stresses are influenced by a wide range of factors including tectonic history, topography, erosion, differences in the elastic constants of the lithological units and the presence of faults and other discontinuities. It must also be remembered that stress is a tensor quantity which requires the quantification of six unknowns in order to define it fully at a point. It may not be assumed that the principal in situ stresses will be oriented horizontally and vertically although, in some circumstances, it may be both reasonable and convenient to do so. Among the first significant measurements of stresses in underground mines were those made in the iron ore mines of
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eastern France in the early-1950s using the then recently developed flat-jack method (Tincelin, 1952). However, it was the pioneering work of Hast (1958) in Sweden which demonstrated that the horizontal stresses in rock could be several times the vertical or overburden stress. The first in situ stress measurements made in Australia made as part of the investigations for the Tumut 2 underground power station in the Snowy Mountains Hydro-electric Scheme (Alexander and Worotnicki, 1957) confirmed this result as have subsequent compilations of measurements in Australia and elsewhere (eg Brown and Windsor, 1990; Hoek and Brown, 1980; Mueller et al, 1997; Zoback, 1992). The magnitudes and orientations of the horizontal components of the in situ stress field can exert major influences on mining excavation costs and performance (eg Gale and Blackwood, 1987; Sandy and Player, 1999). Since the 1950s, a number of approaches have been developed for the measurement of in situ stresses. As well as the original flat jack method which suffers from a number of inherent disadvantages, most emphasis has been placed on a variety of overcoring methods and on the hydraulic fracturing method which, like the flat-jack method makes a number of sometimes limiting assumptions (Brown and Windsor, 1990). For the last two decades, the state-of-the-art method has been the CSIRO hollow inclusion stress cell developed by Worotnicki and Walton (1976). Some ingenious larger scale methods of measuring local stress fields have also been used (eg Brady et al, 1976). Other approaches used to estimate in situ stresses include the resolution of earthquake focal mechanisms, the interpretation of stress conditions associated with young geological features including faults, and the back analysis of wellbore breakouts and excavation behaviour. Overviews of these various methods of measuring and estimating in situ stresses are given by Brown and Windsor (1990) and Amadei and Stephansson (1997). Despite more than 40 years accumulated experience of measuring and estimating stresses at mine sites, the undertaking can still be fraught with difficulty. It has been found that stress magnitudes and orientations can vary markedly with the presence of discontinuities and with changes in rock properties. Even though these effects are often observed, it is still considered essential that every effort be made to measure (from exploration openings) or otherwise estimate the mine- or orebody-scale stresses before decisions are made about the adoption of a particular mass mining method or layout. In the absence of any other source of information, recourse can be made to the several excellent databases and stress maps available on mining district, regional and world scales (eg Amadei and Stephansson, 1997; Mueller et al, 1997; Hillis et al, 1999). An example of a stress map showing the directions of the maximum principal horizontal stresses in Australia is shown in Figure 1. The symbols are those used in the World Stress Map Project (Zoback, 1992). When relying on the data presented in these databases and stress maps, it should be remembered that not all of the results may be assumed to be accurate, particularly at shallow depths where the strains and stress components being measured are low.
ROCK MASS CHARACTERISATION Overview Rock mass characterisation is the process of collecting and processing quantitative and qualitative data which provide descriptors and assessments of the geometrical and mechanical properties of a rock mass. This information is required for most of the purposes outlined earlier. If the required data are not collected and the rock mass(es) involved are not characterised adequately, the likelihood of sound engineering decisions being made is decreased and that of making major and costly errors are increased (Mathews and Rosengren, 1986).
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challenge in the discipline (Glaser and Doolin, 2000). It follows that the assessment of the data collected and the engineering methods which utilise the data must take account of the inherent uncertainties.
Diamond drilling and core logging The proving and delineation of a massive orebodies of the types being considered here is usually carried out by diamond drilling. The core is logged geologically, but not always geotechnically, and then split for assaying. In the authors’ experience, it is common for little or no geotechnical data to be gathered systematically as part of this process. In these cases, a great opportunity to ‘add value’ to the exploration process is lost. Thirty years ago, one of the authors studied the use of diamond drilling for structural or geotechnical purposes (Rosengren, 1970). It is disappointing to note that, despite the advances made since in technology and in rock mechanics knowledge, some of the deficiencies in practice noted then still remain. In order to obtain the maximum value from geotechnical drilling it is necessary that:
• the cores be as large as practicable; • core barrels appropriate to the core size and the rock mass condition are used to ensure maximum recovery; FIG 1 - The Australian stress map (Hillis et al, 1999).
The rock mass characterisation process may involve a number of stages or procedures including:
• core logging and photography; • scanline mapping of exposures; • field and/or laboratory testing of intact core samples and/or discontinuities;
• borehole logging; • storing and analysing data to produce quantitative descriptions of the rock mass geometry;
• developing 3D statistical models which simulate the rock mass geometry; and
• producing quantitative rock mass classifications which enable experientially based estimates to be made of the engineering properties and performance of the rock mass in particular applications. In a mass mining operation, a number of lithological units and conditions of those units are likely to be encountered. It is good practice to divide the rock mass(es) of concern into a series of structural domains each having an approximately homogeneous structure, condition and rock mass properties, and to carry out rock mass characterisation procedures for each of those domains. Although the basic procedures are now well-established and are described in standard works on the subject, new and valuable developments continue to be made. Some of these will be outlined in subsequent subsections. It must be recognised at the outset that whatever techniques are used, the process and the outcomes are both uncertain. Rock masses can be highly variable and only a very small, sometimes a vanishingly small, percentage of the rock mass can be sampled and measured. This applies particularly in greenfield projects where there may be no underground exposures and reliance is placed entirely on drilling. In these cases, serious consideration should be given to establishing an exploration drive or level well above the proposed undercut level in the case of caving operations (Heslop, 2000). At a recent forum on future directions in rock mechanics it was concluded that the characterisation of rock mass properties remains the greatest need and greatest
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• the core be correctly oriented (at least in part); • the core be handled, transported and stored with care; and • the core be logged and photographed when fresh to record the relevant features. Recently, as part of the International Caving Study, Onederra (1999) has carried out a comprehensive review of the available drilling systems, core orientation devices, core recovery and handling techniques and core logging procedures together with case histories. Heslop (2000) provides valuable sets of practical tips for improving geotechnical drilling, core logging and core photography which are similar to those used by the authors. An advance made possible by modern microelectronics and computer technologies is the ability to collect and process a range of data during drilling (Peck and Vynne, 1993). This is sometimes known as Measurement While Drilling or MWD. The results may be used to provide estimates of rock strength and of rock mass quality (eg RQD) and to record major discontinuities (eg faults) and changes in rock type. One of the authors had experience in the application of this approach to tunnelling site investigations some 20 years ago (Brown, 1979; Barr and Brown, 1983) but subsequent improvements in microprocessor technology have made the monitoring and processing of drilling variables much more practicable and effective. Useful geotechnical results are also being obtained from the monitoring of production (percussive) drilling (eg Schunnesson and Holme, 1997). Figure 2 shows a possibly outdated, but nevertheless instructive, example of the output from the instrumented drilling of a short length of a horizontal, 56 mm diameter, diamond drill hole in an underground limestone quarry at Corsham, Wiltshire, England (Barr and Brown, 1983). An Atlas-Copco Diamec 250 drill was instrumented to record rotary speed, head displacement, thrust, torque, inlet water pressure and inlet and outlet water flow rates during drilling. In the record shown in Figure 2, the presence of discontinuities first encountered at depths of 17.25 m and 17.61 m is indicated by the disturbances to the thrust, torque, rotary speed and water pressure records. Both discontinuities were 12 cm wide. The discontinuity encountered at 17.61 m was clay filled as indicated by the increased water pressure resulting from a blocked bit and the associated erratic torque and rotary speed responses. Barr and Brown (1983) also showed how it was possible to use the recorded drilling parameters and an approach developed by Tsoutrelis (1969) to back-calculate values of the
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FIG 2 - Instrumented diamond drilling strip chart record, Great Oolite Limestone, Corsham, UK (Barr and Brown, 1983).
uniaxial compressive strength along the borehole which agreed with those measured in laboratory tests on the core to within ten to 15 per cent. Hood et al (1999) argue that there is great potential to obtain more information from production drilling by the routine monitoring of mechanical drilling variables and by the use of a suite of geophysical sensors developed for the purpose. They suggest that this would provide real-time information on rock properties, discontinuity locations and orebody boundaries that could be used in numerical blasting models to revise blast designs on-line. The Centre for Mining Technology and Equipment is currently carrying out research in this area.
Borehole logging techniques A range of borehole logging techniques are available to obtain quantitative and qualitative data about a number of characteristics of the rock mass. Borehole logging provides additional information which either confirms the data obtained from the core (eg discontinuity locations and orientations) or adds to the data obtained from the core, especially when there has been less than complete core recovery. Borehole logging and testing techniques have been classified as belonging to one of four categories (Van Schalkwyk, 1976):
• observational or visual techniques using borehole cameras some of which allow complete images of the borehole wall to be recorded (eg Kamewada et al, 1990);
• geophysical logging techniques which can collect multiple logs with one pass of a probe, can provide indirect measurements of variations of lithology (density), rock and rock mass properties and the positions and orientations of discontinuities (eg Brown, 1981; Mathews and Rosengren, 1986; Wade et al, 1993);
• mechanical techniques such as dilatometers, borehole jacks, penetrometers, extensometers and inclinometers which can be used to provide assessments of the mechanical properties of the rock mass and to monitor movements; and
• hydrological techniques can be used to measure the in situ permeability of the rock mass, the hydraulic conductivity of a single discontinuity or the water pressure at a point or series of points within the rock mass (eg Elsworth and Mase, 1993).
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Data analysis and presentation In the past, the geotechnical data collected by core logging or scanline mapping were recorded, plotted and processed manually (eg Hoek and Brown, 1980). Over the past 20 years, a range of computerised techniques have been used to collect, but more importantly to store, process and manage the data (eg Hoek et al, 1995; Panek and Melvin, 1987; Onederra, 1999). Pen based computer systems are particularly useful for field data collection. The advantages of storing data electronically in the field include improved efficiency of data collection in terms of time and the sample sizes that can be used, immediate analysis of the data, and immediate storage of the raw data and processed results into a database system (Onederra, 1999). As part of the International Caving Study a program known as JointStats has been developed at the Julius Kruttschnitt Mineral Research Centre (JKMRC and Itasca, 2000). JointStats has three functions or modules:
• data storage and archiving, • simple data analysis, and • detailed data analysis using more advanced statistical methods to produce three-dimensional models of the rock mass geometry. The first function of JointStats is to serve as a discontinuity database for archiving core and face mapping data. The program uses the Microsoft Access database structure and utilities. The data are organised in an hierarchical manner and presented to the user in a tree view. An example is shown in Figure 3. The program also allows for the interactive stereonet plotting of discontinuity orientation data including an on-line Fisher parameter calculation and the associated identification of joint sets. An example of joint set definition with the context menu is shown in Figure 4. Once the joints have been selected and assigned to sets, a range of simple statistical calculations (eg uncorrected and corrected joint spacings, trace lengths, termination statistics) can be made and the results presented or plotted as tables, histograms or graphs. Some of these simple statistics can be used as input to classification systems (see below).
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FIG 3 - Data tree for mine site discontinuity data (JKMRC and Itasca, 2000).
A more complex extension of the statistical analysis of joint or discontinuity data is the development of three-dimensional models which may be used to provide statistical simulations or realisations of the rock mass geometry (eg Priest, 1993; Dershowitz, 1995; Henry et al, 1999). These models may be used to produce simulations of the rock mass geometry from which distinct element stress and deformation analyses, fragmentation assessments and fluid flow simulations may be carried out. The 3D Rock Joint Model in JointStats is used to estimate the joint set parameter values as a maximum likelihood problem (Villaescusa and Brown, 1992). The formulation is statistically rigorous and takes all stereological biases into account. The calculated confidence intervals for the parameter values reflect both the quantity of the data originally available and the extent to which the model fits the observed structure. Information of this type is useful in helping quantify the degree of confidence that may be placed in assessments made from rock mass characterisation data of characteristics of the rock mass such as cavability or primary fragmentation (JKMRC and Itasca, 2000). In common with many other modelling approaches, the 3D Rock Joint Model uses a forward modelling procedure (Dershowitz, 1995) in which the mathematical properties of the model are used in a statistically rigorous way to match the computed values to measured joint data. These methods assume that:
• the joints can be assumed to be flat circular discs or planes of a variety of other shapes (including ellipses, rectangles or polygons) whose centres are distributed randomly in space;
• the normal vector to the joint plane follows some statistical (usually the Fisher) distribution; and
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• the diameters or sizes of the joints follow some statistical distribution (usually negative exponential or log-normal). A refinement of this general approach is used in hierarchical models (Gervais et al, 1995; Harries 2000) in which joints are added sequentially into the simulation from the oldest to the youngest set as determined from structural geological analysis. In the model developed by Harries (2000), as each fracture is added to the simulation, a search is made for intersections with existing joints. Where intersections occur, a probability of termination rule is applied to model termination at the older joint or within a rock block. This modelling approach can produce improved representations of the discontinuity network geometry and of the in situ block size distribution. Figure 5 shows a photograph and a digitised fracture trace map of a 13.5 m high cutting in the Brisbane tuff at Kangaroo Point, Brisbane (Harries, 2000). The influence of joint termination can be seen clearly in this example. It is unlikely that the geometry of this rock mass could be recreated and the block size distribution modelled accurately without incorporating the effects of joint termination.
Classification systems A number of rock mass classification schemes are used to provide a basis for the extrapolation of experience in one rock mass or application to another. These classification schemes seek to assign numerical values to those properties or features of a rock mass considered likely to influence its engineering behaviour, and to combine these individual values into an overall numerical rating for the rock mass. Rating values for the rock
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FIG 4 - Discontinuity set definition using a lower hemisphere equal angle projection, with context menu (JKMRC and Itasca, 2000).
FIG 5 - Photograph and digitised fracture trace map of a 13.5 m high face in Brisbane tuff (Harries, 2000).
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masses associated with a number of projects are then determined and correlated with observed rock mass behaviour. Aspects of rock mass behaviour that have been studied in this way include stable spans, stand-up times for unsupported spans, cavability, stable open pit slope angles, hangingwall caving angles and fragmentation. A number of these assessments made from geotechnical data collected in the exploration or feasibility study phases of a mining project may provide useful guides to the selection of an appropriate mining method. The NGI tunnelling quality or Q index developed by Barton, Lien and Lunde (1974) uses six numerical inputs – the Rock Quality Designation (RQD), a joint set number, a joint roughness number, a joint alteration number, a joint water reduction factor and a stress reduction factor. The other widely used classification is the Rock Mass Rating (RMR) system due to Bieniawski (1976). This scheme uses five classification parameters – strength of the intact rock material, RQD, joint spacing, condition of joints and groundwater conditions. The RMR may be adjusted for the influence of joint orientation on excavation stability. Laubscher (1977, 1990, 1993) extended and modified Bieniawski’s classification for mining applications, notably cavability studies. Laubscher has introduced a number of adjustments for factors such as joint condition, weathering, field and induced stresses, joint orientation with respect to excavation surfaces and blast damage to produce a Mining Rock Mass Rating (MRMR). It is said that: . . . the ratings, details of the mining environment, and the way in which this affects the rockmass and geological interpretation are used to define the cavability, subsidence angles, failure zones, fragmentation, undercut-face shape, cave-front orientation, undercutting sequence, overall mining sequence, and support design . . . (Laubscher, 1994). A major application of the MRMR is for the prediction of the onset of continuous caving using Laubscher’s caving chart which plots MRMR against the hydraulic radius of the undercut to correlate experience of continuous caving (Laubscher, 1993, 1994). Stable, transitional and caving zones are shown on the resulting graph. Recent experience has shown that some difficulty or uncertainty can be experienced in assigning values of the adjustment factors in determining the MRMR (eg Milne et al, 1998) and that the method may not always give reliable cavability predictions in stronger, confined rock masses. A major application of rock mass classification schemes in open stope and more general underground excavation design is Mathews’ stability graph method (Mathews et al, 1980; Stewart and Forsyth, 1995; Trueman et al, 2000). Mathews et al (1981) developed an empirical relation between rock mass quality, mining depth, stope dimensions and stability. They defined a stability number, N, which is the product of Q’, the value of the NGI Q index with the joint water and stress reduction factors put equal to unity, a rock stress factor, a rock defect orientation factor and a design surface orientation factor. The stability number, N, is plotted against the shape factor or hydraulic radius, S (= area/perimeter), of the excavation surface whose stability is being considered. Stable, minor failure and major failure or potentially caving zones may be defined on the stability graph. Over the years since its introduction, the original data base of case histories has been greatly extended and modifications suggested (eg Potvin et al, 1989; Stewart and Forsyth, 1995; Trueman et al, 2000). Variations of this general approach, sometimes using purpose-built classification schemes, have been developed and applied in particular mining environments (eg Villaescusa, 1996).
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Recently, Trueman et al (2000) have extended the Mathews’ stability graph method for application to open stope design by carrying out a logistical regression analysis on a greatly increased database of case histories to re-define the stable and minor and major failure zones. They found that by plotting N and S values on log-log axes rather than the original log-linear axes, the boundaries between the zones could be represented as parallel straight lines (Figure 6). In order to take into account the inherent variability of rock masses and the non-rigorous nature of the method, Trueman et al (2000) also added iso-probability contours to their graph (Figure 7).
Data collection and performance monitoring during mining The uncertainties inherent in the rock mass characterisation process have been outlined above. In geotechnical engineering generally, and in mining geomechanics, one of the means used to deal with the uncertainty is to apply the observational method formalised by Peck (1969). In this approach, further data are collected during excavation and the performance of the excavations, or the outcomes in terms of fragmentation for example, are monitored. The new data are then fed back into the original characterisation results and the models, designs or conclusions revised as appropriate. A generalised illustration of this process is shown in Figure 8. Adding further data during exploration, development and mining and the monitoring of stope or cave performance are essential components of the rock mass characterisation process in mass mining. It is a folly to assume that once the initial characterisation has been done, and design parameters determined, the task is complete. The rock mass characterisation process should be an on-going one.
CONCLUSIONS Rock masses are highly complex and variable engineering materials. The design and operation of modern underground mines using mass mining methods must be such that they are both cost-effective and safe from several points of view. Poor decisions resulting from inadequate information or poor or inappropriate engineering procedures, or from a combination of both, can have disastrous economic and social consequences. It is essential, therefore, that the mining environment be characterised adequately in terms of both the quantity and quality of the data obtained, before key decisions are made on mining feasibility and the mine design is finalised. Adequate data are also required subsequently for detailed planning of the mining operations. The requirements for data used to characterise the mining environment for underground mass mining have been outlined. Emphasis has been placed on the collection and uses of geotechnical data. The techniques used for collecting this data are reasonably well developed. Their effective application and the interpretation of the data so obtained require the use of common sense and a modicum of experience. Many of the methods used to interpret and apply the data are based on simple empirical correlations which may not be universally applicable and so require care for their effective use. In a fundamental sense, the issues involved in geotechnical mine site characterisation and the approaches used to address them have not changed greatly in the more than 30 years in which the authors have been associated with studies of this type. This aspect of mining and geotechnical engineering has seen evolutionary development rather than revolutionary change over that period of time. However, there have been some recent developments, and the prospect of some to come, which show promise of bringing improvements to the collection, interpretation and use of site characterisation data. They include:
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1000.000
STABLE 100.000 Stable-Failure Line 58 % Stable 42 % Failure 0 % Caving
Stability Number, N
10.000
FAILURE AND MAJOR FAILURE
1.000 CAVING
0.100
LEGEND Stable 0.010
Failure Major Failure
Caving Line 89 % Caving 11 % Failure 0 % Stable
Caving Stable-Failure Boundary Failure-Caving Boundary
0.001 1
10
100
Shape Factor, S
FIG 6 - Extended Mathews stability graph for open stopes (Trueman et al, 2000).
1000.000
95% 90% 80% 70% 60% 50% 30% 40% 20%
100.000
10% 5% 0%
Stability Number, N
10.000
95% 90%
1.000
80% 70% 60% 50% 40% 30% 20% 10%
0.100
5%
LEGEND
0%
Stable
0.010
Failure Major Failure Caving
0.001 1
10
100
Shape Factor, S
FIG 7 - Isoprobability contours for open stope stability (Trueman et al, 2000).
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FIG 8 - Generalised application of the observational method (Glaser and Doolin, 2000).
• improved geophysical tools for providing a more detailed knowledge of orebody geometry and rock mass properties;
• the further development of measurement while drilling techniques for both rotary and percussive drilling to improve the range, quantity and quality of the data collected about orebody geometry and rock and rock mass properties;
• computer-based methods of systematically recording, processing and storing geotechnical and other data;
• the development of statistically rigorous three-dimensional models of rock mass geometry which may be used in numerical analyses or simulations of excavation stability, fragmentation and fluid flow through the rock mass;
• revisions to, and extensions of, existing empirical methods of estimating open stope stability and cavability based on rock mass characterisation data; and
• the compilation and publication via the Internet of stress measurement data on mining district, regional, and world scales.
REFERENCES Alexander, L G and Worotnicki, G, 1957. Rock stress tests in Tumut 2 exploratory tunnel, Snowy Mountains Hydro-Electric Authority, Engineering Physics Report SE26. Amadei, B and Stephansson, O, 1997. Rock Stress and its Measurement, 515p (Kluwer: Dordrecht). Barber, J, Thomas, L and Casten, T, 2000. Freeport Indonesia’s Deep Ore Zone Mine, in Proceedings MassMin 2000, pp 289-294 (The Australasian Institute of Mining and Metallurgy: Melbourne.) Barr, M V and Brown, E T, 1983. A site exploration trial using instrumented horizontal drilling, in Proceedings Fifth International Congress on Rock Mechanics, pp A51-57 (Balkema: Rotterdam). Bartlett, P J, 1992. The design and operation of a mechanised cave at Premier Diamond Mine, in Proceedings MASSMIN 92, pp 223-231 (The South African Institute of Mining and Metallurgy: Johannesburg). Barton, N, Lien, R and Lunde, J, 1974. Engineering classification of rock masses for the design of tunnel support, Rock Mechanics, 6:189-239. Bieniawski, Z T, 1976. Rock mass classifications in rock engineering, in Proceedings Symposium on Exploration for Rock Engineering (Ed: Z T Bieniawski), 1: 97-106 (Balkema: Cape Town).
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Brady, B H G, Friday, R G and Alexander, L G, 1976. Stress measurement in a bored raise at the Mount Isa Mine, in Proceedings ISRM Symposium Investigation of Stress in Rock: Advances in Stress Measurement, pp 12-16 (Institution of Engineers, Australia: Sydney). Brown, E T, 1979. CIRIA instrumented drilling trials – background and progress, Ground Engineering, 12 (1):45-52. Brown, E T (Ed), 1981. Rock Characterization, Testing and Monitoring – ISRM Suggested Methods, 211 p (Pergamon: Oxford). Brown, E T and Ferguson, G A, 1979. Progressive hangingwall caving at Gath’s mine, Rhodesia, Trans Instn Min Metall, 88: A92-105. Brown, E T and Windsor, C R, 1990. Near surface and in situ stresses in Australia and their influence on underground construction, in Proceedings Seventh Australian Tunnelling Conference, pp 18-48 (Institution of Engineers, Australia: Canberra). Butcher, R J, 2000. Hazards associated with the mining of diamondiferous pipes, CIM Bull, 93(1037):65-67. Dershowitz, W, 1995. Interpretation and synthesis of discrete fracture orientation, size, shape, spatial structure and hydrologic data by forward modelling, in Fractured and Jointed Rock Masses (Eds: L R Myer, N G W Cook, R E Goodman and C-F Tsang), pp 579-586 (Balkema: Rotterdam). Elsworth, D and Mase, C R, 1993. Groundwater in rock engineering, in Comprehensive Rock Engineering (Eds: J A Hudson, E T Brown, C Fairhurst and E Hoek) 1: 201-226 (Pergamon: Oxford). Gale, W J and Blackwood, R L, 1987. Stress distributions and rock failure around coal mine roadways, Int J Rock Mech Min Sci and Geomech Abstr, 24(3):165-173. Gervais, F, Gentier, S and Chiles, J-P, 1995. Geostatistical analysis and hierarchical modelling of a fracture network, in Fractured and Jointed Rock Masses (Eds: L R Myer, N G W Cook, R E Goodman and C-F Tsang) pp 153-159 (Balkema: Rotterdam). Glaser, S D and Doolin, D M, 2000. New directions in rock mechanics – report on a forum sponsored by the American Rock Mechanics Association, Int J Rock Mech Min Sci, 37(4):683-698. Harries, N J, 2000. Rock mass characterisation for cave mine engineering. PhD thesis (unpublished), University of Queensland, Brisbane. Hast, N, 1958. The measurement of rock pressure in mines, Årsb Sver Geol Unders, 52(3). Henry, E, Marcotte, D, Gaudreau, D and Nickson, S, 1999. Simulation of a 3D fracture network to predict caving potential of a Cu-Mo deposit, in Proceedings APCOM’99, Computer Applications in the Minerals Industries (Ed: K Dagdelen) pp 217-224 (Colorado School of Mines: Golden).
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Heslop, T G, 2000. Block caving-controllable risks and fatal flaws, in Proceedings MassMin 2000, pp 437-454 (The Australasian Institute of Mining and Metallurgy: Melbourne.) Hillis, R R, Enever, J R and Reynolds, S D, 1999. In situ stress field of eastern Australia, Aus J Earth Sci, 46(5):813-825. Hoek, E and Brown, E T, 1980. Underground Excavations in Rock, 527 p (Institution of Mining and Metallurgy: London). Hoek, E and Brown, E T, 1997. Practical estimates of rock mass strength, Int J Rock Mech Min Sci, 34(8):1165-1186. Hoek, E, Kaiser, P K and Bawden, W F, 1995. Support of Underground Excavations in Hard Rock, 215 p (Balkema: Rotterdam). Hood, M, Hatherley, P and Gurgenci, H, 1999. Mining in 2015, Mining Magazine, 181(4):222-229. Hood, M and Brown, E T, 1999. Mining rock mechanics, yesterday, today and tomorrow, General Report presented to 9th International Congress on Rock Mechanics, Paris, 25-28 August. Howell, F M and Molloy, J S, 1960. Geology of the Braden Orebody, Chile, South America, Econ Geol, 55:863-905. Julius Kruttschnitt Mineral Research Centre and Itasca Consulting Group, 2000. Improving the understanding of the mechanics and design processes of block caving, Confidential Year 2 Report, International Caving Study. Kamewada, S, Gi, H S, Tanagucki, S and Yoneda, H, 1990. Application of borehole image processing system to survey of tunnel, in Proceedings International Symposium on Rock Joints (Eds: N Barton and O Stephansson) pp 277-284 (Balkema: Rotterdam). Kendrick, R, 1970. Induction caving of the Urad Mine, Min Congr J, 56(10):39-44. Krstulovic, G, 1979. Influence of tectonic stresses on the caving process – mining by block caving methods, in Proceedings Fourth International Congress on Rock Mechanics, 1:459-466 (Balkema: Rotterdam). Laubscher, D H, 1977. Geomechanics classification of jointed rock masses – mining applications, Trans Instn Min Metall, 86:A1-8. Laubscher, D H, 1981. Selection of mass underground mining methods, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 23-38 (AIME: New York). Laubscher, D H, 1990. Geomechanics classification system for the rating of rock mass in mine design, J S Afr Inst Min Metall, 90(10): 257-273. Laubscher, D H, 1993. Planning mass mining operations, in Comprehensive Rock Engineering (Eds: J A Hudson, E T Brown, C Fairhurst and E Hoek) 2:547-583 (Pergamon: Oxford). Laubscher, D H, 1994. Cave mining – the state of the art, J S Afr Inst Min Metall, 94(10):279-293. Mathews, K E, Hoek, E, Wyllie, D C and Stewart, S B V, 1981. Prediction of stable excavation spans for mining at depths below 1000m in hard rock, Golder Associates Report to Canada Centre for Mineral and Energy Technology, Dept of Energy, Mines and Resources, Ottawa. Mathews, K E and Rosengren, K J, 1986. The integrated planning of geotechnical investigations for new mining projects, in Proceedings Second Project Development Symposium, pp 43-54 (The Australasian Institute of Mining and Metallurgy: Melbourne). Milne, D, Hadjigeorgiou, J and Pakalnis, R, 1998. Rock mass characterization for underground hard rock mines, Tunnelling and Underground Space Technology, 13(4):383-391. Mueller, B, Wehrle, V and Fuchs, K, 1997. The 1997 release of the World Stress Map, available on-line at: http://www-wsm.physik.uni-karlsruhe.de/pub/Rel97/wsm97.html Onederra, I A, 1999. Geotechnical data collection: core logging and face mapping, Confidential Report, International Caving Study.
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Panek, L A and Melvin, M T, 1987. Fracture geometries in three ore bodies mined by undercut caving as determined from oriented drill core and scanline mapping, US Bur Mines Rep Invn 9049, 82 p. Peck, J and Vynne, J F, 1993. Current status and future trends of monitoring technology for drills, in Proceedings International Mining Geology Conference, pp 311-325 (The Australasian Institute of Mining and Metallurgy: Melbourne). Peck, R B, 1969. The observational method in applied soil mechanics – 9th Rankine Lecture, Géotechnique, 19: 171-187. Peele, R, 1941. Mining Engineers’ Handbook, 3rd Edn (Wiley: New York). Potvin, Y, Hudyma, M and Miller, H D S, 1989. Design guidelines for open stope support, CIM Bull, 82(926):53-62. Priest, S D, 1993. Discontinuity Analysis for Rock Engineering, 473 p (Chapman and Hall: London). Rosengren, K J (1970). Diamond drilling for structural purposes at Mount Isa, Ind Diamond Rev, 30(359):388-395. Sandy, M P and Player, J R, 1999. Reinforcement design investigations at Big Bell, in Rock Support and Reinforcement Practice in Mining (Eds: E Villaescusa, C R Windsor and A G Thompson) pp 301-315 (Balkema: Rotterdam). Schunnesson, H and Holme, K, 1997. Drill monitoring for geological mine planning in the Viscaria copper mine, Sweden, CIM Bull, 90(1013):83-89. Sourineni, F T, Tannant, D D and Kaiser, P K, 1999. Determination of fault-related sloughage in open stopes, Int J Rock Mech Min Sci, 36(7):891-906. Stewart, S B V and Forsyth, W W, 1995. The Mathews’ method for open stope design, CIM Bull, 88 (992):45-53. Tincelin, E, 1952. Mesures des pressions de terrains dans les mines de fer de l’est, Ann Inst Tech Bat et Trav Publ, Sols et Fond, 58:972-990. Trueman, R, Mikula, P, Mawdesley, C and Harries, N, 2000. Experience in Australia with the Mathews’ method for open stope design, CIM Bull, 93(1036):162-167. Tsoutrelis, C E, 1969. Determination of the compressive strength of rock in situ or in test blocks using a diamond drill, Int J Rock Mech Min Sci, 6(3):311-321. van As, A and Jeffrey, R G, 2000. Hydraulic fracturing as a cave inducement technique at Northparkes Mines, in Proceedings MassMin 2000, pp 165-172 (The Australasian Institute of Mining and Metallurgy: Melbourne). Van Schalkwyk, A, 1976. Rock engineering testing in exploratory boreholes, in Proceedings Symposium on Exploration for Rock Engineering (Ed: Z T Bieniawski), pp 37-55 (Balkema: Cape Town). Villaescusa, E, 1996. Excavation design for bench stoping at Mount Isa mine, Queensland, Australia, Trans Instn Min Metall, 105: A1-10. Villaescusa, E and Brown, E T, 1992. Maximum likelihood estimation of joint size from trace length measurement, Rock Mech and Rock Engrg, 25(2):67-87. Wade, L, Wang, R and Horton, M A, 1993. Estimation of mechanical rock properties using wireline geophysical measurements, Trans Instn Min Metall, 102:A31-36. Williams, I, 1999. Century Zinc Project, in The Spirit of the Snowy Fifty Years On, Proceedings 1999 Invitation Symposium, pp 183-195 (Australian Academy of Technological Sciences and Engineering: Melbourne). Worotnicki, G and Walton, R J, 1976. Triaxial hollow inclusion gauges for determination of rock stresses in situ, in Proceedings ISRM Symposium Investigation of Stress in Rock: Advances in Stress Measurement, Supplement pp 1-8 (Institution of Engineers, Australia: Sydney). Zoback, M L, 1992. First- and second-order patterns of stress in the lithosphere; the World Stress Map Project, J Geophys Res, 97(B8):11761-11782.
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Method Selection for Large-Scale Underground Mining W Hustrulid1 INTRODUCTION As has been the case since the early Phoenician traders, the minerals used by modern man come from deposits scattered around the globe. The price received is more and more being set by worldwide supply and demand. Thus the price component in the profit equation: Profits = Material sold (units) × (Price/unit – Cost/unit)
(1)
is largely determined by others. The recent price history for two of the major metals, copper and iron, are given in Figures 1 and 2, respectively. As can be seen there has been a steady decline in the prices (when expressed in constant dollars) over the 20 years which have lapsed since the first of the International Conferences on Mass Mining, ‘Design and Operation of Caving and Sublevel Stoping Mines’ was held. If one would look back over the entire 20th century this would also be the trend. Cost containment/reduction through efficient, safe and environmentally responsive mining practices is serious business today and will be even more important in the future with increasing mining depths and more stringent regulations. A failure to keep up is expressed very simply as: Profit < 0
FIG 1 - The price trend for copper (after Edelstein, 2000).
(2)
This, needless to say, is unfavourable for all concerned (the employees, the company, and the country or nation). The mining companies have had to look at ways of reducing the expense part of the profit equation while maintaining safe, environmentally conscious operations. Table 1 summarises the rules for working in today’s economy. TABLE 1 Rules of working (Collahausi, 2000). 1.
High production
2.
Competitive cost
3.
Safety
4.
Teamwork
Production of metals comes both from open pit and underground mines with the predominant source being open pits. In their attempts to hold down costs, the open pit miners have gone to ever increasing equipment sizes. Table 2 summarises the largest available loading and hauling models supplied by several well-known manufacturers for years 1980 and 2000. As can be seen, both shovel and truck sizes have increased by about a factor of 3.5 over this period of time. To keep these large machines operating at greatest efficiency, pit geometries have also increased. Constant or even decreasing open pit mining costs (as measured in constant dollars) has meant that the breakeven depth for the consideration of underground mining methods has steadily increased. Today there are pits of 800 m depth with 1.
Department of Mining Engineering, University of Utah, 135 South 1460 East, Utah 84112, USA. E-mail: [email protected]
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FIG 2 - The price trend for iron ore (after Kirk, 2000).
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TABLE 2 Generalied development of mining production equipment. After Caterpillar (1999), Dietz (2000), Jackson (2000) and Reise (2000). Year
TABLE 4 The possible negative consequences. 1.
The very large automated machines and techniques are incompatible with steep, well-groomed slopes. Precision cutting of bench faces and maintenance of bench widths are required.
2.
All the production comes from a few working places. The plant must be very flexible to handle varying feeds. Incompatible with most current plants.
3.
There is a need for many different stockpiles and re-handling.
4.
There are a great number of safety aspects to be considered with automated machinery and a mixture of automated and non-automated jobs.
Equipment Capacity Shovels (yd3)
Truck Payload (tons)
1980
20
100
2000
67
360
plans to go to 1200 m. Even with mining costs typically below $1/ton moved, the open pit miners are pushing hard to achieve new cost reductions. Some of the potential improvements in open pit mining are listed in Table 3. Unfortunately, as listed in Table 4, there are some accompanying downsides. TABLE 3 Projected ‘advances’ in open pit mining. 1.
Increase in mining geometry, steeper slopes, deeper pits
2.
New, larger production machines/techniques - 450 ton capacity trucks - 100 yd3 shovels - 15″ to 17.5″ drill holes - continuously variable explosive with respect to energy and other properties - electronic blasting caps - blast designs made in the field based on drillability results and fragmentation requirements
3.
Mechanisation/automation - very high accuracy GPS on all machines
FIG 3 - Diagrammatic representation of panel caving (after DeWolf, 1981).
- driverless trucks - remotely-operated drills - remotely-operated shovels 4.
MARC in place, guaranteed availability. Advanced condition monitoring and prediction.
5.
Few working places, with high utilisation
6.
Short time between stripping and production
7.
Automatic sampling of the drill cuttings, remote after blast sampling
8.
Automatic sampling in the dipper/truck bed or along the route. Automatic destination assignment.
9.
Simulation is extensively used to plan production.
Although it may seem strange to begin a paper on large-scale underground mining with a discussion of open pit mining, they are the competition. What they can and will do in the future has a direct effect on underground mining decisions. They are one of the driving forces for going to ever larger scales underground. For the purpose of this discussion large-scale mining will be arbitrarily defined as meaning production greater than 5000 tpd (15 × 106 tpy) involving the use of:
• panel caving (Figure 3), • sublevel caving (Figure 4), and • sublevel stoping (Figure 5).
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FIG 4 - Diagrammatic representation of sublevel caving (after Hamrin, 1982).
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TABLE 5 Technical characteristics of future mining systems. 1.
Low specific development.
2.
Minimum time between development and extraction (just-in-time delivery).
3.
High degree of mechanisation.
4.
Move toward sensible automation.
5.
High machine availability.
6.
High machine utilisation.
7.
Safe working conditions.
8.
High resource recovery.
9.
Environmentally sound.
FIG 5 - Diagrammatic representation of large-hole sublevel stoping (after Hamrin, 1982).
To achieve and maintain such production rates, the orebodies involved are also large. As will be discussed in the next section, the orebodies will be assumed to fall under the ‘massive’ and/or ‘thick, steeply dipping’ category. To a large degree, the cost determination process begins with the selection of the appropriate mining system. The use of the word ‘system’ rather than ‘method’ was chosen deliberately since the ‘method’ is but one part of the proper combination of design/layout, equipment selection, sequencing, etc. Mining scale is obviously one of the important considerations. Once the system is in place, the mining engineer focuses on controlling and hopefully reducing the unit costs. Although the development of new technology is one answer, new technology easily and quickly spreads around the world and soon all operations have the ‘new’ technology. Hence to remain profitable over the long-term, the mining engineer must continually examine and assess smarter and better site-specific ways of reducing costs at the operation. This is done through a better understanding of the deposit itself and the tools/techniques employed or employable in the extraction process. To remain competitive with the open pits, the underground miners have had to aggressively pursue cost reductions including change of mining method and mining scale. Since the cost of panel caving, the least expensive of the underground methods, is three to five times more per ton of material handled than in the open pits, this is a very tough challenge. Table 5 summarises some of the general goals/guidelines to be applied when selecting and designing future underground mining systems. The trend underground has been to increase the scale of mining and, by doing so, to try and adapt open pit techniques. For sublevel stoping and sublevel caving, this has meant the use of longer, and larger diameter blast holes. The result is an increase in stoping dimensions and a reduction in the specific development. New highly productive machines are being used. The short-term cycling of activities has been reduced and a factory-like atmosphere created. The lag time between development and production is kept to a minimum. LKAB, using sublevel caving to mine iron ore in northern Sweden, is a prime example of a company making such changes to remain competitive. Figure 6 and Table 6 summarise some of the key
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FIG 6 - Changes in sublevel caving geometry at LKAB (after Marklund and Hustrulid, 1995).
TABLE 6 Comparison of sublevel caving dimensions. Marklund and Hustrulid, 1995. Unit Sublevel interval (m) Sublevel drift spacing (m) Drift size (m) Ring burden (m)
Year 1983
Year 2000
12
27
11
25
5×4
7×5
1.8
3
Hole diameter (m)
0.064
0.114
Tons/ring
1000
10 000
mine design parameters for years 1983 and 2000. As can be seen the sublevel interval has more than doubled and the tonnage per ring has increased by a factor of ten over this period. The size and the capacity of the machines being used have been increased accordingly. In panel caving, changes have also been made to simplify designs, increase extraction heights, and to increase the plan dimensions. All of these have a positive effect on the specific development. New, larger machines are being used. Applications of the technique to ever harder ore types are being attempted.
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Just as with the developments in open pit mining, there are some potential snakes waiting in the woodpile. With the increase in scale, production comes from only a few places. There is (a) an emphasis on just-in-time delivery of development and production and (b) an increasing dependence upon a small fleet of rather sophisticated machines. The designs must function as designed and the as-built must closely resemble the as-designed. With these changes in scale, there is an increased vulnerability to ‘happenings’. The consequences of misjudging the geologic environment are severe. Hence future underground mining competitiveness imposes a heavy responsibility on those that select a large-scale mining system, those that make the design, and those responsible for carrying out the design. The fallback positions from misjudgments/mistakes are, in general, very few and expensive. In the following sections, this paper addresses method selection, in general, and presents some viewpoints regarding certain aspects of modern, large-scale mining systems.
THE SELECTION OF A MINING SYSTEM Traditionally the selection of a mining system (Peele, 1934) involves three separate, but closely related subjects:
• classification of ore deposits, • classification of mining methods, and • selection of a mining method. Ore deposits are normally classified with respect to their nature, shape and the character of the ore and adjacent country rock. Mining methods are classified on the basis of the method of support. By comparing the orebody characteristics with the method characteristics one arrives at a suitable mining method or suitable mining methods. For many deposits, especially those which are narrow and steeply dipping, the selection process is fairly easy. However, as Peele (1934) states: for massive deposits having large lateral dimensions there is a wider choice. The choice of an appropriate method for such deposits is often much less obvious than for ordinary veins. It is complicated by the greater need for considering the economic factors bearing on the problem in addition to the physical characteristics of the orebody and its mode of occurrence. Although this statement was written more that 65 years ago, the challenge of selecting the correct mining method to be used to extract massive deposits remains. Table 7 is a listing of the factors that should be taken into account when selecting a mining method (Warner, 1934). As can be seen, the list includes internal factors, external physical factors, economic factors in the district, and economic factors in the industry. Specific guidance is provided in Table 8 (Dravo, 1974). As can be seen, all three of the methods which are the focus of this paper appear to be equally-likely candidates for the extraction of massive deposits. Nicholas (1981) has proposed a numerical system to assist in the selection of a mining method. The purpose of the system is primarily to indicate those methods that will be most effective given the geometry/grade distribution and rock mechanics characteristics rather than to choose the final method. The system will allow miners/engineers to focus on the characteristics important for the mining methods being considered. The top candidates from this initial screening process would then progress further for more in depth studies. The technique involves the use of four tables (Tables 9 - 12). The selection process has two steps. In Step 1 the characteristics of the deposit are defined (Tables 9 and 10). In Step 2, using the characteristics of the deposits defined in Step 1, values are selected from Tables
32
11 and 12 for each of the mining methods. Each mining method has been ranked as: Preferred:
the characteristic is preferred for the mining method
Probable:
if the characteristic exists, the mining method can be used
Unlikely:
if the characteristic exists, it is unlikely that the mining method would be applied, but does not completely rule out the method
Eliminated:
if the characteristic exists, then the mining method could not be used with regard to the suitability of its geometry/grade distribution (Table 11), and the rock mechanics characteristics of the (a) ore zone (Table 12a), (b) hangingwall (Table 12b), and (c) footwall (Table 12c). The values assigned to each rank are listed in Table 13. The value selected for the ‘eliminated’ rank was chosen so that if the sum of the characteristic values equalled a negative number the method would be eliminated. A zero value was chosen for the ‘unlikely’ rank because it does not add to the chance of using the method, but neither does it eliminate the method. The values used for the ‘probable’ and ‘preferred’ were chosen so that the characteristics for one parameter could be ranked within a mining method and between mining methods (Nicholas, 1981). In practice, for each method the appropriate points are recorded for each of the following categories:
• • • • • • •
general shape, ore thickness, ore plunge, grade distribution, rock substance strength – ore, hangingwall and footwall, fracture spacing – ore, hangingwall and footwall, and fracture strength – ore, hangingwall and footwall.
The totals are then formed and the candidates with the highest number of points selected for further study.
Closing remarks As indicated, there are a number of factors to be considered when selecting a mining method. For some orebodies, the possible methods are few. For any given massive orebody there are a number of different possibilities to be considered. The qualitative terms used in the past to describe orebody characteristics are certainly a starting point for method selection. Unfortunately with the large-scale techniques being considered today and the accompanying high investments, an expression of these characteristics in terms of expected cavability, expected fragmentation, open span widths, etc is required. Today the decision is not just the selection of sublevel stoping but rather sublevel stoping, for example, based on the creation of ‘super’ stopes. If the ‘super’ stopes being considered are not viable, then perhaps another method must be chosen or perhaps the orebody is (a) not ore or (b) it is much smaller than initially projected. Table 14 outlines the traditional steps in the selection of a mining method. Table 15 outlines an alternate scheme for selecting a mining method and Table 16 yet another way. Lacking good, soundly-based, quantitative arguments in this environment of low prices and fierce competition, the latter two unfortunate approaches could be those upon which the final choice is based. As will be discussed in the following section, some significant efforts have been put forth over the past years to provide quantitative tools which, perhaps, can eventually be used with confidence in the mining method decision-making process.
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PREDICTORS OF CAVABILITY AND NON-CAVABILITY Introduction According to Jackson (1934) one of the earliest mining method classification systems of which he was aware had been developed by a Professor Sperr. As a hobby, which was pursued over a period of many years, Professor Sperr created a very elaborate classification which included 176 different mining systems. Morrison (1976), on the other hand, as paraphrased by Coates (1981), indicated that ‘All underground mining methods may be divided into two groups. One is based on preventing caving and the second requires either the immediate or ultimate caving of the overlying ground.’ In support of this concept, Parker (1973) suggests that the openings in room and pillar mining are usually of maximum breadth without caving whereas in block caving the breadth of the undercut is designed to be the minimum breadth to ensure caving. The quantification of ‘caving ease’ has been the goal of a number of studies. In this section the two concepts:
• shear failure, and • snap-through failure
FIG 7 - Two types of roof failure (after Beer and Meek, 1982).
shown in Figure 7 (Beer and Meek, 1982) will be examined. TABLE 7 Factors influencing choice and detail of a mining system (after Warner (1934).
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TABLE 7 (continued) Factors influencing choice and detail of a mining system (after Warner (1934).
TABLE 8 Geologic and mechanical criteria in large-scale mining methods (Dravo, 1974). Mining method
Orebody characteristics Ore strength Weak
Mod
Strong Weak
Room and Pillar1
X
X
Sublevel Stoping2
X
X
Shrinkage
X
X
Cut and Fill
X
X
Square Set
X
Block Caving3
X
Sublevel Caving Longwall
X X
X
X
Orebody configuration
Waste strength
X
Mod X X
Beds
Strong Thick X
X
Veins
Thin
Narrow
Massive Wide
X
X
X
X
X
X
X
Ore dip Flat
Mod
X
X
X
X X
X X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
X
Steep
X X X
X
X
1. Uniform thickness and grade 2. Regular hanging and foot walls 3. Strong fractured rock also can be caved
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Introduction to the shape factor/hydraulic radius concept
TABLE 9 Definition of deposit geometry and grade distribution. Nicholas (1981)
1.
2.
3.
4.
One of the techniques used for describing opening stability incorporates a surface shape factor (S). In stope stability literature this is often replaced by ‘hydraulic radius’ since it is the area divided by the perimeter. As part of the preparations for this paper, the author attempted to determine the background for the use of the ‘hydraulic radius’ concept. Unfortunately, after numerous conversations and communications with a number of knowledgeable people, the origins are still unclear. In this section, one possibility is presented based upon shear failure. In
Geometry of deposit General shape Equi-dimensional:
all dimensions are on the same order of magnitude
platey – tabular:
two dimensions are many times the thickness, which does not usually exeed 100 m (325 ft)
Irregular:
dimensions vary over short distances
TABLE 10 Rock mechanics characteristics (Nicholas, 1981).
Ore thickness narrow:
15
1.
Plunge Flat:
>20°
Intermediate:
20° - 55°
Steep:
>55°
2.
Depth below surface Grade distribution - uniform
Fractures/m
Fractures/ft
>16
>5
0 - 20
Close
10 - 16
3-5
20 - 40
Wide
3 - 10
1-3
40 - 70
3
= 60°. It is normally adequate if α > = φ1 × SF, where φ1 is the angle of friction on the wall of the discharge hopper and SF is the factor of safety. Normally SF varies from 1.05 for a smooth surfaces to 1.15 for rough surfaces. Typical values for φ1 are given in Table 19.
7.
The arch forms more easily if the mobility of the mixture of coarse material is smaller.
8.
Flat arches can be formed more easily if the particles are more irregular and stronger. Sticky constituents assist in the formation of arches. This is indicated in Table 20.
FIG 27 - The 80 per cent size as a function of fall height for different energy efficiencies.
Drawpoint width Once the layout has been made based upon the size characteristics of the caved material, one can design the drawpoint widths. These widths are also a function of the fragmentation which must be extracted. Figure 28 shows a
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TABLE 19 Typical Values for φ1 (after Kvapil, 1965). φ1 (Friction angle along a concrete wall) (degrees)
Material type
Dolomite, limestone, marble
33 - 43
Granite, greywacke, magnesite
35 - 42
Iron ore
36 - 43
Gypsum
32 - 41
Sandstone
34 - 42
TABLE 20 The risk of hangups as a function of material type (Kvapil, 1965). Class
Material Description
Risk of Hang
1
Uniform size distribution (pieces of the same size and form)
None
2
Mixture of pieces of the same size but different form
Very small
3
Mixture of pieces of different size and different form
Medium
4
Mixture of pieces of different size and form with earthy-clayey constituents
Large FIG 29 - Design nomogram for drawpoint width (after Kvapil, 1965).
9.
Arches of coarse material form more easily if the individual blocks of rock are larger, ie if a smaller number is required to form the arch.
10. A smaller area of outlet opening facilitates arching and an enlargement of the area reduces it. 11. The practical elimination of arching during self-acting discharge through a horizontal outlet opening depends on the application of the so-called minimum area required for the outlet opening. The minimum area required for horizontal outlet openings for coarse material can be calculated (Kvapil, 1965) using Fa = (5 × D)2 k (square opening)
(24)
Fφ = 0.85 Fa (circular opening)
(25)
where Fa
= area of the square opening
Ff
= area of the circular opening
D
= average size of the largest pieces
k
= coefficient derived from the nomogram shown in Figure 29.
where a1
= width of the square opening
d1
= diameter of the circular opening.
The coarse material is a mixture of various particle sizes. A wide range of combinations may be produced and their detailed assessment would be very complicated. To simplify matters the nomogram contains only the major fractions which constitute the coarse material. Group I consists of large pieces (percentage from 25 to 100 per cent) possessing various degrees of angularity (rounded, angular, sharp-edged). This characteristic shape is included by using the appropriate functional line. Group II represents medium size particles (coarse pebbles), Group III the finer fractions(gravel, sand, etc) and Group IV the sticky constituents (moist alumina, loam, etc). One travels through the nomogram following the arrow sequence and eventually arrives at the coefficient k whose value varies from 0.6 to 1.4. The nomogram indicates, for example, that coarse material having a sticky constituent (IV) content of over ten per cent is unsuitable for self-acting flow through the outlet opening. In the nomogram example shown, one begins with the most coarse material. The material in this case is angular and makes up 37.5 per cent of the total. Hence, the lowest of the three particle shape lines apply. One then follows the nomogram through the other material fractions.
The factor 0.85 enters in Equation 25 since the circular opening is 15 per cent more effective than that of the square one. The corresponding width of the outlet opening is
Group
Per cent
I
37 (angular)
a1 = (Fa )1/2 (square opening)
(26)
II
20
III
40
d1 = ( Fφ/0.785)1/2 (circular openings)
(27)
IV
0
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The resulting k value is k = 0.66 Assuming that the average size of the largest fragments is 1.5 m and the opening is square, one finds that a1 = 5 × D × (k)1/2 = 6 m The rule of thumb that is often applied to the flow of ore in orepasses is that to avoid hang-ups the diameter should be three to five times the maximum particle size. This is in agreement with the equation/nomogram suggested by Kvapil (1965). The present author suggests using the guideline W = 4 × K80 (primary fragmentation)
(28)
be used as a first approximation. If the K80 size at the beginning of draw is 1.5 m, then the minimum width of the drawpoint should be: W=6m This is wider than the drifts commonly in use today when panel caving hard rock. It is noted that there will still be blocks bigger than the K80 size which will create hang-ups and handling problems in the drawpoints.
Summary remarks One of the key elements in the selection and design of a successful caving system is a knowledge of the fragmentation and fragmentation distribution as a function of draw. If this is poorly known then major mistakes can be made. In the worst case, the panel caving method may be inappropriate and/or the layout chosen may be wrong. Today there is a major gap in our ability to predict fragmentation distributions based on typical core drilling data. In the hardest, most competent rock it is not unusual to obtain average fracture frequency values of the order of 2 - 3/m which would still suggest relatively small block sizes. Work is on-going to better identify those ‘fractures’ which will be mobilised during the caving process (El Teniente, 2000). Once the fragmentation has been determined, it is then necessary to develop guidelines of the type suggested to arrive at the influence radius (R) as a function of draw, drawpoint widths, etc. Lacking such guidelines, in the rush to satisfy certain economic goals, it is possible that this (theoretically) inexpensive mining method will be poorly selected and applied. This applies both for hard and soft rocks. For soft rocks, recovery may be low and dilution high due to economically-based decisions to widen drawpoint spacings. For hard rock, the continued presence of oversize in the drawpoints will kill the productivity projections and hence the projected costs.
SOME ASPECTS OF LARGE-SCALE SUBLEVEL CAVING Introduction Sublevel caving was initially applied in the early-1900s to extract the soft iron ores found on the iron ranges of Minnesota and Michigan. The sublevel caving as practiced today is significantly different from this early version and should probably be given another name, such as sublevel retreat stoping, continuous underhand sublevel stoping, or something similar that would better reflect the process. Today, the sublevel caving technique is applied in hard, strong ore materials in which the hangingwall progressively caves keeping pace with the retreating rings. In looking at older guidelines regarding the selection of sublevel caving, one often sees that the applicability is to soft ores. This reflects the earliest application of the method and should not be
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confused with modern sublevel caving. The key layout and design considerations are to achieve high recovery with an acceptable amount of dilution. The uncertainties of fragmentation and ore cavability present in panel caving (discussed in the previous section) are absent because each ton of ore is drilled and blasted from the sublevels. Sublevels are created at intervals of between 20 and 30 m, beginning at the top of the orebody and working downward. On each sublevel, a series of parallel drifts are driven on a centre-to-centre spacing that is of the same order as the level spacing. From each sublevel drift, vertical or near-vertical fans of blast holes are drilled upward to the immediately overlying sublevels. The distance between fans (the burden) is on the order of 2 to 3 m. Beginning typically at the hanging wall, the fans are blasted one by one against the front-lying material, consisting of a mixture of ore from overlying slices and the waste making up the hanging wall and/or footwall. This mixture of materials will be assigned the term ‘caved material’ for simplicity but the reader should keep firmly in mind that only a portion is truly caved. Extraction of the ore from the blasted slice continues until total dilution or some other measure reaches a prescribed level. The next slice is then blasted, and the process continued. Depending on orebody geometry, the technique may be applied using transverse or longitudinal retreat. The method has been most used in the mining of magnetic iron ores since the ore can be easily and inexpensively separated from the waste. However, it has been and can be applied to a wide variety of other ore types.
Sublevel caving layout With this method the ore is recovered both through drifting and stoping. Because the cost per ton for drifting is several times that for stoping, it is desirable to maximise stoping and minimise drifting. This has meant that through the years, the height of the sublevels has steadily increased until today they are up to 30 m. Whereas approximately 25 per cent of the total volume was removed by drifting in the early designs, today that value has dropped to about five per cent in the largest scale sublevel caving designs. The sublevel intervals have changed from 9 m up to nearly 30 m. The key to this development has been the ability to drill longer, straighter, and larger-diameter holes. There are a number of factors that determine the design(Bullock and Hustrulid, 2001). The sublevel drifts typically have dimensions (W/H) of 5 × 4 m, 6 × 5 m, or 7 × 5 m to accommodate LHDs. In the example used to illustrate the layout principles, it is assumed that the drift size is 7 × 5 m. The largest possible blasthole diameter (from the viewpoint of drilling capacity and explosive charging) is normally chosen; today, this is 115 mm based largely on the ability to charge and retain explosive in the hole. The largest ring designs incorporate holes with lengths up to 50 m. The distance between slices (burden B) depends both on hole diameter (D) and the explosive used. For initial design when using ANFO as the explosive, the relationship is B = 20D. For more energetic explosives (bulk strength basis), the relationship is B = 25D. Assuming that D equals 115 mm and an emulsion explosive is used, B would equal about 3 m. Typically, the toe spacing to burden ratio is about 1.3. Hence the maximum S would equal 4 m in this case. To achieve a relatively uniform distribution of explosive energy in the ring, the holes making up the ring would have different uncharged lengths. Both toe and collar priming initiation techniques are used. The sublevel drift interval is decided largely on the ability to drill straight holes. In this example it will be assumed that the sublevel interval based upon drilling accuracy is 25 m (Figure 30). Once the sublevel interval has been decided, it is necessary to position the sublevel drifts. In this example, the drifts are placed so that the angle drawn from the upper corner of the
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FIG 30 - Step 1 in the layout of a large-scale sublevel cave - location of the sublevel drifts (after Bullock and Hustrulid, 2001).
FIG 31 - Step 2 in the layout of a large-scale sublevel cave - fan shape (after Bullock and Hustrulid, 2001).
extraction drift to the bottom centre of drifts on the overlying sublevel is 70°. This is approximately the minimum angle at which the material in the ring will move to the drawpoint. The resulting centre-to-centre spacing is 22 m. The inclination of the side holes has been chosen as 55° (Figure 31) although holes somewhat flatter than this can be drilled and charged. The function of holes drilled flatter than 70° is largely: 1.
to crack the ore, which is then removed from the sublevel below; and
2.
to reduce the maximum drill hole length.
Figure 32 shows the final front and side views of the ring design. In Figure 33, an extraction ellipse has been superimposed.
Recovery and dilution The layout shown in Figure 33 is very similar to that which would be obtained using the theory of bulk flow as described by Janelid and Kvapil (1966) and Kvapil (1982, 1992). Their design guidelines came largely from carefully conducted laboratory model studies involving the use of sand and similar materials. A point of importance is that in making such studies the granular material is poured into the models, ie no special compaction is done. This means that typically the in place percent swell is of the order of 50 per cent and it is the same throughout the model. This means that the mobility of the material, defined as the ease of movement, is constant throughout the model. With the start of draw, the material directly above the extraction point comes first. Then, since the mobility is constant, a flow body develops based upon the forces acting on the individual particles. Using this technique, the dependence of the shape of the draw body can be examined as a function of the fragment size, fragment size distribution, width and shape of the drawpoint, etc. In sublevel caving, a sublevel drift is first driven and then from this drift fans of stoping drill holes are drilled to the upper sublevels. The ratio between the ore recovered during development and that recovered by later stoping will be defined as the specific development (SD). Because of the difference in the costs per tonne for drifting and stoping one would like to minimise the specific development. When considering the layout
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FIG 32 - Step 3 in the layout of a large-scale sublevel cave - addition of the hole locations (after Bullock and Hustrulid, 2001).
FIG 33 - Superposition of the theoretical draw ellipse on the fan (after Bullock and Hustrulid, 2001).
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used by LKAB in the early-1980s (see Figure 6) it is observed that the sublevel interval(IN) was 12 m, the sublevel drift spacing (SP) was 11 m and the drift had a dimension (W × H) of 5 m × 4 m. The specific development expressed in tons/ton (t/t) can then be calculated using SD = W × H/(IN × SP – W × H)
(29)
where W
= drift width
H
= drift height
IN
= sublevel interval
SP
= sublevel drift spacing
Substituting the appropriate values into Equation 29 one finds that the specific development in the 1980s was: SD(1980s) = (5 × 4)/(12 × 11 – 5 × 4) = 0.178 t/t In applying the same approach to the design of year 2000 (Figure 3) one finds that:
FIG 34 - Influence of the free-swell volume on the ring as a function of scale.
SD(2000) = (7 × 5)/(27 × 25 – 7 × 5) = 0.055 t/t Thus there has been an improvement in the specific development by about a factor of three over this period. If blasting is done under free confinement conditions, the rock would like to expand about 50 per cent. The result is a material that is easy to dig and remove. With sublevel caving, the stoping blasting is done against caved rock lying to the front. The sides of the ring may be intact rock or caved rock. The available swell volume is largely that provided by the sublevel drift. A small amount may be available through compaction of the caved rock and in flow channels. For the purpose of this discussion, this potential source of additional swell volume will be neglected. One can write the available percent swell (per cent swell) as: per cent swell = 100 SD
(30)
Applying Equation 30 to the specific development values of the 1980s and year 2000 one finds that the available percent swell is: per cent swell (1980s) = 17.8 per cent per cent swell (2000) = 5.5 per cent As indicated, normally rock would like to expand about 50 per cent when blasted. If one allows the rock in the ring to swell 50 per cent during blasting, one can see in Figure 34 the percentage of the 1980s and the year 2000 ring that would be affected. The remaining portion of the ring would have no swell room left. If through careful blast timing one could restrict the swell to only 20 per cent, then more of the rock in the ring would have a chance to swell. In the old design most of the ring would have a chance to swell under the 20 per cent swell condition. For today’s design, it is just the lower third. The overall particle mobility is quite different in the new design than in the old. In the new design the material in the lower part of the ring has a high mobility and that in the upper part a very low mobility. Of most concern is the relative mobilities of the ore in the ring and that of the caved rock. If the mobility of the caved rock is much higher than that in the ring, even if the gravity driving force component is less, it will flow to the drawpoint in preference to the ore in the upper parts of the ring. A very extensive set of field experiments were conducted at LKAB Kiruna to study the extraction pattern of ore from the very large-scale sublevel caving rings (Quinteiro et al, 2001). The waste percentage in the extracted scoops was determined as a function of extraction. Figure 35 shows one example of the resulting curves. As can be
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FIG 35 - Typical extraction curve for large-scale sublevel caving at LKAB, Kiruna (Quinteiro, 2000).
seen in the early part of loading, the material loaded was nearly 100 per cent ore. However, at about bucket number 90 the cave began to enter. The waste percentage steadily increased and then decreased as ore from the ring came in again. Cave rock entered again followed by more ore. This pulsation pattern reflects the relative mobilities of the ore in the ring and the caved rock. In-flow of cave is required to loosen the blasted ore in the upper portions of the ring. The process is illustrated diagrammatically in Figure 36 (Larsson, 1996). A practical difficulty introduced by this behaviour is that it is extremely difficult for the loader operator to know when loading should stop. In the past one would stop loading when the waste present at the face reached a certain limit. However, with pulsation, the percentage of waste at the face can reach high values several times during the extraction process. This dilemma has been overcome by continuously monitoring the extraction of the ring via real-time weighing of the bucket. The termination decision is made based on extraction and total waste. Figure 37 shows one such curve. With strict control, one can achieve recoveries of the order of 80 per cent while maintaining an overall dilution of 20 per cent or below.
Summary The trend to increased scale in sublevel caving has many positive benefits. However in doing this, the reduced swell volume has modified the flow behaviour of the blasted rings. As a result, very careful loading control must be exercised to achieve high ore recoveries while minimising dilution.
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through an incorrect application of a mining method or an inappropriate design. The second is that ‘You can spend a lot of money trying to make a silk purse out of a sow’s ear’. The message is that today more than ever one must realistically describe the mineral reserve that one has and proceed on that basis. It may turn out that the deposit is not ore at the present time. This is nothing new. Once the decision has been made to proceed with a large-scale, highly mechanised/highly automated mining system/layout, there are few fallback positions and those are expensive. Thus it is very important to have clearly and firmly in mind the basis and limitations of the large-scale method. Today, unfortunately, unlike many other things, the selection of a mining system is not something that one can do simply by going into the World Wide Web and clicking on ‘method.’ There still are a number of gaps in our understanding which require a heavy touch of common sense and the technical feeling to bridge. It also requires engineers to evaluate and describe the deposit in ways that perhaps have not been done before. To answer the simple question of the limits between caving and non-caving is not easy. Morrison (1976) states:
FIG 36 - Explanation of the pulsation seen in large-scale sublevel caving (Larsson, 1996).
Mining methods are compromises weighted for the effects of the dimensions of orebodies, the properties of rocks and rock masses, materials and labour costs, the grade of the orebodies, and mechanical and ground control facilities. There are few absolute yardsticks associated with any of these variables and the choice of a method, in the final analysis, thus depends upon judgment. This, in turn, is best based on a perspective over a wide range of knowledge and experience. In this perspective the first priority should be given to ground control facilities upon which the success of every mining method depends. Amen.
ACKNOWLEDGEMENTS The author would like to express his thanks to: Doug Milne, Rimas Pakalnis, Glen Heslop, Dennis Laubscher, Evert Hoek, Ken Mathews, Bill Forsyth, Byron Stewart, Rudolf Kvapil, and Carlos Quinteiro for helpful discussions during the course of preparing this paper. Diane Christopherson of the Mining Engineering Department of the University of Utah was of great assistance. I have appreciated the opportunity of working with various underground mining companies throughout the years. In that regard, special thanks go to LKAB and to CODELCO.
REFERENCES
FIG 37 - Recovery and dilution curves as a function of percent extraction in large-scale sublevel caving at LKAB, Kiruna (Quinteiro, 2000).
CONCLUSIONS There is an old saying (Flavell, 1993) that ‘You can’t make a silk purse out of a sow’s ear’. Today companies must face and accept the possibility that a given deposit is, in fact, a sow’s ear. Accepting that fact, one then makes the best mine possible out of the sow’s ear. If one succeeds, it may turn out to be, in fact, a silk purse. There are some related sayings. The first that comes to mind is that ‘You can make a sow’s ear out of a silk purse’
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Kvapil, R, 1965. Gravity Flow of Granular Material in Hoppers and Bins, Int J Rock Mech Min Sci, 2(1):25-41 and 2(3):277-304. Kvapil, R, 1982. The Mechanics and Design of Sublevel Caving Systems, in Underground Mining Methods Handbook, (Ed: W A Hustrulid) pp 880-897 (SME: New York). Kvapil, R, 1992. Sublevel Caving, in SME Mining Engineering Handbook (Senior Ed: H L Hartman) 2nd Edition, Vol 2, Ch 20.2, pp 1789-1814 (SME: Littleton, Colorado). Lacasse, M and Legast, P, 1979. Change from Grizzly to LHD Extraction System, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D R Stewart) Ch 10, pp 107-118 (SME AIME: New York). Larsson, L, 1996. Personal Communication. Laubscher, D H, 1975. Class Distinction in Rock masses, in Coal, Gold and base metals of South Africa, August, pp 38-50. Laubscher, D H and Taylor, H W, 1976. The importance of geomechanics classification of jointed rock masses in mining operations, in Exploration for Rock Engineering (Ed: Z T Bieniawski) Vol 1, pp 119-28 (A A Balkema: Cape Town). Laubscher, D H, 1977. Geomechanics classification of jointed rock masses – mining applications, Trans Instn Min Metall, (86):A1-8. Laubscher, D H, 1981. Selection of Mass Underground Mining Methods, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 23-38 (SME/AIME: New York). Laubscher, D H, 1984. Design Aspects and Effectiveness of Support Systems in Different Mining Conditions, Trans Instn Min Metall, Sect A: Min Industry, 93(4):A70-A81. Laubscher, D H, 1990. A Geomechanics Classification System for the Rating of Rock mass in Mine Design, J Sth Afr Inst Min Met, 90(10):257-73. Laubscher, D H, 1994. Cave Mining – The State of the Art, J Sth Afr Inst Min Met, 94(8):270-293. Lawrence, B W, 1982. Considerations for Sublevel Open Stoping, in Underground Mining Methods Handbook, (Ed: W A Hustrulid) pp 364-374 (SME/AIME: New York). Lewis, R S, 1958. Elements of Mining, 3rd Edition, pp 404-418. MacMillan, P W and Ferguson, B A, 1982. Principles of Stope Planning and layout for Ground Control, in Underground Mining Methods Handbook, (Ed: W A Hustrulid) pp 526-530 (SME/AIME: New York.). Mathews, K E, Hoek, E, Wyllie, D C and Stewart, S B V, 1981. Prediction of Stable Excavations for Mining at Depth Below 1000 meters in hard Rock. CANMET Report 802-1571. Marklund, I, 1982. Vein Mining at LKAB, Malmberget, Sweden, in Underground Mining Methods Handbook, (Ed: W A Hustrulid) pp 441-442 (SME/AIME: New York). Marklund, I, 1995. Large scale underground mining, new equipment and a better environment – result of research and development at LKAB, Sweden, Trans Instn Min Metall, Sect A: Min Industry) 104(Sept – Dec):A164-168. McCormick, R J, 1968. How wide does a drawpoint draw?, E/MJ, 169(6):106-116. McMurray, D T, 1976. Sublevel Caving Practice at Shabanie Mine, Rhodesia, Trans Instn Min Metall, Sect A: Mining Industry, 85(Oct):A136-143. Milne, D, Pakalnis, R C and Felderer, M, 1996a. Surface geometry assessment for open stope design, in Rock Mechanics (Eds: Aubertin, Hassani and Mitri) pp 315-322 (Balkema). Milne, D M, Pakalnis, R C and Lunder, P J, 1996b. Approach to the Quantification of hangingwall Behavior, Trans Instn Min Metall, Sect A: Mining Industry, 105( Jan-April):A69-A74. Milne, D and Pakalnis, R, 1997. Theory behind empirical design techniques, in 12th Colloque en controle de terrain de l’Association Miniere du Quebec, Val d’Or, Quebec. Milne, D, Hyett, A and Bawden, W, 1998. Interpreting deformation data for support optimization. CIM 100th Annual General Meeting, Montreal. Morrison, R G K, 1976. A Philosophy of Ground Control – A Bridge Between Theory and Practice. Dept of Mining and Met, McGill University, 182 p. Mottahed, P and Ran, J, 1995. Design of the jointed roof in stratified rock based on the voussoir beam mechanism, CIM, October, 88(994):56-62.
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Newman, M F, 1999. Stability Graph Design Method – A Mining operator’s Guide, in CIM Mine Operator’s Guide, Bathurst, New Brunswick, February. Nicholas, D E, 1981. Method Selection – A Numerical Approach, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 755-754 (SME, AIME: New York). Nicholas, D E. Selection Variables, in SME Mining Engineering Handbook, (Senior Ed: H L Hartman) 2nd Edition, Chapter 23.1, pp 2051-2057 (SME: Littleton, Colorado). Nilsson, D, 1992. Surface vs Underground Methods, in SME Mining Engineering handbook, (Ed: H L Hartman) 2nd Edition, Chapter 23.2, pp 2058-2069 (SME). Obert, L, Munson, R and Rich, C, 1976. Caving Properties of the Climax Ore Body, Trans SME/AIME, 260(June):129-133. O’Hara, T A, 1970. Factors affecting feasibility of underground mines, Trans Instn Min Metall, Sect A: Mining Industry, 79(761):A54-A57. Owen, K C, 1979. Block Caving at Premier mine, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) Chapter 15, pp 177-187 (SME/AIME: New York). Pakalnis, R C T, 1986. Empirical stope design at the Ruttan mine, Sherritt Gordon Mines, Ltd, Vancouver. PhD Thesis. University of British Columbia. Pakalnis, R, Miller, H, Vongpaisal, S and Madill, T, 1987. An Empirical Approach to Open Stope Design, ISRM Proceedings, Montreal, pp 1191-1196. Pakalnis, R and Vongpaisal, S, 1993. Mine Design an Empirical Approach, in Innovative Mine design for the 21st Century (Eds: W F Bawden and J F Archibald) pp 455-467 (A A Balkema: Rotterdam). Pakalnis, R, Nickson, S, Lunder, P, Clark, L, Milne, D and Mah, P, 1996. Empirical Methods for the design of Mine Structure, 11th Colloquium on Ground Control, The Mining Association of Quebec, Val d’Or, Quebec. March. Pakalnis, R, 1998. Empirical Design Methods – UBC Geomechanics. Proceedings, 100th CIM Annual Meeting, Montreal, May. Panek, L A, 1982. Geotechnical Factors in Undercut – Cave Mining, in Underground Mining Methods Handbook, (Ed: W A Hustrulid) pp 1456-1465 (SME/AIME: New York). Panek, L A, 1979. Comparative Cavability Studies at Three Mines, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) Chapter 9, pp 99-106 (SME/AIME: New York). Parker, J, 1973. What can be learned from surface subsidence, E/MJ, 174(7):70-73. Peele, R, 1934. Discussion of the paper Selection of a Mining System by R K Warner, AIME Trans Met Min Nonmet Min, 109:22. Peele, R, 1961. Mining Engineers’ Handbook, 3rd Edition, Vol 1, pp 5-16. Pine, R J, Lay, S, Randall, M M and Trueman, R, 1992. Rock Engineering Design Developments at South Crofty mine, Trans Instn Min Metall, Sect. A: Min Industry, 101(Jan-April):A13-A22. Pettersson, A, Halonen, T and Hedstrom, O, 1998. The planning and design of the new 1000m haulage level at LKAB Malmberget. Annual SME Meeting, Orlando: Florida. 9-11 March. Potvin, Y and Hudyma, M, 1988. The Stability Graph Method for Open Stope Design. Paper 45, 90th Annual CIM Meeting, Edmonton. Potvin, Y, Hudyma, M R and Miller, H D S, 1989. Design Guidelines for Open Stope Support, CIM Bulletin, 82(926):53-62. Potvin, Y and Milne, D, 1992. Empirical Cable Bolt Support design, in Rock Support in Mining and Underground Construction (Eds: Kaiser and McCreath) pp 269-275 (Balkema). Quinteiro, C R, 2000. Personal communication. Quinteiro, C R, Larsson, L and Hustrulid, W A, 2001. The theory and practice of very large-scale sublevel caving, in International Case Studies in Underground Mining (SME) to be published. Richardson, M P, 1981. Area of Draw Influence and Drawpoint spacing for block caving mines in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 149-156 (SME/AIME: New York). Riese, M, 2000. Bucyrus International, Inc. Personal Communication. Rossouw, P A and Fourie, G A, 1996. Classification of Underground Mining Methods, Trans Instn Min Metall, Sect A: Mining Industry, 105(Sept-Dec):A162-A165. Shellhammer, D, 2000. Atlas Copco Rock Drills AB. Personal Communication.
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Shoemaker, D R, 1981. Method Selection at Questa, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 67-74 (SME-AIME: New York). Snyder, M T, 1994. Boring for the Lower K, in Eng Min J, 195(4):20ww-24ww (SME/AIME: New York). Stewart, S B V and Forsyth, W W, 1995. The Mathews method for open stope design, CIM Bulletin, 88(992):45-53. Taylor, H K, 1986. Rates of Working of Mines – A Simple Rule of Thumb, Trans Instn Min Metall, Sect A: Mining Industry, 95(Oct):A203-A204 Taylor, H K, 1991. Ore Reserves – the Mining Aspects, Trans Instn Min Metall, Sect A: Mining Industry, 100(Sept-Dec):A146-A158. Thomas, L J, 1973. An Introduction to Mining (Hicks Smith and Sons, Sydney). Tobie, R L and Julin, D E, 1982. Block Caving, General Description, in Underground Mining Methods Handbook (Ed: W A Hustrulid) Ch 1, pp 967-972 (SME/AIME: New York). Tobie, R L, Thomas, L A and Richards, H H, 1982. San Manuel Mine, in Underground Mining Methods Handbook (Ed: W A Hustrulid) Ch 3, pp 982-989 (SME/AIME: New York). Trotter, D A and Goddard, G J, 1981. Design Techniques for Sublevel Caving Layouts, CIM Bulletin, 74(825):92-100. Trueman, R, Mikula, P, Mawdesley, C and Harries, N, 2000. Experience in Australia with the application of the Mathews’ method for open stope design, CIM Bulletin, 93(1036):162- 167. Vera, S G, 1979. Caving at Climax, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D R Stewart) Ch 14, pp 157-176 (SME/AIME: New York).
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Villaescusa, E, 1996. Excavation design for bench stoping at Mount Isa mine, Queensland, Australia, Trans Instn Min Metall, Sect A: Mining Industry, 105(Jan-April):A1-A10. Wang, J, Pakalnis, R, Milne, D and Lang, B, 2000. Empirical Underground Entry-type Excavation Span Design Modification. 53rd Canadian Geotechnical Conference. Montreal, October. Ward, M H, 1979. Technical and Economical Considerations of the Block Caving Mine, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) Ch 11, pp 119-142 (SME/AIME: New York). Warner, R K, 1934. Selection of a Mining System, AIME Trans Met Min Nonmet Min, 109:11-24. Weiss, P F, Fettweis, G B, Moschitz, I M, Olsacher, A and Reidler, H, 1981. Relevant Factors for Development and Draw Control of Block Caving, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D R Stewart) pp 705-714 (SME/AIME: New York). Weiss, P F, 1981. Development System for Block Caving Under Severe Conditions, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D R Stewart) Ch 12, pp 143-146 (SME/AIME: New York,). White, T G, 1992. Hard-Rock Mining: Method Advantages and Disadvantages, in Mining Engineering Handbook (Ed: H L Hartman) Ch 21.2, pp 1843-1849 (SME). Woodruff, S D, 1966. Methods of Working Coal and Metal Mines, Vol 3, pp 383-387 (Pergamon Press: Oxford).
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The Use of Evaluation Surfaces to Assist in the Determination of Mine Design Criteria R M Kear1 ABSTRACT
CASE STUDIES
One of the problems with mining projects is the setting of the design criteria that will best match the requirements. Normally this is expressed in economic terms but can be other factors. With massive orebodies these problems are often compounded by the general lack of distinct mining boundaries. A technique using evaluation surfaces was developed at Palabora for the underground mine design. This technique has subsequently been applied on several other mining projects to assist in the determination of the design criteria. This paper discusses the method, illustrated by some of the results of these studies.
The following paragraphs are case studies of evaluations which have been undertaken and have utilised evaluation surfaces. The results discussed are only a few of the many options that were investigated.
INTRODUCTION Mining projects, both as a ‘green fields’ or an extension to present operations, normally have a long life and involve considerable sums of capital. The project itself can have a life of several years. With these considerations it is especially important to ensure that the best possible design criteria are set to ensure ease of operation and the best returns. Too often the design criteria are arbitrarily set, either as a continuation of present practice, something that has been observed on a visit to another property, or dictated by the present infrastructure. If the initial calculations using these arbitrary criteria show reasonable results the project can move into detailed design without the initial design criterion being adequately tested. A technique to assist in the determination of design criteria was developed on the Palabora underground mining project. The technique utilised evaluation surfaces and has subsequently been found to be useful on other projects. This paper gives a brief overview of this technique and shows examples of projects where its use has been of assistance during the various study phases.
EVALUATION SURFACES Evaluation surfaces consist of three-dimensional graphs showing the relationship between two variables and the resultant. This is not a particularly difficult concept and although three-dimensional matrices are subsets of multi-dimensional matrices, the multi-dimensional matrices are not easily visualised. The ability to evaluate effect of the interaction of two parameters on a resultant allows the planning team to determine the most likely combination of parameters that will give the best results. However, the determination of the results matrix requires a detailed and sophisticated model. This model then requires a calling routine to step through the variables under consideration and generate the matrices for the three-dimensional graphs. The model development and the generation of evaluation surfaces have been described in previous papers (Kear and Kirk, 1997; Kear, Fenwick and Kirk, 1996). This paper incorporates some of the results obtained on various mining projects where this technique has been utilised.
1.
Mining Advisor, PO Box 4512, Edenvale 1610, South Africa.
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Palabora The Palabora underground study investigated the possibility of extending the mine life by transferring production from the existing open pit to an underground mine. During this study it was appreciated that the design criteria would have a major impact on the economics but the problem was how to determine the best combination of these parameters. It was then that the idea of using evaluation surfaces came to the fore to ascertain the parameters that would give the best economics. From the many issues under consideration the following three are included.
The height of the cave draw column The height of the cave column determines the elevation of the extraction horizon. Variations in this elevation affected the sinking schedule, the tonnage available, draw-point repairs and dilution envelopes.
The cave grade boundary The Palabora orebody has radially decreasing grades and the cave boundary would be designed to a copper grade rather than a lithological boundary. The larger the cave footprint, the more tonnes are made available for extraction, but at a lower grade. Increasing the footprint boundary grade decreases the area and the tonnes but increases the grade. Very small footprints would limit production rates and eventually become too small for cave initiation. The project team needed to determine the boundary grade that would give the best economics.
Production rate The higher the rate of production, the shorter the mine life, with lower unit operating costs. The maximum rate of production for a particular footprint has an upper limit depending of the maximum draw rate allowed. Higher rates of production would also affect the required infrastructure.
Palabora results The results pertaining to the column height, grade boundary and production rates are shown below. Of interest was to determine the extraction level elevation or column height. At 20 000 tonnes per day the cave boundary and the column height were varied and produced the graph shown in Figure 1. In the figure the X and Y axis represent the column height and the design boundary. The Z axis is the resultant internal rate of return (IRR). This indicated that a design grade boundary of 0.9 per cent and a column height of 500 metres, would give the best economics. The graph has not peaked at a 500 metres column height but is flattening. Placing the extraction level at a lower elevation gave concern with respect to the dyke invasion from the east and also the life of the draw-points.
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FIG 1 - Palabora: relationship at 20 000 tpd between design grade boundary, column height and IRR.
A similar graph, Figure 2, was developed for a production rate of 30 000 tonnes per day.
FIG 3 - Palabora: relationship between design grade boundary, production rate and IRR at 500 metre column height.
Argyle Diamond mine The Argyle diamond pipe has weak rock which is suitable for block caving. However, the host rock is weaker than the ore, and this along with the pipe geometry presented several geotechnical challenges. The orebody is inclined and two scenarios were investigated:
• inclined draw in which the draw followed the footwall contact; and
• vertical draw as shown in the Figure 4.
FIG 2 - Palabora: relationship at 30 000 tpd between design grade boundary, column height and IRR.
Here the best boundary was indicated as 0.8 per cent contained copper with the 500 metre column height. Subsequently, a comparison was generated between the production rate and the cave boundary setting the column height at 500 metres (Figure 3). This showed that except at 0.9 per cent boundary the 30 000 tonnes per day production rate had better economics than the 20 000 tonnes per day option. The reason for the reduced economics at 0.9 per cent was the shorter mine life due to the reduced tonnes. However, the 0.8 per cent boundary at 30 000 tonnes per day had by far the best economics and these parameters were used as the design criteria. Now that the Palabora Underground mine is under construction, a similar exercise was completed to determine the most promising future extensions and expansions to the mine and the timing of these options.
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FIG 4 - Argyle Diamond mine, schematic cross-section.
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Both scenarios were modelled but for the purposes of this paper only the vertical draw is considered. One of the first issues to be examined for the vertical draw option, was the requirement or not for footwall draw-points. These draw-points would be placed in the footwall waste in an attempt to obtain the ore from higher elevations. These draw-points would initially produce waste prior to ore. This would have an adverse affect on the initial head grade but would increase the overall recovery. Figure 5 shows the effect of including one footwall draw-point, in each drive, at the column heights under consideration. The inclusion of two draw-points on each drive reduced the economics and this was not considered further during the investigation.
It could be seen that the 9.5 million tonnes per year option showed the best NPV, except at the 250 metre column height where the six million tonnes per year option showed better results. The peak NPV was at 9.5 million tonnes per year at a 200 metre column height. However, geotechnical modelling showed there was considerable risk of excessive dilution at 200 metres which was dramatically reduced at 150 metre column heights. From the graph it could be seen that the NPV reduction at 9.5 million tonnes per year from 200 to 150 metres was not large and therefore this column height was chosen. Had the NPV drop off been severe then this would have justified further geotechnical modelling.
A large copper open pit – underground investigation The mine has been successfully operated as an open pit for many years. Extensions to the present pit had been evaluated and it was required to compare these economics with those of an underground operation. Owing to the rock characteristics and the large production potential, caving was seen to be the most suitable mining method for the majority of the ore.
Orebody geometry
FIG 5 - Argyle: relationship between column height, footwall draw-points and NPV.
It was concluded that one footwall draw-point per drive had a positive effect on the NPV at all column heights. The production rate and column heights were then examined. The graph in Figure 6 shows the variation in the column height in metres and the annual production rate against the resultant NPV.
FIG 6 - Argyle: relationship between column height, production rate and NPV.
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The orebody is extremely complex with the grades being fairly widely disseminated. At the bottom of the pit the copper grades form an annulus around a barren centre. With this grade distribution it was extremely difficult to determine where the best positions were for the mining blocks. To assist in identifying the blocks the ore reserve was examined to determine the columns of ore which would pay for the operating costs and the direct capital for the extraction level and undercut. These columns were generated for a range of probable operating costs . These columns were then used as guides to blocking out for the various operating cost scenarios. At low operating costs, four potential block caves were identified. These reduced to two at the high operating costs. It was seen that two of the four potential caves had little impact on the economics and were removed from the initial evaluation with a view to re-examining their contribution at a later stage.
Results Many evaluations were required for the remaining two caves which lay in the north eastern and south eastern section of the mine. Of interest was the determination of the design boundaries for these caves. Figure 7 shows the results at 80 thousand tonnes per day production rate and shows a rapid fall off in the NPV at the higher design boundaries for the North East Cave. The surface is relatively flat around the peak values which is useful information during the final blocking out. The North East Cave was shown to be a major contributor to the economics and hence was the first cave planned to be opened. The initial period of cave production, during the ramp up, can have significant effect on the economics. Therefore, it was of interest to determine if the economics were sensitive to the cave initiation sequence. The draw program was modified to allow the cave to commence at the centre or alternatively at the North or South perimeter. This was combined with an investigation into variations in the extraction level construction rate and the results are shown in Figure 8. In this evaluation the South East Cave was fixed and the NE cave parameters varied. Initiating the cave on the South perimeter showed a lower NPV than the North and Centre options which showed a marginal improvement for the North perimeter initiation. The rate of bell construction shows improved economics as this increases. Here geotechnical input will be required to determine the fastest rate allowable.
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FIG 9 - Venetia: view to north east of the open pit geometry and diamond pipes.
FIG 7 - Copper open pit: relationship between design boundaries between two cave and NPV at 80 000 tpd.
From an extrapolation of the vertical zones of the current reserve, beyond 400 metres, studies have indicated that the open pits can be mined to depths in excess of 600 metres and still show positive economics. However, at these depths the stripping ratios become enormous and an investigation into the timing of the transfer of production from surface to underground was undertaken.
Surface to underground transition An open pit mine normally progresses through a series of pit extensions to the final pit shell. Each extension usually requires a period of waste stripping followed by the mining of the exposed ore. Therefore, once a shell extension has been started resources will have been expended for the future ore. Consequently, it is expected that an underground transition should occur at the end of a pit shell extension. A preliminary model was developed to accept the long-term open pit plans for each extension and these data were combined with a conceptual underground mine commencing production at the completion of a particular extension or cut. Figure 10 shows the effect on the mine NPV by commencing underground production at the completion of the various cuts for the K1 and K2 pipes.
FIG 8 - Copper open pit: relationship between construction rate, point of initiation and NPV for one cave.
Venetia Diamond mine The Venetia Diamond mine has several diamond bearing kimberlite pipes. Planning to-date has been concentrated towards the extraction of the K1 and K2 pipes, around which the present open pits are centred.
Open pits Figure 9 shows the planned open pits with the diamond bearing kimberlite pipes.
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FIG 10 - Venetia: effect of underground mine timing and open pit cuts on NPV.
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The rapid fall off in value with the smaller cuts is due to the hiatus in mill feed between the end of these cuts and the start of underground. It can be seen that, although the mine NPV remains positive for the largest cuts evaluated, commencing underground production after Cut 3 for both pipes produces the highest NPV. This initial evaluation was required to determine how urgent was the commencement of underground mine planning. Now that this has been determined, work is continuing on the evaluation of satellite pipe mining, alternative pit geometry and the utilisation of ‘split shells’. These studies are required to ascertain if the economics of Cut 4 could be improved and hence delay the start of underground mining until the completion of this cut.
Premier Diamond mine The Premier Diamond mine has been in production since the early-1900s. The mine has progressed from an open cast operation to underground mining, where several mining methods have been used. For the last several years the mine has successfully used block caving but is now reaching the point where the present infrastructure, especially the shafts, needs to be deepened. This has opened up several possibilities and the drop down under investigation has been named the C Cut.
FIG 12 - Premier: impact of block timing on the NPV.
It is of interest to note that this type of analysis could just as easily be applied to an existing operation as to a project.
C Cut components The schematic diagram, Figure 11, shows the major mining components of the C Cut.
Other applications There have been several occasions where evaluation surfaces have been used to determine the best fit of parameters ranked by non-economic resultants. Figure 13 shows a graph of the effect of two different mining systems feeding one recovery plant. It was required to determine what mix of these systems kept the plant at full capacity.
FIG 11 - Premier: exploded schematic view to the north west of the proposed C Cut.
The dilution comprises waste and ore that has accumulated over the years. The mine is presently producing from a portion of the pipe, which results in varying column heights for the drop down. Geotechnical input has advised that the drop down should be mined in three block caves and evaluations showed that these should progress from block 1 to block 3. FIG 13 - Relationship between mining methods and percentage feed.
Block timing One of the evaluations required was to show when the various blocks should be started. Owing to the increasing ingress of dilution the grade decreases over time. There comes a point when the following block must be started to maintain the highest NPV. The graph in Figure 12 shows the effect on the NPV of the different start dates for the two subsequent blocks after block 1.
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The graph was based on the outputs of a discrete simulation and is interesting in that there is a ‘false’ peak in the surface. This sort of result and possible discontinuities in the evaluation surface make it difficult to use numeric methods to mathematically solve the multi-dimensional matrices of results.
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CONCLUSION
ACKNOWLEDGEMENTS
It is extremely important in today’s competitive environment that the best economics for a project are determined rather than a ‘satisfactory’ return. In order to achieve this the various aspects and properties of an orebody and the associated processes must be matched to achieve the best fit. Excellent tools are available to evaluate various options. The use of evaluation surfaces assists in combining these outputs to determine the best mix and match of these alternatives to give the best results. Not only does the technique assist in determining the design criteria for further study, but also gives an indication of the sensitivity to the various parameters. This information is most useful to project management in that the ‘sharpness’ of the surface peak can easily be seen. If the peak is in a relatively shallow plateau, then there is some flexibility in changing the parameters, whereas a sharp fall off in a particular, or in all directions, will indicate the importance of those particular parameters. It is hoped that this paper has demonstrated how that this approach has been of assistance in the development of design criteria on various projects and has provided an extra tool for project evaluation.
This paper has been made possible by the kind consent of the various people involved with the projects described. In particular the author would like to thank: Argyle Diamond Mine: Gordon Gilchrist: Managing Director, Argyle Diamond Mine; Palabora Mine: Mr K Calder: Project Director; Premier Diamond Mine: Mr H Gastrow: General Manager and Mr A Croll: Project Manager C Cut; Venetia Diamond Mine: Mr S Webb: General Manager; and Mr E Harvey: Ore Extraction Manager. Discussions with colleagues and clients have been most helpful in developing the logic for project evaluation. Specifically, the input of Ken Owen and Mike Pryor during the initial work on evaluation surfaces at Palabora and continuing discussions with Frank Russell and Dr Peter Gash have been appreciated.
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REFERENCES Kear, R M and Kirk, R L, 1997. The Determination of the criteria required for the successful sizing of mine projects: Case study Palabora underground mine, in Proceedings of the Anglo American Group Mining Symposium, Johannesburg, paper 20. Kear, R M, Fenwick, F and Kirk, R L, 1996. The Sizing of the Palabora Underground Mine, in the Proceedings of the South African Institute of Mining and Metallurgy, Massive Mining Colloquium, Johannesburg. The Use of Evaluation Surfaces to assist in the Determination of Mine Design Criteria (R M Kear).
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Analysis and Management of Mining Risk J Summers1 ABSTRACT
CONTEXT AND DISCUSSION
Within the mining industry, project risk analysis has not yet found universal application because of misunderstanding of the types of risk analysis performed and confusion with the terminology. This is usually the fault of practitioners and has led to scepticism, especially in the validity of quantified results because they are not interpreted before submission to management who have little experience in their application. Because of the general lack of trust in the method there is often an unwillingness to accept risk analysis as a management tool and an erroneous belief, in some quarters, that risk analysis is of value only in safety. This paper attempts to correct some of these shortcomings describing the different application of risk analysis throughout a mining project and drawing on experience from other industries; and concludes that practitioners and managers should promote formal risk analysis out of the safety and environmental departments.
Risk and its analysis is perceived and defined in many different ways, depending on the context in which it is applied. For example, the life insurance industry uses methods based on large amounts of historical data; actuarial risk analysis. This is the first conclusion that we reach when defining risk analysis in the mining industry; we very rarely have a full set of statistical data on which to base our probabilistic judgement. Two exceptions to this statement are the cases of the ore reserve, where there is usually a body of statistical data often enhanced by specific geostatistical techniques; and geotechnical data which is extracted from borehole logs and presented as probabilistic distributions. An example of risk analysis in ore reserve estimates is given by Ravenscroft (1992) in which he describes the technique of conditional simulation to incorporate reserve variability and sampling error into mine scheduling. Perhaps the most appropriate source of inspiration for the mining industry comes from the civil engineering sphere, where risk analysis has been adopted in three main areas; operational safety, construction performance (from design to completion), and in the selection of successful tenderers. In the UK, the operational safety aspect has been driven, in part, by the Construction (Design and Management) Regulations (1995) which have made hazard identification and risk analysis an integral component of safety systems on sites; large and small. The mining regulations in the UK, South Africa, and Australia have followed this trend by introducing the requirement for hazard identification and safety risk analysis. In construction performance, risk analysis has become a tool used by engineers and contractors to determine the most reliable construction options, for instance, in the prevention of damage to existing structures from ground movements. As an example, Powell and Beveridge (1998) provide a useful summary of the design of suitable NATM methods in civil tunnelling, where they emphasise the integration of risk analysis into design decisions. The recent report of the UK Health and Safety Executive investigation into the Heathrow Express Tunnel (2000) collapse in October 1994, has highlighted that: ‘Risk assessment should be a fundamental step in the procedures adopted by all parties: it is inappropriate wholly to leave the control of risk to contractors’. The report criticises all parties in the project and describes a catalogue of errors and substandard construction over a period of three months prior to the collapse. The contractor was fined £1.2M, the designer £0.5M, and both had to pay costs of £100 000 each. Risk analysis methods applied to the selection of contractor, examines the risks to the owner from the different options proposed by contractors, and attempts to balance the relative risk of each option against the price offered. Much is written about risk sharing and risk transfer in large civil engineering projects, but the owner is always left with the difficulty of having to determine how much money to pay in order to share risk with the contractor. Obviously, the contractor is not taking the risk without imposing a cost and the owner must decide whether the price demanded (Kampmann et al, 1998) is fair for the risk being shared. This is the second conclusion in risk analysis – risk must be assessed by the party or person who is going to bear the risk and, as we shall see later, risk perception is a key element in decisions.
INTRODUCTION Project risk analysis has not yet found universal application within the mining industry despite its adoption in other, related disciplines. For many in the mining industry their exposure to formal risk analysis is confined to safety and the environmental disciplines. The aim of this paper is to remove some of the mystique from project risk analysis and risk assessment and to show that different types of analysis exist and have their place; but that the place should be defined. Misunderstandings have arisen because of the differing types and application of risk analysis coupled with the different aims and benefits of the process. Confusion is not always the fault of the mining engineer but often the fault of practitioners who fail to acknowledge that risk analysis is not a technique invented by them but is a methodology built up over a number of years from a range of different sources. Risk is present in all projects, whatever their type, and understanding and controlling risk is an essential component of project management. The key to controlling risk lies in having a clear understanding of what risk is, the risks relevant to the project under design, and the risk acceptance thresholds set by the sponsors or owners of the project. Whilst these three requirements are easy to demand they are far more difficult to implement in a real project. This paper will attempt to clarify some of the different applications in common use, describe the status quo within the mining industry, with comparisons to other industries; and highlight some limitations in risk analysis and suggest improvements. To achieve this the paper addresses the following issues in the context of a broader discussion of the application of risk analysis.
• • • •
definitions,
1.
CGSS, 9 Devereux Road, Windsor, Berkshire, SL4 1JJ, United Kingdom. E-mail: [email protected]
risk perception and acceptance, risk analysis application, and risk management tools.
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Definitions Definitions of some terms used in risk assessment are presented in Table 1. This is not a comprehensive glossary but defines terms used in this paper that have been modified from a number of different sources. Other terms will be defined as they are introduced. TABLE 1 Definition of some risk analysis terms. Term
Definition
Hazard:
A potential occurrence or condition that could lead to injury, damage to the environment, delay, or economic loss.
Risk Analysis:
A structured process that identifies both the likelihood and the consequences of hazards arising from a given activity or facility.
Risk Assessment:
Comparison of the results of a risk analysis with risk acceptance criteria or other decision parameters.
Risk Reduction Measure:
An action that could be adopted to control a risk by either reducing the likelihood of occurrence, or by mitigating the consequences of an occurrence.
Risk Management:
The process by which decisions are made to accept known risks, or the implementation of actions to reduce unacceptable risks to acceptable levels.
Perhaps the most important definitions to examine are those of ‘risk analysis’ and ‘risk assessment’ because they are often confused or used synonymously. There is an important difference between the two terms, which is more than a pedantic adoption of jargon. Risk analysis is the process by which risks are identified, examined, defined, and their magnitudes determined. Risk assessment, on the other hand, is the process by which the outcome of a risk analysis is compared to the risk acceptance criteria established for the specific project. It follows, therefore, that a risk assessment requires a clearly defined set of risk acceptance criteria that are understood by all parties to the process. Risk perception and risk acceptance are discussed later but, from the definition of risk assessment, it can be seen that two organisations could make different decisions relating to risk management simply because of differing risk acceptance thresholds. The concept of risk acceptance must always be considered when examining other people’s responses to risk. It is not the role of risk analysts to make decisions relating to risk acceptance but it is their role to help elucidate risk acceptance criteria.
Uncertainty
RISK PERCEPTION AND ACCEPTANCE Risk perception and risk acceptance are, in some senses, the opposite sides of the same coin. People are often willing to accept a high risk that they feel able to control but will not willingly accept a lower risk that is imposed, especially through government organisations. The classic example of this must relate to smoking, where the risk of death is five times higher than for coal mining (see Table 3). Risk acceptance can be difficult to determine, especially for economic risks, because of competing considerations. The literature often refers to the risks associated with a new or expanded mining project but very rarely, if ever, do we see the acceptable risk documented and used as a criterion to accept or reject a particular project. In the case of safety risk, however, the literature is well supplied with descriptions of risk to the individual, risk acceptance levels that have been determined by governments on behalf of their citizens, and risk to individuals from accident and disease, based on statistical data. Acceptance of economic risks is a decision that is usually under the control of the board of directors who are often unwilling to divulge the reasons for decisions whether or not to proceed with a potential investment. In some instances this reluctance is because the decision relates to wider business strategy rather than the perception of risk inherent in the project itself. It is usually pointless to back analyse decisions to try and identify the risk acceptance threshold within any particular organisation as many decisions will be clouded by other considerations not related to the project itself.
Risk perception
Uncertainty is the source of risk and dominates almost all forms of human activity. Confining our consideration to mine design, uncertainties can be grouped as either: i) uncertainty as to the mechanism or working of a process, or ii) uncertainty in the actual value of a parameter or property of interest. These are referred to as conceptual uncertainty and parameter uncertainty, respectively. Using examples from block caving, the ability of the rock mass to cave is governed by the prediction of a minimum required hydraulic radius of the undercut for a specific rock mass strength (Laubscher, 1994). The rock mass parameter MRMR is
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determined and plotted on the caveability relationship whereby the minimum hydraulic radius to induce caving, is predicted. For this relationship to be absolutely valid, the link between MRMR and hydraulic radius must be robust. However, the relationship is purely empirical, ie it does not follow some immutable law of physics but is based on human observations and assumptions; thus, there is a conceptual uncertainty – the relationship might not be valid for all circumstances of geology or geometry – and this constitutes conceptual risk. The measured properties of rock are often quoted as indices (eg Rock Quality Designation) and cover a range of natural variability, yet we often derive unitary values for them, such as RMR and MRMR. Clearly, the single value cannot be universally applicable across the whole region, because there will be some situations where the property lies within a tail of the distribution used to determine the ‘most likely’ value. This is parameter uncertainty – we can never be entirely certain that the representative value chosen will be correct for a given location. Risk analysis in design requires an examination of all areas of uncertainty and an assessment of the magnitude of the resulting risk. Without this, a sensible risk management process cannot be adopted. However, each mining design and mining operation is unique, and we operate in the area of ‘sparse data’ where most risk analyses are based on expert judgements, elicited from specialists and other ‘experts’, and this poses additional risk, as discussed later.
There are clear indications that many people do not fully appreciate the risks that effect their daily lives. Hambly and Hambly (1994) has estimated (Table 2) the approximate risk to life from accidents during various daily activities, compared to the risk from disease for the individual’s age group. Few people would realise that the risk of death from cycling is about ten times that of car travel, only three times that of helicopter travel, and equal to that of dangerous work. From these published statistics, we can assume that the risk from being killed working underground is equal to that of cycling to work.
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TABLE 2 Risk to life from accident and disease. Death by accident or misadventure (per million)
0.03
0.1
- At home being active
3h
10 h
6 min 3h 1h 20 m
20 min 10 h 3h 1h
2 min
- Travelling on foot beside road by bus by train by car by aeroplane by bicycle by motorcycle by helicopter - Working office work manual work dangerous - Leisure football skiing or boxing canoeing mountaneering smoking cigarettes Death by disease (per million) At age
0.3
1
3
1h
3h
10 h
6 min 2 min 2 min
10 h 3h 1 flight 20 min 6 min 6 min
10 h 3 flight 1h 20 min 20 min
3h 1h 2 min
10 h 3h 6 min
10 h 20 min
6 min 2 min
20 min 6 min
0.03
0.1
10
30
3h 1h 1h
10 h 3h 3h
10 h 10 h
1h
3h
10 h
1h 20 min 2 min 40 s 3 cigs
3h 1h 6 min 2 min 10 cigs
10 h 3h 20 min 6 min 30 cigs
10 h 1h 20 min
0.3
1
3
0 - 24
25 - 34
35 - 44
Today, many mining companies are setting goals of zero accidents at work. Whilst this is a laudable position, demonstrating a good motivational approach to safety management, are such companies claiming that they will endeavour to reduce the risk of fatality to the individual below that incurred when driving to work? Risk perception is a difficult issue to grapple with because human beings apply different criteria to determine the acceptability of risks. On a generalised level, risks to the individual can be classed as either; voluntary risk or imposed risk. The UK Health and Safety Executive (1992) has estimated that, in Britain, a 20 year old man has roughly a 1 in 1000 chance of dying within a year; for a man of 40 the chance is around 1 in 500. At 60 years old the chance of death within a year is 1 in 50 for a man and 1 in 100 for a woman. The important point to remember here is that this risk is the aggregate of all sources and in risk management we are often trying to reduce the risk from only one or two sources. Nevertheless, most people feel concern about accepting risks that are beyond their control and demand that these risks are very low compared to those that they choose to accept.
Risk acceptance The fatal accident risk (FAR) is the measure usually used to indicate the probable number of fatalities from 1000 people working 100 000 hours (their approximate lifetime). Hambly has determined the relative FAR for various activities in the UK (Table 3), both voluntary and imposed. The tolerable risk levels shaded in Table 3 have been determined by the UK Health and Safety Executive and shows that many of the activities that we accept and take for granted, such as car travel, have a FAR almost twice that experienced in the UK coal mining industry. Having gained some perspective about risk levels in general and the attitude of some regulators to the risk levels that they impose on society as a whole, we need to examine the process of
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100
300
1000
3h 1h
10 h 3h
10 h
10
30
100
300
1000
45 - 54
55 - 64
65 - 74
75 - 84
85 +
TABLE 3 Fatality accident risk (FAR) in the UK. Activity
Risk of death × 10 hours (FAR) -8
Rock climbing
4000
Helicopter travel
500
Motorcycle travel
300
High rise construction erection
70
Tolerable limit of 1/1000 per year at work (assume 2000 work hours pa)
50
Smoking
40
Walking beside a road
20
Air travel
15
Car travel
15
Coal mining (UK)
8
Train travel
5
Construction (average)
5
Metal manufacturing
4
Bus travel
1
Tolerable limit of 1/10 000 per year near a major hazard
1
Tolerable limit of 1/100 000 per year near a nuclear plant
0.1
Terrorist bomb in London street
0.1
Note: Shaded rows are tolerable risk levels defined by the UK HSE
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risk acceptance within the mining industry. It is apparent that the attitude of all responsible mining companies, in the developed world, is that serious injury and fatalities are quite unacceptable, and that every effort will be made to prevent such injuries. Unacceptability is certainly the principle where operational safety risk is concerned, and when procedures and inspections can be adopted to help manage risk but, is this really the case when major re-design of a mine is required to control risk? Under such circumstances, directors and senior management may resort to other, less onerous, acceptance criteria such as the ALARP principle (as low as reasonably practicable) shown in Figure 1, on the basis that it would be too expensive to implement all the risk management strategies required to remove the risk. ALARP may be used to justify existing levels of risk and could be considered to be an expediency applied in the absence of either a thorough investigation or a justifiable risk threshold. There is opportunity for interpretation as to what would be considered reasonably practical under the specifics of each risk, even though the ALARP principle states that all possible measures should be adopted unless the costs are grossly disproportionate to the benefits gained. Having briefly set the scene, provided a few definitions and looked at some principles of risk analysis in other industries, we can move on to examine the specifics of risk in mining projects.
The route by which a mining project develops, from the earliest scoping study to the final construction stage, reflects an ever increasing level of complexity with more and more detail added. Most mining projects go through three or four stages of analysis prior to construction: a scoping study, feasibility and investment studies, and a detailed design. The improvement in detail and understanding implicit in this progress is illustrated in Figure 2.
Scoping study
Detail
Feasibility study
Time Detailed design
Knowledge
Investment study
RISK IN A MINING PROJECT Sources of risk in a mining project are many and varied and no project is without risk. Risk can be managed, minimised, shared, transferred, or accepted but it cannot be ignored.
FIG 2 - Stages of a mining study.
Risk unjustified except in extraordinary circumstances
Unacceptable region
Tolerable only if risk reduction impracticable or if the cost grossly disproportionate to the risk reduction gained ALARP or Tolerability region where risk is accepted only if the benefits are desired Tolerable if costs of reduction would exceed the improvement gained
Risk management must ensure risk remains at this level
Acceptable region
Negligible risk FIG 1 - The ALARP principle (after HSE).
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As the project progresses through the various studies, the degree of financial risk is progressively reduced, as shown in Figure 4. However, risk is never totally eliminated and some risks manifest themselves only during the later stages of the design. To examine how risk sources can be identified throughout the design stages of a developing mine, we need to consider the life cycle of a typical mining project (Figure 3).
TABLE 4 Some sources of risk in mining projects. Life cycle stage
Risk from changes or uncertainty in:
• Concept
Resource uncertainty, Mining method, Political risk, Macro-economic risk,
Processability, J V partner(s), Collaborators, Project economics
• Design
Mine design, Process design, Infrastructure design, Licensing and permitting,
Scale of operation, Project finance, Early cashflow, Production build-up
Design
• Construction
Schedule and milestone risk, Contractor’s safety, Cost over-run,
Mineability, Processability, Rock mass performance.
Construction
• Operation and Improvement
Production schedule, Process recovery, Head grade,
Safety performance, Environmental issues
• Disposal and Closure
Waste management, Acid generation, Rehabilitation design, Rehabilitation costs,
Groundwater recovery, Post-closure maintenance, Community issues
Concept
Operation Improvement Modification
Design risk analysis Disposal & Closure FIG 3 - Life cycle of a mining project.
Data gathering is generally conducted in parallel with analysis and design, and often driven by the requirements of analysis and computer modelling. Risk exists in data gathering, as in all aspects of engineering and, because of the timing of sequential activities, managers are generally compelled to address this risk within the analysis and design process. The risks associated with the execution, or implementation of the project depend upon the analysis and design process, but cannot be analysed until after the design is complete. Based on the life cycle stages defined above, some generalised risks to most mining and recovery processes are listed in Table 4. The UK Association for Project Management (1997) states that, in general, implementing project risk analysis and management is better in the early stages of a project’s life cycle where it is more effective and useful in guiding the development of the project. Introducing risk analysis late in the life cycle is difficult and with few compensating benefits, because: contracts have been placed, equipment purchased, commitments made, reputations are on the line, and resulting management of change is difficult and unrewarding. To examine the various types of risk analysis commonly used in the mining industry, the typical life cycle of a project can be divided into five generalised groups, more or less corresponding to those in Table 4:
• • • • •
design, project economics, construction, operation and safety, and environmental and closure. These are examined in more detail later in this paper.
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The purpose of a mine design is to develop a mining, recovery, and infrastructure plan that maximises the exploitation of the resource and shareholder value, while minimising adverse impacts to the environment, the local community, and the work force. The aim of a design risk analysis is to examine the design in order to uncover and document sources of risk that could defeat the design purpose. All forms of human activity, or inactivity, contain elements of risk. In a mine design, the sources of risk include some or all of the following: 1.
optimistic or unsupported assumptions,
2.
limited or poor quality data,
3.
unjustified extrapolation of past experience,
4.
use of inappropriate computer models,
5.
models and parameters that have not been properly validated,
6.
unsuitable or misdirected ‘expert’ advice,
7.
unwillingness to acknowledge previous failures,
8.
unexpected changes in conditions,
9.
natural variability,
10. aggregation of risks, and 11. external hazards. Two of the greatest sources of design risk are the use of inappropriate models or input parameters, and undue reliance on the untested opinion of specialists and experts. Figure 2 depicts the progress of a mine design over time, showing the increase in information and data. An additional source of risk is the inability to manage the transfer of knowledge between the different stages of the process. In other words, potential problems (or risks) identified in the feasibility study are not correctly captured or addressed in following design stages and the risk manifests itself later in the mine’s life. The analysis of design risks is usually conducted as a qualitative process, where hazards are identified, entered into a risk register, and a risk index or risk rank determined. This process is discussed below in greater detail.
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Risk analysis of project economics The earliest application of risk analysis in the mining industry was in the area of financial project risk evaluation, where boards of directors and their advisors sought methods by which they could evaluate the risks of investing in certain areas (both commodity areas and geographical areas), and gain an understanding of the reliability of the estimates presented to them by project sponsors. There are a number of techniques in use to determine economic risk and contingency provision, a selection of which are discussed below. The reduction of cost uncertainty over the stages of a project can be illustrated by a torpedo model or
diagram (Figure 4). The aim of economic risk analysis is to determine, and then reduce, the range of uncertainty at each stage of design.
Contingency estimating Contingency estimating methods of risk analysis are used routinely in many mining companies to provide a justifiable estimate of the expected cost of a project (defined as the base cost estimate). These estimates are applied to both the capital cost and the operating cost estimates. The typical outcomes, or results, are illustrated in Figure 5 where there are three main elements to the final cost; the base cost, the contingency risk cost, and the risk event cost.
COST ($) Uncertain Maximum
Less Uncertain
Certain
Final Cost Most Likely
Minimum
Scoping Study
Feasibility
Detailed Design
Investment Study
TIME
FIG 4 - Torpedo diagram of cost uncertainty.
Most likely Median (central estimate)
Key BC = Base cost CRC = Contingency risk cost REC = Risk event cost
Probabilistic Estimate
$
Contingency Item:
BC
Authority Level:
Project Manager
CRC
REC
G/M
MD or Board
FIG 5 - Risk and contingency cost.
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The base cost is derived by building up a spreadsheet model of single line items each of which is represented by a point cost determined from quantity and unit cost. This is the classic estimating technique. Often the individual point costs are judged to represent a range of uncertainty and so terms such as ±30 per cent estimate for pre-feasibility study levels and ±10 per cent for detailed design studies are often used. Contingency risk cost is estimated by a variety of techniques and added to the base cost. The resulting total is usually assumed to represent the median cost, where there is equal probability of the true cost being above or below. The risk event cost is the sum of all costs associated with unexpected events such as inundation of a sinking shaft. In estimator’s language, these represent changes in the project scope, ie the estimator had not expected (or been warned of) a flood in the shaft. Figure 5 also shows typical authority levels for contingency and risk expenditure. The statement that an estimate is accurate to ±25 per cent indicates a symmetrical distribution of uncertainty – in other words, the chance that the true figure will be less than the estimate is equal to the chance that it will be greater. Experience suggests that there is a greater chance that the true cost will be higher than the estimate because costs tend to increase and problems rarely reduce the cost of a project. Coupled with the general desire of project teams to keep the estimates low enough that the project moves on to the next stage, it is probable that the uncertainty in such an estimate should be described as +25 per cent and –10 per cent. Sensitivity analyses are conducted where individual components of a spreadsheet model are varied within defined ranges and the sensitivity of the model determined. The results are usually illustrated by a spider diagram (Figure 6) to show which of the components has the greatest effect on the overall outcome. While this information may be of use in the early stages of a project, the basic assumption in the analysis, that each component will vary independently of another, is potentially a severe limitation. Stochastic modelling (Monte Carlo simulation)
can examine the numerous ‘intermediate’ cases, where individual components in the model vary at the same time, thus overcoming the idealised assumptions. This technique is described below.
Project valuation Project valuations are determined using net present value (NPV) and the internal rate of return (IRR) of a project. Both are calculated from a cash flow model constructed to represent the expenditure and income stream for the life of the project. Such a prediction is subject to varying degrees of uncertainty depending upon the type and location of the project, and the commodity being mined. It is generally accepted that an expansion project has less uncertainty than a new project, where the location imposes political and macro-economic risks. Furthermore, the volatility of the price of the commodity also influences the perception of the risk, depending on who is judging the risk. In the past, the usual method of assessing these uncertainties has been to increase the discount rate applied so that for projects with a perceived higher risk the discount rate is progressively raised for each adverse factor, which reduces the NPV making the project less attractive. With the arrival of tools capable of modelling uncertainty there has been a trend to explicit modelling using stochastic simulations (Monte Carlo modelling) of NPV calculations. Some modellers have argued that the ‘explicit’ modelling of risk in stochastic simulations, means that lower discount rates should be used. Davis (1995) has refuted this argument and shown that no amount of analysis can reduce the risk to an investor. The only way that investor risk can be reduced is by action, such as hedging the price of the commodity. A quantified risk analysis of a project economics has a number of advantages, as expounded by Davis. Firstly, when the correct discount rate is applied, simulation gives a better estimate of the project NPV than does a deterministic analysis because the uncertainties in the model may be non-normal or correlated. Explicit modelling of these distributions gives a better understanding of the expected NPV.
Change in NPV (%)
40
Operating Cost Exchange Rate
20
Fuel Costs
0
Commodity Price Labour Rate
-20
-40 -40
0
40
Change in Variable (%) FIG 6 - Spider diagram.
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Secondly, the variance in the modelling yields an uncertainty around the expected NPV value from which the probability of the true NPV being greater than a nominated value, can be determined. For example, the probability that the project will have a positive NPV can be calculated. Thirdly, uncertainties in a number of project NPVs can be compared and choosing the preferred option may be easier if the breadth of uncertainties are known for options with essentially identical NPVs.
Process and petro-chemical companies have developed these methods to a high level of sophistication which incorporate risk acceptance criteria. The user can see immediately if the risk level determined is above or below a response threshold and is often directed to the sort of response that is applicable. Such methods work well in the relatively constant environment of a process or petro-chemical plant but are not always well suited to the variable conditions in a mine.
Construction risk analysis
Environmental and closure risk analysis
Construction risk analysis includes a number of related risk disciplines, such as safety and economic risk which are addressed elsewhere in this paper. The issue of concern here is that of schedule risk, our ability to predict the construction time and, more importantly, the point at which the cash flow becomes positive. Many of the techniques available to the mining industry have been borrowed from project management in general and civil construction in particular. In evaluating the risk to estimated project construction time, two main sources of uncertainty become apparent; firstly, our ability to determine the duration for each of the tasks in the schedule model, and secondly, the degree of confidence that the resulting critical path is correct. The critical path risk is important because project managers appreciate that the construction duration can be controlled only if the critical path items are well managed, thus these items receive greater attention than non-critical path items. However, a minor hold-up in part of the project may change the critical path and brings other tasks, which may have received only minimal management attention, onto the critical path. In large projects, the planning software often contains an uncertainty module which allows the duration (and resources) for each task to be specified as an uncertainty distribution. For smaller projects, third party tools are available, such as Risk+ (Risk and Program Management Solutions, Inc), an add-in for Microsoft Project, which provide the capability to specify uncertain ranges for each task. The output of these types of program are uncertain completion dates for the project and individual milestones, plus an estimate of the probability that a particular task may be on the true critical path. The latter output is, perhaps, the more important result from the point of view of the project manager.
Closure risk and environmental risk analyses are very similar techniques; the difference being that closure analysis is directed to the longer term issues of mine closure and post closure maintenance. Environmental risk in the mining industry can be classed as either; health risks (physical, chemical or biological) or ecological. Historically, the mining industry has dealt with health risks in conjunction with safety but more recently, with the advent of environmentalists on the staff of mining companies, health risks have become a bridge between safety and the ecological environment. In the future, as safety issues are examined in greater detail, it seems likely that health and ecology will become more closely linked. Environmental risk assessment methodology is discussed by the European Environment Agency (1998) who summarise the steps as follows:
Operational and safety risk analysis This is, perhaps, the area of risk analysis that is the most familiar to mining engineers. The fundamental requirement for any risk analysis is the identification of a comprehensive and exhaustive set of hazards. Without such a list, it is unlikely that the risk analysis will be judged to be a success, especially if an unidentified risk occurs. The techniques used by safety analysts include hazard and operability studies (HAZOPs) hazard analyses (HAZANs) and job hazard analyses (JHAs). All of these are very similar in approach and rely on systematic techniques to identify specific hazards. Identification is usually conducted under the assumption that the task will be performed using established practices and methods. It is usually assumed that an incident will result in an injury or a fatality and so the definition of consequences of an event is more limited than in other forms of risk analysis. In order to determine the level of risk the concept of ‘exposure’ is introduced and a mathematical relationship is determined, or assumed, to calculate the risk. The relationship may be non-linear (Eisenberg, 1975) of the form cmtn where c is consequence and t is exposure time. In most applications, the mathematics is incorporated into a series of charts where the user selects the consequence level and exposure deemed appropriate and reads the risk index from a table.
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• • • • • •
problem formulation, hazard identification, release assessment, exposure assessment, consequence assessment, and risk estimation.
The problem must be clearly defined and include a specification of the source. For example, is it a single chemical, what transport mechanisms are involved or is it a disposal hazard? Additionally, the regulatory framework must be understood including the acceptable limits, any licensing limits specific to the site, and any special aspects of the end-point recipient (is it a children’s hospital?). Hazard identification follows, to determine the mechanisms and circumstances under which a release could occur and the receptors of concern, which could be flora, fauna, people, water courses, etc. The identification should also examine the methods by which the release could occur and the conditions necessary for it to happen. A release assessment is the study of the potential for a risk source to release the hazard into the environment and this is often a probabilistic analysis. The exposure assessment involves the quantification of the intensity, duration and frequency of exposure of the receptors of interest to the hazard under examination. A consequence assessment will quantify the effects of the release on the population of the receptors being examined. For human health, the consequences examined are usually death and illness. The data examined cover toxicity levels, epidemiology and modelling, such as dose-response predictions. All these individual strands of the study are brought together to assess the overall risk from the defined hazard to the specified receptor group. This could be an estimate of the number of people likely to experience health effects over given time periods, from releases within the likely and the less likely magnitude ranges. In other words, the effects on people of ‘routine’ releases or licensed discharges and the effects of larger, accidental or uncontrolled releases.
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This type of environmental risk analysis is equally relevant to operational conditions, such as routine cyanide discharges, and to post closure releases such as acid drainage from waste rock dumps.
RISK ANALYSIS METHODS All risk analysis methods can be classed as either quantified, using probability distributions within numerical models; or subjective methods which usually use classification systems to derive a risk index or risk rank. Although quantified risk analysis is used extensively in certain industries, and has a place in mining, it is common to find that a subjective analysis is conducted as a pre-cursor to determine the major influences and issues to be examined. Often quantification of risk is a time consuming process and, unless sufficient justification for its use can be found from the outcome of a subjective analysis, considerable time and resources can be wasted.
Criteria for Risk Assessment
Conceptual Model of Project
Subjective and quantitative risk analysis employs a variety of tools, some of which have been developed specifically and others that have been adapted from other disciplines. The risk analysis literature contains numerous examples describing the tools that can be applied, especially to safety risk investigations. Computer based systems are also available but care must be taken in their application to guard against the analysis being driven by the capability of the technique used. The stages of a typical risk analysis are shown in Figure 7, which illustrates the relationship between a subjective and a quantified analysis.
Subjective risk analysis There are a range of techniques used by practitioners in the execution of a subjective risk analysis and, because a subjective analysis often precedes a quantified analysis, such techniques have indirect application to probabilistic analyses.
Identification of Unwanted Events
Processes Leading to Events
Factors Leading to Accidents Identify Risk Reduction Measures
Quantitative Determine Consequences of Occurrence
Type of Analysis ?
Qualitative
Determine Probability of Occurrence
Assess Likelihood of Occurrence
Determine Dependencies, Relationships
Assess Consequences of Occurrence
Determine Project Risk
Assess Overall Risk
Risk Acceptable ?
YES
Proceed With Current Plans & Schedules
YES
NO
Risk Acceptable ? NO
FIG 7 - Risk analysis flow diagram.
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The main stages of a risk analysis can be summarised as follows:
• • • •
problem definition, identification, risk determination, and risk management.
Each of these is examined in a little more detail. The main output of a subjective risk analysis is a risk register that documents each hazard. The risk register is discussed later in more detail.
Problem definition The definition of the scope of the analysis is a prerequisite to ensure that all parties understand the issues to be examined and have a clear definition of the expected outcome. Some of the tools available to gain an overall understanding of the components and structure of an underground mine design are: 1.
Concept Map
2.
Influence Diagram
3.
Logic Tree
- a map of the main stages, phases or controls of the extraction cycle - an unstructured representation of links and dependencies between properties, processes and design controls
- a hierarchical presentation of the properties, processes and design controls in the extraction Of these, the concept map and the logic tree are the two that tend to be used most extensively. Influence diagrams are mostly used to collect thoughts and ideas (brain-storming) which are transformed into the more structured representation of a concept map and a logic tree. Using the concept map and the logic tree, the problem can be divided into manageable sections.
Identification Identification is the most important component of risk analysis; a risk that has not been identified cannot be ranked or adequately managed. Figure 7 shows the relationship between identification of the hazards and potential risk reducing measures. There are a number of techniques used to aid identification exercises, some of which are complimentary; as follows:
• • • • • • • •
HAZOP, brainstorming and what-if, HAZAN,
Risk determination Risk is determined from the relationship between likelihood and consequences. Mathematically, this is expressed as Risk = Likelihood × Consequences, where likelihood is expressed as a probability. In a subjective analysis, likelihood and consequences are assessed according to suitable classifications such as those shown in Tables 5 and 6.
Checklists, TABLE 5 Likelihood classification.
failure mode and effect analysis (FMEA), fault tree analysis (FTA), event tree analysis (ETA), and
Likelihood
task analysis.
HAZOP (hazard and operability) is widely used in process applications and so has a place in mining process plants but can also be used as a routine ‘toolbox’ procedure to examine potential safety shortcomings, especially in unusual tasks such as maintenance interventions. Brainstorming takes a variety of forms, and is essentially a structured forum where participants examine a system, or subsystem, to identify potential hazards. Brainstorming sessions work well only if guided by a facilitator, knowledgeable in the specifics of the system under examination. HAZAN (hazard analysis) is similar to HAZOP except that it tends to be carried out by specialists external to the immediate
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organisation, using experience brought from outside and from relevant historical data, such as accident statistics. Checklists specify known problems relevant to a type of plant and are designed to encourage designers to address known risks. The technique is similar to HAZAN in that it relies on industry wide information and experience from other locations. Failure mode and effect analysis (FMEA) is a method of examining how a component or system can fail or be incorrectly operated or used. This technique is more suitable for process plants where the analysis can systematically examine every component of the system. It has application in mining where technology levels are high, such as hoisting. Fault tree analysis (FTA) is used to examine the events and conditions necessary for a particular hazard to develop. The incident of interest (often called the top event) is de-convolved and the individual causes or conditions required are determined. Each cause or condition can be assigned a probability from which the total probability of the top event can be calculated. The system lends itself to computerised analysis and has become more popular in recent years because computers have made it easier to apply. An event tree analysis (ETA) examines the potential string of consequences from a particular event, such as the failure of a pipeline. It is similar in concept to an FTA, except that it examines the likely outcome of the top event. Event tree analyses are used in environmental studies to examine the effects of a system failure that leads to a release. As with the FTA, it lends itself to probabilistic modelling and, thus, to computerisation. Task analysis is used to examine the human contribution to a system or operation. Human error, or failure to perform as expected, is a major cause of accidents large and small, Chernobyl is a particularly well known example of the potential for humans to behave irrationally. Each task is examined to determine where breakdowns might occur and experience from related sites may be used to focus participant’s attention. Some or all of these techniques are used by the proficient facilitator to develop a comprehensive list of hazards and possible risk reducing measures that becomes the basis for the subjective risk analysis.
Very Unlikely
Unlikely
Probable
Highly Likely
Descriptive
Almost Impossible
Possible Sometime
Isolated Incidents
Repeated Incidents
Frequency Interval (Multiple events)
Within 20 years
Within 5 years
Within 1 year
Within 6 months
Probability (Single events)
< 1/2000
1/2000 to 1/100
1/100 to 1/10
> 1/10
The level of risk is then determined either as a risk index, or from a matrix similar to that shown in Table 7.
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TABLE 6 Consequence classification. Consequences Very Low
Low
Moderate
High
Environmental Localised Widespread Severe Catastrophic Impact Degradation Degradation Degradation Degradation Personnel Safety
No Injuries
Minor Injuries
Serious Injuries
Fatalities
Lost Time (shifts)
0
0 to 500
500 to 6000
> 6000
Operating Cost
< $A 0.5 M $A 0.5 M to $A 2.5 M to $A 2.5 M $A 10 M
> $A 10 M
Ore Milled (tonnes)
< 30 000
30 000 to 200 000
200 000 to 500 000
> 500 000
Total Mined (tonnes)
< 200 000
200 000 to 1M
1 M to 2 M
>2M
TABLE 7 Risk determination matrix. Most serious consequence Very Low
Low
Moderate
High
Very unlikely
Level 1
Level 2
Level 3
Level 4
Unlikely
Level 2
Level 3
Level 4
Level 5
Probable
Level 3
Level 4
Level 5
Level 6
Highly likely
Level 4
Level 5
Level 6
Level 7
QRA allows the user to specify each uncertain variable as a range of values, defined by a probability distribution. Monte Carlo based QRA has gained a broader application in the past decade because of the availability of powerful computers and the universal adoption of the spreadsheet as a basic modelling tool. Any problem that can be established as a deterministic spreadsheet model can be examined as a QRA, however, considerable thought and skill is required to derive a justifiable and meaningful model. Simply to accept the software’s default triangular distribution of ±10 per cent of the most likely value as the upper and lower limit for each variable; which results in a solution showing a ±10 per cent range, is an exercise in futility. The uncertain distribution used for each variable must be clearly thought out and the selected input justified. Most modelling tools, available as add-ons to Microsoft Excel, offer a range of stochastic distributions to describe uncertain parameters some of which are well known to most mining engineers. One of the most important issues in any uncertain input is the shape of the distribution. Too often, inexperienced users assume a ‘plus-minus’ uncertainty assigning an equal probability that the true value will be greater than, or less than the most likely estimate. This is often not the case; capital projects do not under-run with the same likelihood as they over-run. Vose (1996) provides an excellent description of QRA, and summarises the benefits of Monte Carlo simulation as follows:
• the distribution of the model’s variables do not have to be approximated in any way;
• correlations and other inter-dependencies can be modelled; • the level of mathematics required to perform a Monte Carlo simulation is quite basic;
• the computer does all of the work required to determine the outcome distribution(s);
• software is commercially available to automate the tasks involved in the simulation;
• greater levels of precision can be achieved simply by
Risk management The risk analysis process in Figure 7, shows a loop around a decision point where the user determines if the risk is acceptable. Risk acceptance is discussed elsewhere in this paper and, assuming that a suitable set of risk acceptance criteria have been established, the risk manager is able to determine if risk reduction measures need to be implemented. Risk management is the process by which conscious and informed decisions are made to accept known levels of risk or to implement a set of actions to reduce the unacceptable risks to acceptable levels. Once the risk is reduced to an acceptable level, possibly using the ALARP principle shown in Figure 1, the risk management process ensures that the risk remains beneath the acceptability threshold. This is often achieved by the use of inspections, audits, and the examination of formal risk indicators.
Quantified risk analysis Quantified risk analysis (QRA) uses Monte Carlo simulation and is related to ‘what-if’ scenario modelling. However, QRA goes further than simple what-if because it examines all possibilities within the ranges specified; whereas what-if scenarios examine only a limited number of points and interpret between them. What-if scenarios give equal weight to each scenario and are unable to estimate how likely one scenario is compared to another. In addition, what-if models deal only with explicit, pre-determined combinations and cannot examine all the possible interactions between a number of uncertain variables. Monte Carlo based QRA examines all possible combinations of interactions between input variables, and faithfully reproduces the ranges of uncertainty specified in the input.
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increasing the number of iterations performed;
• complex mathematics (power functions, logs, IF statements, etc) can be included with no added difficulty;
• Monte Carlo simulation is widely recognised as a valid technique and so the results are more likely to be accepted;
• the behaviour of the model can be examined with ease, and • changes in the model can be made quickly and the results compared with previous models. The results of a Monte Carlo simulation are illustrated in Figure 8a and 8b showing some of the main features of the output.
RISK MANAGEMENT TOOLS Risk can be eliminated, transferred, shared, reduced, or retained; but it cannot be ignored. There are a number of different techniques for managing risk some of which overlap: 1.
Eliminated. The cause of the risk can be removed completely by banning, such as the use of aluminium in coal mines; or in design by choosing to eliminate a method of working, such as hand raising in stopes. The mining industry has used both methods. One of the drawbacks of risk elimination is that other materials and methods must be available for substitution and they, in turn, carry their own risks.
2.
Transferred. Two methods of risk transfer typical in mining are risk sharing arrangements with contractors, or the purchase of insurance. The limitations of both of these
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methods is that the outcome of any incident can never be determined in advance as contract conditions are always liable to dispute and interpretation. With insurance policies, claims are only interpreted by the claim assessor in the light of the actual incident. 3.
such as agreements with contractors; and in others it is implicit, such as where royalty payments are based on profit or revenue formulae. 4.
Shared. In mining, risk sharing occurs in; i) joint venture agreements covering the running of an operation, ii) with contractors in the execution of certain works, or iii) with the authorities in the establishment of formulae for royalty payments. In some instances the risk sharing is explicit,
Reduced. This is usually often the manner in which most mining risks are managed. The risk is analysed and feasible risk reducing measures identified. The risk management process involves selecting measures that are cost effective and able to reduce the risk below an acceptance or manageability threshold, or tolerable under the ALARP principle.
Overlay Chart Project Value .029 .022
Lognormal Distribution 10% = 29.00 90% = 47.00
.014 .007
Total Cost
.000 15.00
26.25
37.50
48.75
60.00
FIG 8a - QRA output showing overlain Lognormal distribution.
Forecast: Total Cost 2 500 Trials
Frequency Chart
5 Outliers
.029
72
.022
54
.014
36
.007
18 Mean = 37.67
.000 15.00
26.25
37.50
0 48.75
60.00
Certainty is 35.12% from 40.00 to +Infinity $ FIG 8b- QRA output showing probability that Total Cost will exceed $40.
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5.
Retained. Risk acceptance can occur knowingly or unknowingly. Known risks are usually accepted after a process of risk analysis and risk reduction so that the retained risk is assumed to be manageable with the resources available. Unknown risks are another matter entirely; these usually develop in the absence of a thorough or systematic risks analysis and, depending on the nature of the risk and the size of the incident, could be extremely damaging to company’s reputation or financial security. Sometimes known risks are under estimated and erroneously accepted. Usually there is some aspect of the risk that was unknown and led to the incorrect decision to accept the risk.
Risk management has been described by the Royal Society (1992) according to what it does rather than what it is; using the three basic elements of organisational control: i) setting of goals, ii) gathering and interpretation of information, and iii) actions. This philosophy has been extended to form the basis of the operation of ISO14001 (1996) where the cycle is summarised according to the five basic actions shown in Figure 9; policy, planning, action, checking, and review. This cyclical approach to risk management embodies continuous improvement and depends on the updating of the risk analyses.
Discussion In their publication on risk, The Royal Society presents a useful appraisal of the debate between practitioners, academics, and regulators governing the general identification and management of risk. Following the structure of the Royal Society’s discussion, risk management approaches can be examined under seven headings: 1.
anticipation,
2.
liability and blame,
3.
quantitative risk analysis,
4.
corporate response,
5.
cost of risk reduction,
6.
levels of participation, and
7.
regulatory targets.
Anticipation Risk anticipation relies on systems of detection and warning. In the civil engineering industry this is termed the ‘observational method’ where complex instrumentation is used to signal a deterioration in some measured condition. In some circumstances there is a clear application for this method, for example in the case of cyanide discharges to tailings dams where routine monitoring of cyanide levels, coupled with a cyanide balance within the plant, are capable of detecting abnormalities in the process. Advocates of the observational approach use disaster investigations, where it is often shown that an accident was waiting to happen, to show that by measuring and monitoring the key factors, risks can be adequately managed. The anticipatory approach relies on a number of parallel requirements. Firstly, it is necessary to monitor the right indicator. While this may be self-evident, in practice it does not always happen. For example, monitoring of cyanide in the tailings discharge line may be too late in the process to enable adequate and timely action to be taken to prevent a breach of discharge permits. Secondly, it is necessary to set alert levels low enough that action can be taken in sufficient time before a dangerous condition develops. Thirdly, a pre-determined response plan is required. To be aware that the system has gone wrong but unable, through lack of procedures or lack of training, to respond immediately invalidates even the most sophisticated real-time monitoring system. In summary, from a mining point of view, anticipation is potentially a useful approach, but it needs to be reinforced by a well designed monitoring and action response procedure.
RISK POLICY Intent Principles Acceptance criteria
REVIEW
PLANNING
Independent Audit Internal Audit
Risk Analysis Risk Assessment
CHECKING Safety inspections Accident analysis Health measurement Environmental monitoring
IMPLEMENTATION Risk elimination Risk reduction
FIG 9 - Risk management cycle (after ISO14001).
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Liability and blame This approach advocates the use of incentives or punitive actions to ‘encourage’ compliance with rules. At a societal level, many of our laws are founded on this principle – that punishment will follow inappropriate actions. Within mining companies this is often the situation where individual employees are liable to instant dismissal for breaches of safety codes and standards. Blame allocation has a number of limitations. Firstly, it tends to focus the investigation on the immediate actions leading up to the incident, especially in the event of an accident, and the so-called ‘hunt for the guilty’. This often leads to underlying shortcomings in the overall system remaining undetected, or deliberately hidden, and blame apportioned to the immediate supervisor. Secondly, this approach is in direct conflict with the principle of a no-blame reporting system, where individuals are encouraged to report near-misses and to own-up to ‘negligent’ actions. Thirdly, at a corporate level, where regulations are framed to provide financial or other sanctions against the individual or the organisation; investigations after incidents are stifled because of the risk of self incrimination. There are arguments that a clear investigation as to the real cause of major incidents can be defeated by the requirements of individuals to protect themselves at all costs.
Quantitative risk analysis There is considerable debate as to the extent to which risk management should rely on quantification of the risk as opposed to subject (qualitative) risk analysis methods. In the past, quantification of risk in the mining industry led to a reticence by managers to adopt the methodology partially because they were unable to interpret and apply the results of the analysis. Quantification requires the ability to compare quantified risk against pre-determined acceptance thresholds. However, as discussed previously, in the mining industry we often work in a situation of sparse data which makes quantification of risk as a routine process difficult, if not impossible. This is not to say that quantification does not have its place – it does, especially in economic risk evaluation but it requires intelligent interpretation of the results so that management can make correct decisions having clearly understood the meaning of the results. In summary, within the mining industry, quantitative risk analysis has a place in economic evaluation of projects, but the lack of applicable data and the inherent differences in operations between mine sites makes its universal application difficult. The need to understand the causes and sources of risks strongly suggests the need for a subjective analysis before a quantified analysis is contemplated.
Corporate response Advocates of the corporate response argue that there is a knowledge base covering risk management within the confines of good practice in the corporate environment. This approach has led to initiatives borrowed from the petro-chemical and process industries, such as written commitments to safety and environmental behaviour made at the highest levels. The result is initiatives such as zero accident targets across organisations, with internal and external audits to enforce the action. This ‘safety culture’ has been a focus of mine safety for a number of years and has worked its way downwards through the operating companies with total quality management (TQM) principles in middle management and tool box safety huddles at operator level. While there have been clear benefits in some organisations, some international companies find great difficulty in applying the corporate directives to an equal standard irrespective of the country of operation. Thus, it is easier to apply these techniques in Australia or the USA, where the safety culture is already
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amenable to such methods, than it is in some developing countries, where the safety culture is not so well developed or where there is an attitude that accidents reflect God’s will. There is one area of corporate response where some mining companies are probably deficient; that of minimum design rules or standards for their operations. This approach is, in general, to be encouraged but only if the rules are well thought through, specify minimum performance requirements rather than prescriptive designs, and do not exonerate the designer from adequate consideration of the conditions and allow the unthinking application of a ‘code’.
Cost of risk reduction Perhaps the biggest issue in mining risk management is the cost of risk reducing measures. At its lowest level, this issue hinges on the risk acceptance criteria applied by the operating company or the main organisation. In some areas it is reasonably simple to apply, for example zero fatalities, but there is a financial cost above which no manager will incur expenditure. For example, if the risk of fatality to each loader operator from falls of ground has been calculated as very low yet is deemed to be unacceptable; and if the only feasible risk reducing measure is to introduce remote loader operation at a total capital cost of A$10M, there will be considerable debate about the cost-benefit of such action. Despite the fact that there is a real probability of a fatality in the mine, management will probably be very reticent to commit itself to such a level of capital expenditure. If a fall of ground did occur and led to the fatality it would be hard to argue, in this example, that the fatality could not have been foreseen and that it could not have been prevented. This leads to the principles of BATNEEC (best available technology not entailing excessive costs) and ALARP (as low as reasonably practicable) which allow risk managers to trade-off the apparent risk benefits against the costs involved, rather than forcing them to apply corporate accident targets directly. In many cases these risk acceptance principles are justifiable, especially in the absence of more formal acceptance criteria. However, they can be misused to minimise the costs of risk reducing strategies and to avoid applying more costly risk management rules set at a corporate level. In summary, the application of sound risk management practices across an organisation usually indicates good business practice with links between expenditure on risk reduction and good training, management and operational practices.
Levels of participation The size of the groups involved in risk management can be debated on a number of planes. At a societal level, it is argued that as broad a range of the ‘peer community’ will result in improved risk management. The example used is that amateur ornithologists first detected the decline in the populations of the peregrine falcon, resulting in the appreciation of the detrimental effects of DDT in the environment. Within a mining operation, the issues are the involvement of the work force, especially safety risk management, and the involvement of the immediate community in decisions that have the potential to effect their lives. The involvement of the work force in safety risk assessment is now well established on both the formal and the informal levels, with the introduction of hazard analyses (HAZAN) as part of the general work procedure and the implementation of safe operating procedures (SOPs) governing activities that are considered to be hazardous. The involvement of the wider community around the mine, as well as other stakeholders is not so well established. Some mining companies have suffered significant costs and delays due to the interference of these groups and, although their involvement may have been of overall benefit to the operation,
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community and environment, their contribution is not always acknowledged by management. In general, it is probable that wide participation and consultation leads to the best risk management because of a perception of involvement and ownership. Ownership of the solution is often as important as ownership of the risk.
TABLE 8 Contents of a risk register. Topic
Description
Hazard
Definition and description of the unwanted event, condition or threat.
Regulatory targets
Likelihood
Assessment of the probability of occurrence.
At a higher level in society, discussion of regulation as a method of risk management is valid. For example, many countries now have rules governing the shelf life of perishable foods to minimise the risk of widespread food poisoning. These regulations are based on an analysis of the probability of food poisoning increasing as the storage time increases but, a number of simplifying assumptions need to be made including the method of storage and the state of the ingredients before they were packed. No amount of regulation governing the shelf life of food will compensate for the initial ingredients being tainted. In the mining sphere, regulations have changed over the past decade moving away from a prescriptive set of minimum standards to a position of responsible action on the part of the appointed person, the manager, and the mining company. It is probable that this approach to regulation will continue into the foreseeable future with the possible exception of the introduction of regulatory targets within a probabilistic framework. In Australia, for example, there is considerable discussion regarding the definition of an acceptable probability of failure of pit slopes. The drawback of this approach is that the risk is a product of both the probability of failure and the consequences of that failure and so the regulations can only be defined within specific scenarios, such proximity to an active haul road.
Consequences
Assessment of the most severe expected consequence of an occurrence.
Implementation
Risk Level or Index Determination of the risk based on likelihood and consequence. Project Impact
Whether this is a risk of a deleterious effect or if there is upside potential which can be viewed as an opportunity.
Other Impacts
The objective or milestone effected by the risk or other areas impacted by the hazard, e.g. environment, mine closure, or underground extension.
Risk Manageability The ability of the organisation to manage the risk. For example, earthquakes are an unmanageable risk. Risk Indicator(s)
Physical or other measurements, or observations, that might signal that this hazard is about to develop.
Risk Reduction Measures
Possible measures that could be adopted to reduce the likelihood of occurrence or mitigate the consequences if the hazard occurred.
Risk Ownership
The individual or department who is responsible for managing the risk. In some cases, this might be a later stage of the project, e.g. feasibility design.
Secondary Risks
Additional risks that might arise from the risk reduction measures identified.
There are three main tools suitable for the implementation of risk management in mine design and operation, assuming that a risk analysis has been, or is about to be, conducted: i) risk register, ii) action management, and iii) risk updates.
1.
identify the unacceptable, or the highest priority, risks to be tackled first;
Risk register
2.
agree a set of actions to address the risk;
The risk register has been described briefly as a component of subjective risk analysis and was described as the main deliverable. To be of maximum value, the risk register should contain at least the topics defined in Table 8. Without the acquisition of this level of information for each risk, a subjective risk analysis is incomplete and not capable of leading to effective risk management. The ability to complete entries for the topics in Table 8 for each hazard, indicates that the risk analysis has been thorough and that each risk is understood. Risk analysis is not a goal in itself, it is a tool and a pre-requisite to good risk management. Without the ability to demonstrate an understanding of the risks, the ability to succeed in risk management becomes much more difficult.
3.
provide resources and a schedule to accomplish the remedial actions;
4.
manage the implementation of the remedial actions; and
5.
monitor the remedial actions to verify that the desired outcome has been achieved, ie that the risk has, in fact, been reduced.
Action management Once a risk register has been constructed, and the levels of the individual risk determined, the requirement is for action to reduce the unacceptable risks to acceptable levels. The risk register documents the potential risk reduction measures that can be adopted so that the manager has the option to implement one or more of these or to identify other actions. These actions require both tracking and management. ISO 14001 (Figure 9) provides an outline of how such management fits into the overall scheme of risk management. Expanding on the principles in ISO 14001, the following sequence of tasks can be derived:
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The management of the remedial actions to ensure that the outcome of the risk analysis is implemented becomes the most important aspect of risk management after the risk analysis itself.
Risk updates The last of the three main risk management tools most applicable to the mining industry is the risk update. Risk analysis and management is a continuous process by which risks are identified and reduced; and other risks are identified. Changes in operations – especially through continuous improvement – mean that all risk analyses need to be updated on a regular basis. The risk update serves two main purposes:
• to ensure that the risk register is updated considering the latest operational and management controls that are in place; and
• that new risks are identified and dealt with. Both of these tasks can be viewed as an update of the risk register itself and the needs for an update becomes pressing when the risk register no longer contains the latest information.
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Although it may seem self-evident that risks analyses should be updated, it is not always the case that such updates are conducted as a routine process. Despite the lack of a risk update, managers will often claim that they are using risk management techniques routinely across their operations.
BENEFITS AND LIMITATIONS OF RISK ANALYSIS The Association of Project Management has defined the benefits of risk analysis as either:
• hard – contingencies, decisions, controls, statistics, indicators, etc; or
• soft – people related issues, attitudes, commitment, etc. Not all projects will accrue all the benefits listed, but most will gain the advantage of at least half of those listed below. Hard benefits can be listed as follows:
Transfer of risk ownership There is an ever present danger that responsibility for managing risk will be transferred to either the risk analysis process or to the risk analyst leading the project. After the risk analysis itself, risk ownership is one of the basic requirements for effective risk management.
Validity fades with time A risk analysis is not a one-off project. Risk profiles change over the life of a project and, in rapidly developing projects, frequent updates of the risk analysis are required. The risk register should be viewed as an active tool for risk management that requires routine updates.
Effectiveness of the process difficult to prove
1.
enables better informed and defensible plans, budgets and schedules;
2.
increases the likelihood that the project will follow the plan;
3.
better type and structure to contracts;
4.
provides a rigorous assessment of contingencies;
5.
reduces the likelihood that economically flawed projects will be accepted;
Where uncertainty in intrinsic, it is impossible to demonstrate the effectiveness of risk analysis. If the risk failed to develop was the risk analysis deficient or was the risk management process effective? Risk analysis cannot predict where, when and if a particular risk will occur – it can only assess the likelihood and consequences; management must decide whether or not to accept the risk. There is often a general reticence to acknowledge low probability high consequence events and to act accordingly. When such an event occurs it is viewed as an indictment of risk analysis not of management practice, despite the fact that no risk reducing actions were taken.
6.
contributes to an informal database of project experience useful corporately;
The process can antagonise or threaten staff
7.
enables objective comparisons of alternatives; and
1.
formalises corporate experience and improves general communication;
Overselling risk analysis by management or the risk analyst leads to disillusionment. A lack of credibility in those on the risk team, especially the risk analyst, leads to scepticism of the benefits. There can be a lack of co-operation, especially from middle management who feel threatened by the process and intimidate their staff who are asked to participate.
2.
improves understanding between disciplines and team spirit;
CONCLUSIONS
3.
helps to distinguish between good luck and good management;
4.
develops the ability of staff to assess risks in everyday tasks;
Within the mining industry risk analysis is slowly gaining ground but still has many hurdles to jump. This paper opened by suggesting that:
5.
focuses management attention on the real issues;
6.
demonstrates a responsible approach to staff, community and the environment;
8.
identifies and allocates responsibility to the best risk owner. Soft benefits relate to improved communications, as follows:
7. 8.
allows justifiable economic risk taking form a position of understanding; and
confusion with the terminology;
• scepticism of the results because they are not interpreted before submission to management;
• a general lack of trust in the method because it is not understood;
• unwillingness to accept risk analysis as a management tool; and
encourages participation at all levels in the organisation.
Risk analysis is not the universal solution to all problems, and set against the benefits noted above are several limitations. The severity of the limitations can often be reduced, or even eliminated, by a well designed and properly managed risk analysis project. Limitations can be grouped under five headings.
Garbage in and Gospel out The results of a risk analysis are only as good as the quality of the information provided. Poor quality models or an inexperienced modeller will result in misleading results. The risk analysis is only as good as the analyst. All risk analyses require evaluation and interpretation; simply presenting the results to management is not good enough.
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• there is misunderstanding of the types of risk analysis and
• an erroneous belief, in some quarters, that risk analysis is being used correctly. This paper has attempted to correct some of these issues by describing the types of analysis typical in mining projects, using examples from other industries; and has reached a number of conclusions to be considered when applying risk analysis and management in mining projects: 1.
risk analysis is not a goal, it’s a process which can assist management if properly conducted, interpreted and reported;
2.
risk analysis has a place at all levels in a mining operation and its benefits are not confined to safety and environmental management;
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3.
risk analysis in mining is based on sparse data, requires expert participants, pools the collective opinions of all specialists, and does not replace the need for retained experts;
4.
quantified risk analysis is usually less useful that subjective analysis;
5.
risk definition and risk ownership are the keys to effective management of risk;
6.
risk analysis is a specialist task requiring experienced practitioners who should be ‘external’ to overcome project ‘group thinking’;
7.
a risk register is the management tool emanating from a risk analysis; and
8.
risk analyses need frequent updating to maintain their value.
Risk analysis practitioners and managers who have used the methods must do their utmost to ensure that this useful technique receives wider application across the industry and is not left as a tool for the safety department and the environmental section.
REFERENCES Davis, G A, 1995. (Mis)use of Monte Carlo simulations in NPV analysis, Mining Engineering, Jan:75-79. Eisenberg, N A, 1975. Vulnerability model - a simulation system for assessing damage resulting from marine spills, National Technical Information Service AD-A015-245.
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Environmental Management – Specification for guidance and use, 1996. BS EN ISO 14001, 4. European Environment Agency, 1998. Environmental risk assessment, Copenhagen, Environmental issues series No 4. Hambly, E C and Hambly, E A, 1994. Risk evaluation and realism, Proc Instn Civ Engrs Civ Engng, Vol 102:May. Health and Safety Executive, HMSO, 1992. The tolerability of risk from nuclear power stations. Health and Safety Executive, HMSO 1995. Designing for health and safety in construction – A guide for designers on the construction (Design and Management) regulations 1994. Kampmann, J, Summers, J W and Eskesen, S D, 1998. Risk Assessment Helps Select the Contractor for the Copenhagen Metro System, World Tunnel Congress ’98, 24th ITA Annual Meeting Sao Paulo, April. Laubscher, D H, 1994. Cave mining – the state of the art, J S Afr Inst Min Metall, October. Powell, D B and Beveridge, J, 1998. NATM in Context, Tunnels and Tunnelling International, January. Ravenscroft, P J, 1992. Risk analysis for mine scheduling by conditional simulation, Trans IMM, Vol 101, May. Risk + Program Management Solutions Inc, Santa Barbara, CA, Tel: +1 805 898 9571. Royal Society study group, 1992. Risk analysis, perception and management. Simon, P, Hillson, D and Newland, K (Eds), 1997. Project risk analysis and management, Assn Proj Mgmnt. The collapse of NATM tunnels at Heathrow Airport, 2000. (HSE Books) ISBN 0717617920, June. Vose, D, 1996. Quantitative risk analysis (J Wiley).
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Planning and Scheduling Simulation Modelling of Mining Systems
B E Hall
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Open Pit to Underground — Transition and Interaction
T R Stacey and P J Terbrugge
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Quantifying Geotechnical Risk in the Mine Planning Process
T P Horsley and T P Medhurst
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Dilution Control in Southern African Mines
R J Butcher
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Reflections of a Mine Scheduler
J Luxford
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Simulation Modelling of Mining Systems B E Hall1 ABSTRACT This paper discusses the requirements for a successful simulation modelling study, and the steps involved in the modelling process. Advantages and disadvantages of simulation are described, and a case study highlights how not to go about it. Two other case studies, for a decline truck haulage and a block caving production operation, are used to illustrate the types of variables whose impact may be investigated. Typical results are presented to indicate the sorts of presentation styles the author has found useful for conveying study results to project sponsors. Standard spreadsheet charting techniques can be very useful in this respect, but the possibilities are limited only by the imagination of the study team.
INTRODUCTION Simulations of mining operations have been carried out for many years. In many cases, end-users of simulation projects have been disappointed with the results obtained. Often this is because the strengths and weaknesses of the tool have not been fully understood, and the aims of the simulation project were not clearly defined. This paper discusses the requirements for a successful modelling study of any kind, but with particular reference to simulations. It considers when simulation is and is not the appropriate tool to use, looking at some of its advantages and disadvantages. Examples of simulations of (a) block caving production operations, from drawpoints to final delivery of ore, and (b) operation of a truck haulage fleet in an underground mine, from the source of the rock to its final destination, are described. These illustrate the types of parameters which might be investigated, and how results can be presented to facilitate decision making based on sound engineering processes.
AN OVERVIEW OF ‘MODELLING’ There are many types of models that can be built, such as physical models, computer spreadsheet models, and animated discrete event Monte Carlo simulation models. The last of these is the main topic of this paper. This description indicates that this type of simulation takes account of both random variation and changes in the system being modelled over time, and can be animated to provide a visual appreciation of the behaviour of the system. All models have certain common features. They are approximations of the real thing, and are built with a particular purpose in mind. They should produce a good representation of the particular aspects of reality relevant to their purpose, but are unlikely to be good representations of other aspects of the systems that have not been modelled. The validity of a model is determined by the accuracy of its outputs, not by the reality of its inner workings. A very accurate representation of reality is a replica, not a model, and a lot more time and effort than was necessary will have been expended to obtain useful results. For example, in mining, linear elastic models are often used to analyse rock stresses. Although in many cases it is known that the ground does not in reality behave in a 1.
MAusIMM, Principal Mining Engineer, Australian Mining Consultants Pty Ltd, Level 19, 114 William Street, Melbourne Vic 3000.
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linear elastic way, the results are close enough not to alter the decisions that would be made. Additional accuracy will not change the outcome, so the added complexity of non-linear calculations is not warranted. The crucial question to ask about a model is not ‘How accurately do its calculations mimic the real interactions?’ but is rather ‘If I feed in realistic input data, does the model produce realistic outputs?’ There is a problem here in identifying the boundaries and the level of detail of the modelling project. If these are too wide or too great, unnecessary work is done. But if they are not broad or detailed enough, there is the risk of being lulled into a false sense of security: work has been done and a result obtained, but a feature not modelled may yet have an adverse impact. There is often no easy way of identifying where the bounds should be. For the sake of conservatism, it is this author’s opinion that it is better, and potentially less expensive in the long run, to err on the side of doing too much than not enough (though one must be careful to avoid ‘paralysis by analysis’).
Steps in the modelling process All modelling projects have a number of process steps (Pegden, Shannon and Sadowski, 1995). These apply to all types of modelling, and are very relevant to discrete event simulations. The following subsections briefly describe these steps, and the reasons for them.
Identification of the problem The model will generally be purpose built for the problem. The problem to be solved must therefore be specified. This may be vague at first, but become clearer as initial project scoping proceeds.
Statement of project objectives No model can do everything, or it will become too large and cumbersome. The scope must therefore be defined to ensure that all key issues are covered, and non-essential complications do not distract from the key purposes. Together with the identification of the problem, this is one of the most critical parts of the project. It may require revisiting as the project progresses, to maintain focus, and/or to formally revise the objectives.
Collection and preparation of data Input data, both in the description of the system and for the numbers to be used for various parameters, must be accurate. Real data should be used in the model as soon as possible for reality checking as the model is being built. Non-availability of data may compromise the validity of results. The old adage ‘Garbage In - Garbage Out’ is very relevant. If data is not available, assumptions need to be clearly stated, and sensitivity analyses may be necessary to ascertain how critical that data might be, to determine whether more detailed investigation is necessary.
Formulation and construction of the model Generally it is preferable to construct a model as a group of structured modules. This aids subsequent debugging and modification.
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It is important that the potential ultimate model requirements are known at the beginning of the project, even if all features are not constructed initially. That way, allowance can be made in the overall design for later incorporation of more detailed features. If this is not done, at best there may be a loss of time and/or efficiency as new items are grafted into something which was not designed to accommodate them. At worst, it may be necessary to start again from scratch, as the methodology used may not be amenable to modelling the new features.
Verification and validation of the model There are two distinct tasks to be done at this step. Verification ensures that the model works as intended, for the full range of input conditions. Validation tests that the model results are realistic for known real world cases where possible. It is only after this stage that it is possible to use the model with confidence to make predictions about cases which do not yet exist in reality. Animation can be a very useful debugging tool at this stage of the process with simulations.
Modification and/or refinement of the model Feedback loops at all stages of the modelling process help to ensure that the model is constructed as efficiently as possible. Early results may cause a rethink of model logic, or even of the objectives of the study.
Using the model After gaining confidence from the verification/validation phase, the model can then be used to make predictions. However, it may happen that, where a model is being developed to identify problems with existing systems, the act of fully investigating and describing the current system to enable it to be accurately modelled leads to a realisation of where the problem lies. When a modeller who is not fully conversant with the system starts asking probing questions to gain understanding, those who know the system may investigate certain aspects more thoroughly, and discover something previously unknown. Animations of simulations can also be very useful for understanding the behaviour of a system. However, with an animated simulation, as in real life, it is difficult or impossible to guarantee that a problematical situation will arise while the animation is being viewed, and that, if it does, it will be recognised as such.
Applying the results It may appear self-evident, but it needs to be recognised that, if there is no intention to use the results of a modelling project, it is a waste of time and money doing the job in the first place. The project sponsor needs to be in a position to influence implementation of the study recommendations.
ADVANTAGES AND DISADVANTAGES OF SIMULATION Simulation has many benefits. It can provide realistic estimates of the behaviour of a system. It is ideal for evaluating the effects of changes in complex dynamic and interrelated systems, such as changes to the number or types of machines, processing rates, machine availability and breakdown parameters, the process flow, allocation of operators to various tasks, product mixes or blends, material handling methods and stockpile sizes. Simulation can provide results when simple analytical solutions cannot be calculated, components in the overall system interact in a complex way, and what is ‘lost on the swings’ may
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not necessarily be ‘gained on the roundabouts’. No matter how complex the system, if it can be described, it can be modelled to the requisite degree of accuracy. By utilising the stochastic random behaviour features of the software, both average performance and likely variability can be estimated. Experiments can also be conducted without disrupting the real system, or before the real system exists. Animations make visualisation of the system easier. They can often be crucial in ‘selling’ the validity of the study to influential sceptics. As noted above, they can also be very useful at various stages of the modelling process. On the downside, brief viewing of an animation has the potential to unduly influence an analyst about the relative importance of an observed feature. To obtain sufficient results to support rigorous engineering decisions, it is generally necessary to do a large number of runs without the animation. Output that can be analysed dispassionately must be generated, tabulated, plotted and compared. Despite its power there are a number of things that simulation cannot do. For example, it cannot automatically optimise the system. However, it will give answers to various ‘what if?’ questions which can point a team towards the optimum answer. It cannot give accurate results from inaccurate data nor describe system characteristics which have not been explicitly modelled. Nor, importantly for clients’ expectations, can it provide fast easy answers to complex questions. Simulation may, however, be the only technique which is capable of producing the desired result at all.
PITFALLS TO AVOID WITH SIMULATION Simulation may not be the appropriate tool When a problem is identified, there are frequently many ways it can be solved. At times, simulation may be used for problems where it is not the most appropriate tool. This can result in an end user unfamiliar with the strengths and weaknesses of simulation becoming dissatisfied with it, and subsequently rejecting its use for a problem where it is the ideal tool. At best, the problem solving process may be slower than necessary if simulation is applied inappropriately. Two areas which this author believes are not the province of simulation are the development and analysis of a long-term mining plan, and detailed scheduling of mining activities over time. The author believes that these two classes of problems are perhaps best handled by ‘spreadsheet’ and ‘project management’ software respectively. Stochastic add-ins for ‘off-the-shelf’ software permit evaluation on the impact of random events in these types of analyses. Results of discrete event simulations of the state of the mine at various times and with various constraints applied may give invaluable input data with respect to production capabilities and bottlenecks to use in the other analyses. But the intricacies of determining optimum mine scheduling rules, for instance, are at the cutting edge of research. They are not easily incorporated into ‘off-the-shelf’ simulation packages and models developed using them.
Sufficient analysis may not be done For valid conclusions to be drawn, a statistically valid sample must be obtained. It is beyond the scope of this paper to discuss this in any detail. However, to ensure that low probability events or combinations of events are able to occur, and that their impact on overall predictions is neither too great nor too small, the simulation modeller must give careful consideration to both the length of a simulation run and the number of replications done.
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It must be recognised that, because of the stochastic nature of events both in real life and in a simulation model, results of observations may not plot on a straight line or smooth curve such as might be predicted. With a relatively simple simulation, it may be that all random variations are averaged out, and ideal results are obtained. But with a more complex system, it may be impractical to achieve this. The more complex the model, the more dangerous it may be to draw conclusions from a short run, or from analysis of a single run. Longer runs, and for a number of scenarios, may be required to interpolate trends. A regression line may give a better estimate of the value of a dependent variable than the actual observed value with a given value of an independent variable. Similarly, with simulations, the trends, rather than the results actually obtained for that case, may give the best estimate of the most likely results for a given set of circumstances.
Inadequate project control may compromise the study A small case study is used to illustrate this potential pitfall. Modifications to the treatment plant at a certain mine were reducing significantly the amount of ore storage available. At the same time, the mine was considering increasing its output. The questions posed by senior management were apparently simple: Can the mine produce at the proposed target rate, and is the reduction in surge capacity a constraint? A study team was established, comprising an experienced modeller and a relatively junior mining engineer. The modeller was skilled in the use of the simulation software package, but was not familiar with the mine. The mining engineer was too junior to seriously challenge what the modeller was building into the model. He was also unaware of management’s interest in changing working rosters for both operating and maintenance staff in both the mine and the treatment plant. Management had not ensured that the need to evaluate such changes, should they be necessary for the achievement of the goals, had been communicated to the study team. After several weeks of gathering data, building a model, and generating results - all without further management involvement - a draft report was prepared, concluding that the target production rate was achieveable, and the surface surge capacity was not a constraint. However, some comments in the report rang alarm bells, and the author was called in to audit the study. Investigations revealed that:
• The study did not make use of random effects. All calculations were fully deterministic.
• No allowance was made for the fact that the treatment plant worked 24 hours a day, seven days a week, while the mine had a weekend break in excess of 24 hours.
• Though the report was being published with all the authority of: ‘A simulation study has determined that …’, the real calculations in the model affecting the results could be fully and accurately summarised as follows: 1.
Shaft availability x hours per month x hoisting rate = tonnes hoisted per month, > target production per month ∴ target is achieveable
2.
(Hoisting rate - treatment rate) x hours per month = increase in stock in one month < surge bin capacity ∴ surge bin is not a constraint.
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A complex model, built using a powerful tool, had done nothing more than ‘back of envelope’ calculations. The first calculation was trivial. Management knew that the shaft itself was not the constraint, but was aware that patterns of ore production in the mine (affected by availability of ore and equipment availabilities) and treatment in the mill (affected by breakdowns and maintenance rosters) were the crucial issues, yet these were not investigated at all in the study. The second calculation was also trivial, but was compounded by a serious logical flaw. The net rate of increase in stocks was positive, and there would eventually come a time (after some two months with the data used) when the surge bin would be filled, and the hoisting rate would have to drop back to the treatment rate. The analysis had an implicit assumption that somehow the surge bin emptied itself at the end of each month. Clearly there are a number of problems highlighted by this example. If they had not become apparent in time, serious wrong decisions could have been made. The study was subsequently re-done, requiring a complete re-build of the model to enable it to answer management’s real questions. Certain workplace changes were identified as necessary, and ‘sold’ to the workforce in part by using the simulation results. The key lessons for a successful simulation study are:
• the
sponsoring managers must communication with the study team;
remain
in
regular
• management must ensure that the problem to be solved and the objectives of the study are clearly understood by all concerned;
• management must ensure that the model is capable of answering all the questions they may wish to ask of it;
• the team working on the study must consist of at least: - a person familiar with the software, and able to -
appreciate how real world situations may be translated into model constructs; and a person familiar with the system being modelled, to ensure that all relevant real world behaviours are accounted for by the model.
These two may at times be one and the same person. If not they must both have such personal qualities and comparable levels of seniority as are needed to ensure that each is able to examine, understand and challenge the work of the other to ensure the final result is reliable.
TYPICAL SIMULATION RESULTS There are many types of systems which can be investigated, as indicated in many papers on simulation over the years (Sturgul, 1997). In many of these, a substantial proportion of the paper is devoted to reasons for the selection of the software, and/or a detailed description of the mining system being studied. The types of results obtained and presented often receive little attention. In this paper, the author’s aim is to indicate to potential users of simulation the types of things which might be investigated, and a few ways in which the results might be presented. This is done by way of two case studies. In each case, a description of the system investigated is given in broad terms only, to provide basic understanding. The types of variables included in the model, and thus able to be changed for evaluation of their impact on the system, are described. Some typical results from each study are presented, showing a number of presentation formats that this author finds useful. It should be noted that, if something can be measured in real life, then it can be measured and recorded in a simulation model. The only limitations are disk space on the computers, the imagination of the modelling team, and the time available to run the models and draw useful conclusions.
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CASE STUDY 1 – TRUCK HAULAGE IN A DECLINE Description of system A study was undertaken on behalf of the Kanowna Belle Gold Mine in Western Australia, 100 per cent owned by Delta Gold Limited. The main objective of the project was to identify the truck fleet requirements for achieving production targets over a number of years as the average depth of production increased. The study investigated the truck haulage system in some detail, with simulation modelling taking into account such factors as planned maintenance and breakdowns, shift working rosters (times and durations of shift changes, meal breaks and vehicle servicing), queuing and congestion in the decline passing bays and at loaders, and the total duty requirements of the truck fleet at various stages of the mine life. The impact of other vehicles in the decline was also accounted for. The model was not developed to allow significant decisions to be made about loaders and loading operations, but simple modelling of loaders was used to define the number and locations of truck loading points, taking into account planned and breakdown maintenance of loaders and their operational availability after allowing for meal breaks and the like. The main features of the trucking system were as follows. The underground haulage occurred in a fairly typical decline, commencing near the bottom of an existing open pit. Passing and overtaking were possible only at defined passing bays. Mining levels branched off the decline at various defined locations. For each of the four major mining blocks defined at the mine, representative levels were defined for production loading, development loading and fill tipping. All material hauled to surface was tipped either directly into the surface crusher, onto the main run-of-mine stockpile near the crusher, onto a small ore stockpile in the open pit, near the portal, or onto the waste dump. Provision was also made for underground tipping at a possible shaft location.
Simulations evaluated truck requirements to meet scheduled tonnages of production and development ore and waste, and two types of fill, for three specified years of the mine’s schedule. Productivities were also assessed for the three major activities of production, development, and backfill separately, for each mining block. Two truck types, ‘tippers’ and ‘ejectors’, were modelled. All available ejector trucks were allocated to filling if there was fill to be hauled. If not, ejector trucks were able to be allocated to production or development tasks, as were all tipper trucks. A simple truck despatching algorithm was implemented to ensure that loaders and trucks were allocated to production and development rock sources so that the simulated production remained in line with the schedules. To maximise truck fleet utilisation, queuing at one loader while another was idle was also avoided.
Typical investigations and results A number of production and equipment parameters were measured in the model and output in a simple spreadsheet file, which was subsequently processed to generate a variety of results graphs. For each case run, the following types of plots were typically generated, along with the detailed data tables which they summarised.
• Daily ore production from each stoping and development source (Figure 1).
• Cumulative frequency distribution of daily ore tonnages (Figure 2), which summarises Figure 1.
• Daily split of operating, maintenance and idle ‘activities’ for each type of truck and loader (Figure 3). These could also be summarised into a ‘pie’ chart if desired.
• Histograms of the numbers of trucks in the decline or on levels at any time (Figure 4). Figures 1 - 3 show the quite significant fluctuations in daily performance which can be expected in this particular system. These result purely from the system characteristics, rather than ‘good’ or ‘ bad’ management practices. Figure 4 was useful for identifying ventilation requirements.
Fig 1 - Daily ore tonnes trucked - year 3 targets.
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FIG 2 - Cumulative frequency distribution of daily ore tonnes trucked.
FIG 3 - Daily fleet availability and utilisation - year 3 targets.
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FIG 4 - Proportion of the time various numbers of trucks are in the decline.
For each mining block, cases were run with varying numbers of trucks, for production, development, and filling separately. For each product hauled, a number of plots were generated. For each mining block, typical plots were:
• Productivity (tonnes per hour and hours per tonne) as a function of mining depth, at target production rates (Figure 8). The truck operating hours per tonne curves are essentially linear with depth, indicating that the interference effects with larger fleets are also linear with depth.
• Tonnes hauled vs number of trucks in fleet, showing the daily averages and a range of percentiles above and below the average (Figure 5). The averages were generally relatively linear until the output approached the amount of rock actually generated by the model, when a levelling off occurred. (It was specified for the study that ore should always be available for trucking. The model therefore generated rock to be trucked at some 200 per cent of the target rates.)
• Truck fleet availability and utilisation components vs number of trucks (Figure 6). The non-working times increase slightly with increasing fleet size while output is increasing, but when the maximum rock output is approached, these idle times increase more rapidly. Figure 5, like Figure 1, indicates the level of normal daily production variability to be expected. It can also indicate the size of fleet required to be able to produce at least the daily target for a specified proportion of the time. The results on which Figure 6 is based can indicate realistic availability and utilisation targets for maintenance and operating sections, taking account of both the size of the fleet and the duty required of it. Comparisons for all mining blocks were generated, such as:
• Production and productivity curves for each block, for various fleet sizes (Figure 7). Production curves for Blocks A and B, relatively close to surface, show significant flattening of the curve as the rate of ore generation used is approached. Deeper Blocks C and D exhibit a more gentle curvilinear behaviour as the increasing truck fleet results in some increasing interference.
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Figure 9 shows the effect of changing passing bay spacing in a truck decline. The Kanowna Belle model had the facility to evaluate this behaviour, but this was not done for that study. This figure was generated for another similar study. It can be seen that, as the spacing of passing pays decreases, the truck productivity approaches that of a one-way loop haulage, at significantly less cost. For simulations of the total materials movements, two distinct modes of filling operation were identified. When cemented fill was to be run, the operators would target nominally full-time operation of the cement slurry plant and fill loading system with no work breaks. At other times, dry rock fill would be run at its average required daily rate with ‘normal’ work rosters. This indicated two possible operating strategies, which were both assessed: 1.
operate a truck fleet which is large enough to produce the average daily ore target, even when cemented fill is being hauled at the maximum rate, aiming to do scheduled maintenance when cemented fill is not required and the total trucking demand is lower; and
2.
operate a truck fleet so that the overall daily average target is met by producing more than the target when cemented fill is not required, and less when cemented fill is required.
Study conclusions The main conclusions of the study were as follows:
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FIG 5 - Daily ore tonnes trucked - Block A - production ore only.
FIG 6 - Truck fleet availability and utilisation - Block A - production ore only.
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FIG 7 - Daily ore tonnes trucked - production ore only.
FIG 8 - Truck productivity vs haulage depth - production ore only.
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FIG 9 - Trucking capacity as a function of fleet size and passing bay spacing.
1.
If variations in ore production were acceptable depending on the filling duty at any time, a fleet of a certain number of trucks was required.
and downstream conveying of the DOZ ore. The model was built as a standalone model, but could also be incorporated into an overall site ore flow model developed independently for PTFI.
2.
If it was required that daily production should average the target even while filling with cemented fill, up to two additional trucks would be required in later years.
Description of system
3.
Significant daily fluctuations in output were likely in any case.
• An extraction level with: - up to 30 extraction
4.
‘Over-trucking’ and consequent significant queuing idle time was necessary to meet cemented fill targets, which effectively required the fill loading stations to work non-stop.
5.
There was a reducing contribution made by additional trucks as the rock generation limit was approached. Therefore, if the development, drilling and blasting capability were not significantly in excess of the daily targets, additional trucks would be required to provide additional instantaneous trucking capacity, to ensure that average targets were met despite there being no ore to haul at times. (This effect was not quantified for this particular study.)
CASE STUDY 2 – BLOCK CAVING PRODUCTION OPERATIONS PT Freeport Indonesia (PTFI) commissioned a simulation of the underground block caving production system for the planned ‘DOZ’ operation in the Ertsberg Mine in Irian Jaya, Indonesia. The simulation modelled the mucking and secondary blasting activities on the block caving extraction level, trucking operations on the truck haulage level, and simplified crushing
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The mining system modelled consisted of:
-
-
panel cross-cuts, aligned approximately north east - south west; an orepass dump near the centre of each panel, effectively breaking each panel into northern and southern sections, and thereby allowing two loaders to operate in each panel simultaneously; and three additional dumps off the northern perimeter drive, servicing the northern sections of a number of the central panels.
• A truck haulage level with: - a one-way loop haulage, outbound empty in the north, -
and inbound full in the south; truck loading chutes in cross-cuts connecting the inbound and outbound haulage drives; an additional haulage loop to the north to access the three northern ore passes; and a central dumping area, with east and west accesses from the haulage, and three truck dumping locations into the dump pocket above the crusher.
• Crushing and conveying facilities consisting of dump pocket, crusher, surge bin, and a generic ‘conveyer’ drawing from the surge bin, representing the entire ore handling system downstream of the surge bin.
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The main features that were taken into account in the model were:
• prioritisation of ore passes, taking account of ore pass
On the extraction level:
• truck loading chute hang-ups and failures.
• a number of different loader and hang-up drill types; • loader movements between drawpoints and dumps, between cross-cuts, and to and from the workshop;
• delays due to planned work breaks, breakdowns and scheduled maintenance, and full and blocked dumps;
• prioritisation of drawpoints and cross-cuts, taking account of drawpoint daily draw limits, maximum time between successive drawings, and maintenance of even draw;
• three types of drawpoint hang-ups; • temporary sterilisation of drawpoints caused by hang-ups in adjacent drawpoints;
• drilling and blasting of hang-ups; • different hang-up, sterilisation and blasting parameters for different muck types;
• blasting outside scheduled times as an option; and • drawpoint failures and repairs. On the truck haulage level: • a number of different truck types; • truck movements between chutes and dumps, including the effect of broken down trucks in the haulage;
• delays due to planned work breaks, breakdowns and scheduled maintenance, full/blocked dumps, empty/inaccessible ore passes, and queuing at chutes and dumps;
contents, and numbers of trucks already allocated to passes; and
Crushing and conveying system: • planned maintenance shutdowns and unplanned breakdowns; • existence of surge capacity at certain locations, and availability of ore to draw or space to dump in surge bins; and
• maximum processing rates for each item of plant. Typical investigations and results The model was delivered to PTFI, who subsequently performed all their own analyses. To facilitate this, all inputs to the model were via text files. These were generated by Microsoft Excel spreadsheets for ease of use by PTFI personnel. The model also wrote all its output to a spreadsheet file. A template Excel spreadsheet was provided for processing the results from each model run into a useable format. Sample results have been generated by the author, independently of PTFI’s studies. Figures 10 and 11 show two different views of production with varying numbers of loaders and trucks. The figures show clearly the regions where the operating constraint is either the number of loaders, the number of trucks, or neither of these, but rather the existing state of the mine. The flat but irregular horizontal region of the surface in Figure 10 is a good example of the situation, described above, where even long simulation runs may not result in a smoothing of all the random effects. This is the region where the ‘state of the mine’ rather than the size of the equipment fleets, is the
Fig 10 - Daily production as a function of loader and truck fleet.
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FIG 11 - Contours of daily production as a function of loader and truck fleet.
constraining factor. Investigation of these irregularities indicated that, although the long run average number of unserviceable drawpoints was consistent in all cases, the pattern of location of these over time affected the production rate to varying degrees. Runs of sufficient length to account for all possible variations in drawpoint failure patterns would be impractically long. It is therefore necessary to estimate the expected maximum possible production from the mine in a particular state as the average of those cases which are judged to be limited by the state of the mine. Figures 10 and 11 were generated with a secondary blasting drill fleet too large to be a constraint. Having identified a limited set of suitable truck and loader fleet sizes, similar plots could be generated for various numbers of drills. This would then assist in deriving the balance of trucks, loaders and drills best able to meet production targets. Also, using fleet sizes and production rates as shown in Figure 10, together with productivities and utilisation statistics from the model (similar to those in Figure 6) and applying various operating and capital costs and revenue assumptions, the optimum economic operating point could be determined. The PTFI data specified different characteristics for different ore types. In particular, various categories of wet muck would hang up in drawpoints less frequently than similar dry ore, but would require longer periods of delays to allow stabilisation of the hang-up before secondary blasting. Figure 12 shows the decline in maximum production rate as the proportion of drawpoints with wet muck increases. The values shown for each data point are the estimated average for the ‘irregular mine-state limited’ cases in each scenario, as described above, with the ranges of the irregularities indicated by ‘error bars’. It can be seen that there is initially a significant effect as the number of wet drawpoints increases, but this effect reduces as the proportion of wet drawpoints increases further.
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A number of other parameters were reported by the model. Loader and truck ‘activities’ were reported as 8 and 13 categories respectively. Similarly, 20 different drawpoint states were identified and reported, and summarised into five main categories. Records were also kept and plotted for truck chute status and drawpoint and roadway repairs. These were all presented in plots similar to Figure 3 for showing the various effects over time, and as pie charts to summarise the results for each run. Stacked bar charts, similar to Figure 6, were also used to show trends across a number of runs. As noted above, the information able to be recorded and reported is virtually unlimited. Sample results were also generated to test the ‘unscheduled blasting’ feature of the model. For the situation assumed, it could be seen (Figure 13) that the number of unscheduled blasts is small until there are excess loaders available. However despite this increase in the amount of unscheduled blasting, the overall benefits are negligible. This would suggest that factors such as maintaining an even draw eventually limit the overall production rate, despite the apparent ability to blast more frequently and muck higher tonnages in the short-term. The model constructed permitted the effect of a large number of parameters to be assessed. The author is aware that PTFI has continued to use the model to investigate a large number of possible mining scenarios.
CONCLUSION Simulation is a powerful tool for the mining engineer. When used in appropriate applications it is able to provide insights into system behaviour in a way that few, if any, other techniques can. However, if improperly used, serious mistakes can arise. This is perhaps exacerbated by the fact that results are not always intuitive, and can rarely be checked by random audits of reported results only, or by a person not familiar with the intricacies of the particular software package used.
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B E HALL
FIG 12 - Block caving production potential as affected by drawpoints with wet muck.
FIG 13 - Impact of unscheduled blasting on production.
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SIMULATION MODELLING OF MINING SYSTEMS
With a complex system, a large number of runs may be necessary to adequately investigate the response of the system to variations in a number of parameters, both individually and simultaneously. Necessary skills in the simulation team are the abilities to assess what cases actually need to be evaluated, and to present the results in ways which can be readily understood. Spreadsheet (and other) plotting and graphing software can assist with this. The information reported by the model as it runs must meet the needs of the project sponsors, but is generally only limited by the imaginations of the modelling team.
The managements of Delta Gold Limited and PT Freeport Indonesia are thanked for their permission to describe the models and results pertaining to their respective operations.
REFERENCES Pegden, C D, Shannon, R E and Sadowski, R P, 1995. Introduction to simulation using SIMAN, 2nd edn, pp 8-24 (McGraw Hill: New York). Sturgul, J R, 1997. Annotated bibliography of mine system simulation, in Mine Simulation - Proceedings of the First International Symposium on Mine Simulation. (Balkema : Rotterdam).
ACKNOWLEDGEMENTS The author wishes to thank the management of Australian Mining Consultants Pty Ltd for permission to prepare and present this paper, and also the company’s secretarial staff for assistance in its preparation.
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Open Pit to Underground — Transition and Interaction T R Stacey1 and P J Terbrugge2 ABSTRACT Many ore deposits have considerable vertical extent which is not always known at the time of commencement of mining. In many of these cases, initial mining is by opencast methods, and as ore reserves are proved to greater depths, the pits are often planned to go deeper than had been originally envisaged. In such cases, surface plants and critical underground facilities such as conveyer tunnels, access ramps, ore passes, hoisting shafts and ventilation shafts are often located much closer to the pit rim and the orebody than is desirable for the deeper mining situation. This usually raises the question of the stability of such underground excavations, which may be critical for the viable longer-term mass underground operation of the mine. In this paper the key considerations in the evaluation of the stability of the overall mining environment for ‘deepened mines’ are defined. Aspects which are dealt with are: • open pit stability, • shaft stability, • dilution, • mudrushes and airblasts, and • selection of mining method. The necessary stability evaluations are put into the appropriate context by considering two case histories. Critical parameters and considerations which need to be addressed during the evaluation of excavation interaction are also discussed.
INTRODUCTION Many mining operations commence using open pit methods. At the planning stage, mining to a defined depth is considered, and the location of facilities such as crushers, reduction plant, waste rock dumps and tailings dams are determined accordingly. With the progress of mining, payable ore reserves may be proved to greater depths and a deeper pit or underground mining may be required. In such cases the surface installations may be closer to the mining operations than may be desired. In some cases these facilities may be within the area of deterioration due to the mining - the risk to their safety may be such that some have to be moved. In numerous mining operations, underground mining often commences after the completion of the open pit operation. In some cases a crown pillar may be left between pit bottom and underground workings, and in others the underground operations may proceed from the pit bottom or be exposed to the pit bottom. In caving operations, caving will take place through to the pit, and the cave muckpile will separate the underground workings from the open pit above. In the last two cases the underground workings are exposed to dilution by waste from the pit and mud and water inflow from the pit. In the first case, the crown pillar should protect the underground workings from these undesirable occurrences. The risk associated with the crown pillar is its instability, which could lead to an air blast, and allow water inflow and dilution. Airblasts are a potential risk in most underground massive mining operations. Mudrushes are a potential risk when there are clay forming minerals and water present, and entry points into the underground workings, for example, drawpoints exposed in the old pit, and caving 1.
Director, SRK Consulting, P O Box 55291, Northlands 2116, South Africa. E-mail: [email protected]
2.
Director, SRK Consulting, P O Box 55291, Northlands 2116, South Africa.
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drawpoints. Airblasts can also be associated with the occurrence of mudrushes. A reverse of the above can also occur, in which underground operations have been carried out first and opencast activities make take place many years later. A particular example of this is in coal mining. A seam may be mined by bord and pillar methods. Subsequently, other seams may become economically mineable, often by opencast methods, owing to the improvement in opencast mining equipment and commensurate improvements in mining efficiency and lowering of mining costs. In some mining operations, underground and open pit mining take place concurrently. In such cases there may be on-going interaction between the two - underground operations may influence the stability of pit slopes adversely, and the open pit may increase the risk of water, gas and mud inflows into underground workings. More commonly, concurrent open pit and underground activities take place during the transition phase from full open pit to full underground mining. This transition requires careful planning and it is considered appropriate to summarise some of the main issues which should be taken into account.
PLANNING CONSIDERATIONS — OPEN PIT TO UNDERGROUND The efficient transition from open pit mining to an underground operation requires extensive planning. For a large mining operation, this planning period could be as long as 20 years. The main factors and activities which should be taken into account in the planning are:
• Definition of the orebody: this will define the orebody shape, the orebody plan outline (footprint), the depth extent and shape variations, the three-dimensional geometry, and the maximum potential mining depth.
• Rock mass characterisation: the purpose of this is to obtain details on the major geological structures, the jointing in the rock mass, the rock types, the rock mass quality, and the groundwater conditions. These all affect the stability of any proposed underground excavations.
• Definition of the boundary conditions: the boundary conditions that must be taken account of include the in situ stresses, the topography, the three-dimensional geometry, any superloading, which could be in a positive or negative sense, the groundwater, and blasting and seismicity.
• Investigation of suitable mining methods and associated factors, for example, caving or other mining methods, combined open pit and underground mining, in which formation of wedge zones could occur, lift height, and long-term depth.
• Mining strategies such as: extensions of open pit operations beyond the planned time and depth. In the intermediate stage, should open benching take place? Is caving the best mining method in the long-term? What should be planned for remnant draw recovery?
• Underground infrastructure: decisions are required on the locations and sizes of shafts, orepasses, crushers, silos, chambers, ramps and declines, pump chambers, conveyor excavations and so on.
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• Surface infrastructure must be planned so that it is secure
IMPORTANT ASPECTS AFFECTING INTERACTION BETWEEN SURFACE AND UNDERGROUND OPERATIONS
from the effects of underground mining, and that underground openings are secure from any surface effects (such as floods, slumps from open pit slopes, and tailings dam failures, which could lead to mudrushes). Items of infrastructure that need to be considered are dams, tailings dams, rock dumps, plant, shafts, roads, rail lines, pipe lines.
Although, as indicated above, there are many factors which will affect the interaction between open pit and underground facilities, several of them, which are considered to be most significant, are dealt with in the following sections.
• On-going open pit factors of influence: stability of slopes, the effect of underground mining on stability, potential dilution of underground reserves from the open pit, and the potential for mudrushes and airblasts.
Open pit slope stability The implication of planning an open pit operation for optimum efficiency is that pit slope architecture will have been customised to the particular geotechnical environment present. The term geotechnical implies all of the geological, structural and hydrogeological in situ conditions which combine to determine the ultimate rock mass characteristics. The term architecture refers to the spatial relationships between, and face inclinations and geometrical dimensions of, each of the individual elements of a particular sector of the mine shell. These individual elements are as follows:
• Underground mining and layout considerations, such as fragmentation, layout geometry, sizes of drawpoints, orepasses, drifts, spacing of drawpoints, undercut location and geometry, undercut dimension and shape, support requirements.
• Underground and surface effects of the underground mining: cave angles and cave crack locations, which will affect the location of surface installations, underground location of cave cracks, which will affect the location of shafts, and the effect on the pit.
• individual bench face - height and face angle; • bench stack or inter-ramp slope - height versus angle
• Surface water and groundwater effects: probable maximum flood and water volume, groundwater inflow, potential rates of inflow, sump and pumping requirements, risk of mudrushes, and environmental effects.
relationship;
• spill berm width; • ramp or geotechnical berm width; and • limiting overall slope by design sector.
• Risks: a detailed risk assessment should be carried out to identify all relevant factors, and their degree of influence, which could impact on the profitability and successful implementation and operation of the underground mine.
In determining the optimum design for each sector of the mine shell, geotechnical units will have been evaluated separately before reconstructing them in their correct spatial relationship with each other. This final ‘sandwich’ of individual materials will have been analysed to determine the performance with regard to magnitude, rate and direction of potential slope displacement during the mining of the deposit. Adjustments will have been made to various elements of the slope architecture in order for the resulting slope displacements to be within acceptable levels. These levels will be in accordance with the required safety
• Project timing: as indicated above, a period of up to 20 years could be necessary from initial conceptual planning to full production. An indicative guideline-planning schedule is shown in Figure 1. The above points can be regarded as a checklist. Not all aspects will be relevant in all cases, in particular for smaller operations.
Activity
Time (years) 1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
Exploration Conceptual Design Stage 1 : Prefeasibility Board Review Stage 2 : Feasibility Board Review Stage 2A Stage 2B Feasibility Study Stage 2C Stage 3 : Final Design Stage 4 : Implementation
FIG 1 - Open pit to underground planning schedule.
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factors or probabilities of failure which have been established for the operation. Part of the design therefore involves the consideration of the frequency and magnitude of likely bench or slope failures, and whether they are occurring in critical sectors of the mine shell, such as in the vicinity of an access ramp or above a zone of high grade ore. These types of potential failure can be accommodated provided that monitoring of the actual conditions is carried out on a regular basis and that suitable, cost effective remedial measures can be implemented without delay. Of importance is a comparison of the predicted conditions with the actual performance of the resulting mine shell. It is important to note that the design of pit slopes also involves at least the following economic issues:
• selectivity of mining, ie high-, medium- and low-grade ore may be mined at differing bench heights due to the distribution of the ore grades; and
• selection of mining equipment to suit the scale of mining, ie daily production requirements may demand 15 to 30 m bench heights in order to produce the tonnage. Typically, bench geometries will have been designed to a probability of failure of 40 per cent (corresponding with a factor of safety of about 1.05 to 1.10) and the corresponding expected failure volume from the bench scale used to determine the spill berm width. When these benches and spill berms are stacked, the bench stack geometries are defined. Probabilities of failure for bench stacks will range from less than five per cent below major access ramps or areas containing mining infrastructure, to less than ten per cent above major access ramps, and then to between ten per cent and 30 per cent on slopes with no access ramps or infrastucture. These considerations make up the design of the overall slope geometries containing all of the required architectural elements - ramps, benches, special geotechnical berms, etc. The implication is, therefore, that in an optimum operation, pit slopes will have been designed to be very close to the limit of their stability. Consequently, there will be little scope for steepening of slopes to avoid impact on surface facilities if the effects have not been taken into account in the original planning and design.
Shaft stability In cases in which both open pit and underground mining is taking place simultaneously, and in which a shaft has been sunk in preparation for underground mining, but open pit mining has been carried on to greater depths than previously planned, shaft stability may be affected by deformations resulting from relaxation of, or increases in, stresses due to the deepening of the pit and the consequent encroachment of the pit rim. In the majority of mining situations the problem of interaction between stoping and underground service excavations (and between an open pit and underground operations) is a fully threedimensional problem. Usually the in situ stress field, the orebody, the open pit and stope geometries, and the layout of service excavations are all three-dimensional. In rare cases, for example with some tabular deposits, it may be possible to consider the operation as two-dimensional. Apart from the application of experience and the consideration of precedent practice, the only way in which the possibilities of occurrence of the above instabilities can be evaluated realistically for planning purposes is by using numerical stress or deformation analyses. Good application of such analyses depends on thorough knowledge of the detailed geology and geological structure of the rock mass, as well as of the deformation and failure behaviour of the rock mass. It can confidently be stated that such total knowledge never exists, and the results of such analyses can only be estimates of possible behaviour. The accuracy (or inaccuracy) of such analyses were demonstrated clearly by an exercise reported by Kay et al (1991). In this case several methods of
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analysis were used to predict subsidence behaviour in advance of mining. The results showed a wide variation in predicted behaviour, and also indicated that predictions using the same analysis program could be different for different users. The recommended approach is to repeat the analyses several times using the likely range of rock mass properties to obtain different sets of results or predictions of behaviour, and then to apply engineering judgement to these data. Analysis of an existing situation, where conditions and behaviour are known, preferably quantitatively, should be carried out as part of this exercise. These analyses or ‘back analyses’ serve to calibrate the model. A further suggestion is to use a combination of different methods of analysis. This philsophy of the application of numerical analyses to mining problems was described by Diering and Stacey (1987). In this way the most probable behaviour will be bracketted by the results of the range of analyses. Too often analyses are carried out without the requisite engineering understanding and interpretation. The analyses indicated above would involve the modelling of various mining stages, if appropriate, and hence the determination of deformations in shaft locations at these stages, and determination of the stresses which would be induced along the shaft locations. Comparison of these data with those from the calibration analyses, and with specified limits of stress and deformation, will allow decisions to be taken on the stability of the shafts and the measures which must be taken to ensure either their stability or their replacement. The same approach can be adopted for the assessment of the stability of crown and other pillars. Failure of a pillar can result in an airblast, which can be extremely hazardous to safety, and may lead to major damage to mine infrastructure and significant dilution of the ore. Failure of a crown pillar may be particularly disastrous since, in addition to these aspects, the underground workings will be exposed to much greater water inflow from surface as well as associated dilution from fine, weathered surface material. The potential for mudrushes then increases substantially.
Dilution Dilution considered here is in the context of interaction between excavations, and excludes consideration of the dilution which commonly occurs, for example, in caving mines or due to scaling of waste and low-grade ore from stope surfaces in many other types of mining. In mining operations which have progressed from an open pit to underground activities, the risk of dilution is mainly that of failure of the pit walls, which are in waste material, into the base of the pit and underground workings. Evaluation of the risk of dilution, and the potential volume of dilution in interaction situations, is based on the assessment of stability of pit slopes, and of pillars, as described above.
Mudrush and airblast potential The occurrence of mudrushes requires the simultaneous presence of four factors – mud forming materials, water, disturbance in the form of mining or drawing, and a discharge point. If any one of these is not present, a mudrush cannot occur. Warning signs of the potential for large flows are the occurrence of small mudflows, variability in the volumes of underground water being handled, and unexplained reduction in the quantity of water being pumped. The last could indicate that the muckpile has been plugged by mud, with water being retained in the muckpile above. Poor draw control can also lead to mudrushes. A thorough risk assessment will determine whether a mine is mudrush prone or not, and thus determine whether preventative measures are required. Such measures would include defined procedures, monitoring of pumping rates, reporting of any minor mudflows, and strict draw control.
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For an airblast to occur an open excavation must exist, with the potential for a crown pillar or stope back to collapse suddenly into a stope which contains a large volume of air. This type of situation can develop in underground open stoping mines, or caving mines in which over drawing takes place and the stope back hangs up over large spans. In such cases, sudden collapse of the pillar or back leads to the displacement of the volume of air beneath the back. This occurrence can be extremely hazardous and may result in severe damage - mine equipment such as LHD’s may be annihilated, ventilation doors and seals destroyed, and conveyor systems and shaft steelwork and equipment ripped out. It should be noted that airblasts can also be associated with mudrush occurrences - sudden displacement of air occurs when the mud flows out rapidly. No publications have been encountered which predict the effects of airblasts, such as the velocities of air flow in various tunnels and other openings within the mine as a result of a back collapse. A possible method for this is suggested as follows:
• assume that high air pressure is instantaneously generated by
compressed air. As a result of pressure differences, air begins to accelerate from the main stope along the tunnels, and into other excavations, until either the speed of sound, or the limiting velocity given by Atkinson’s formula, is reached. The mass of air leaving each exit from the main stope is calculated for a small time increment and the changes in mass, density and pressure calculated in each of the elements of the model and hence the air flow velocity. The cycle is repeated for a series of time steps to model the development of air flow throughout the mine.
1
Tunnel 1 Tunnel 2 4
0 (Main Excavation)
2 (Atmosphere)
Tunnel 5
a change in volume of air caused by the falling rock: P1V1 = P2V2
3
where
Tunnel 3
P1
= average air pressure before start of air blast
P2
= average air pressure developed at start of air blast
V1
= volume of main excavation
V2
= remaining air volume at start of air blast
Conservation of mass: Change of mass in any excavation = Volume entering – volume leaving Newton’s laws of motion: Velocity (t + Dt) = Velocity (t) + a.Dt where t
= time
Dt
= time increment
a
= acceleration = resultant force F / mass M
M
= length of tunnel L × cross sectional area A × air density w
F
= driving force of pressure difference – resisting force of friction
Atkinson’s formula for air flow: For steady state air flow, driving and resisting forces are equal in a long tunnel: A.DP = K.C.L.V2.w/1.2 where DP = pressure difference between the ends of the tunnel A
= cross-sectional area of the tunnel
K
= friction factor
V
= air velocity in the tunnel
L
= length of the tunnel
w
= air density
Tunnel 4
FIG 2 - Model of mine for airblast analysis.
• assume that the following laws determine the flow of air:
The results of the simple model are given in Figures 3 and 4. The air pressure in the main excavation (0) decreases gradually as the air flows out (Figure 3). Excavation 3 is relatively small and the pressure builds up rapidly since air is entering the excavation faster than it is flowing out. Excavation 2 represents the flow of air out of the mine into the atmosphere. The air pressure in ‘Excavation 2’ therefore remains constant. The air pressure in the remaining excavations builds up gradually. After approximately two seconds, the pressure in all excavations has equalised, and the pressure gradually reduces until all the air that was compressed has flowed out of the mine. Figure 4 shows the velocity profiles in the tunnels. The velocity in Tunnel 1 increases rapidly owing to its large cross-sectional area and the slow build-up in pressure in Excavation 1, and the peak velocity of 300 m/s is reached. Air rapidly flows through Tunnel 3, the velocity peaking above 200 m/s. The velocity drops quickly, however, due to the rapid build up of pressure in Excavation 3. Less dramatic increases in air velocity are observed in the remaining tunnels. The velocity in Tunnel 2 reaches almost 100 m/s and this is sustained for a long period of time. These velocities are a serious safety hazard and are likely to cause extensive damage to equipment and mine infrastructure. The model assumes that the air pressure due to the airblast is generated instantaneously. In practice, the air pressure will develop rapidly, but not instantaneously, and therefore the model probably overestimates the air velocities. Errors are expected to increase as pressure differences and velocities increase, but it is believed that the approach is useful for order of magnitude calculations and assessing potential risks and risk areas.
Selection of mining method
assume that the air velocity is limited to 300 m/s, the speed of sound at 1 bar, since an increase in velocity above this value would probably be limited by turbulence. The above assumptions can be applied to the modelling of air flow through a mine. As an example, a simple model is shown in Figure 2. A stope back collapse results in the primary volume of
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2 (Atmosphere)
The underground mining method chosen can have a very important influence on the existing mine surface and underground infrastructure. Caving methods will definitely cause a cave crater to be formed. This crater may grow in size as mining depth increases, and hence its influence will extend. Cut and fill and stable open stoping methods preserve the stability of the rock mass and therefore of the mine infrastructure. However, collapse of unstable open stopes may lead to development of
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7.E+05
6.E+05
5.E+05
Pressure (Pa)
4.E+05 Excavation irn 0
3.E+05
1 2
2.E+05
3 4
1.E+05
0.E+00
-1.E+05
-2.E+05 0.01
0.1
1
10
100
Tim e (s)
FIG 3 - Air pressures due to the airblast.
400
300
200
Velocity (m/s)
100 Tunnel 1 0
2 3 4
-100
5
-200
-300
-400 0.01
0.1
1
10
100
Tim e (s)
FIG 4 - Air velocities due to the airblast.
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further failure within the rock mass, with subsequent impact on mine infrastructure. The potential influence of such behaviour must therefore be analysed carefully as part of the evaluation of alternative mining methods.
CASE HISTORIES OF INTERACTION PROBLEMS In the following section several case histories of interaction between various mining excavations are described briefly to illustrate the approaches to the problems, the provision of information on which decisions could be made, and the importance of such planning decisions to the mining operation.
Open pit/ block cave interaction Large-scale operation The mine in this case study is comprised of both an underground block cave and an open pit, mining the same contorted tabular orebody. The open pit has traditionally mined the upper levels of the orebody while underground have mined the lower levels of the orebody. As the pit has moved in an easterly direction, the caving operations have followed beneath the pit with the open pit slopes subsiding substantially. Cave cracks from underground are expressed on the surface of the pit and have at times resulted in large-scale failures. Although mining in the pit has generally moved ahead of the caving operations, haul roads and ramps can be seriously affected. Of prime importance in controlling the stability of the slope, is the draw control, and if managed correctly, ie an even draw, the effect on slope stability can be minimised and the slopes managed accordingly. Where cracks and sink holes did occur within the open pit, and depending on accessibility, these areas were filled with dump rock in order to prevent and retard ingress of surface water. In considering the mine design, care must be taken to ensure that cave cracks do not intersect water courses such as streams or perimeter drains in order to prevent ingress of water and/or mud rushes to the workings. With the block cave encroaching on the open pit, and in particular the open pit sump, control of surface water to the pit is of paramount importance. During the early-1990s, an exceptionally heavy rainy season was experienced, during which time the pit perimeter drainage system was inundated with large flows into the pit. Production during this period was suspended from certain underground areas, while others were closed off until the end of the rains. As caving operations approach the open pit, consideration must be given to blasting activities in the pit as it has been found that the large pit blasts have been instrumental both in damaging underground development and releasing nitrous fumes to the underground excavations.
Slope monitoring was carried out on a regular basis in order to ensure that displacements were within the design limits, while cracking was visually identified and monitored accordingly. Mining of the open pit was successfully completed with minimal disruption to the operations while the down dip cave continues to produce. Major failure and sloughing of the hangingwall pit slopes took place during the first rainy season after completion of pit operations.
Shaft and crown pillar stability in an open stoping and sublevel stoping mine Although the mining operation involved in this case study did not involve any open pit mining, it is included here since a collapse through to surface created a large open hole, with underground facilities in close proximity – effectively an open pit and underground mining in close proximity. Early production from this mine was from open stoping operations which resulted in the formation of large open excavations close to the surface (Stacey et al, 1991). The 18 m thick crown pillar above the largest orebody collapsed suddenly, resulting in an airblast, which caused considerable damage to the shaft system. A crater was formed at the surface. Following this event there was concern that, with mining progressing to deeper levels, the stability of the shaft system might be at risk, and that collapse of the crown pillar above one of the other orebodies might also occur. Consequently an investigation was carried out with the following aims:
• to back analyse the nature and cause of the crown pillar failure;
• to assess the stability of the crown pillar above a secondary orebody;
• to evaluate the changes in stress in the footwall rock adjacent to the shaft region and on a major geological contact as a result of future mining, and hence to assess the stability of the footwall;
• to estimate possible shaft deformations and potential misalignment due to future mining. The orebody host rock is talc-carbonate schist, which is massive though strongly foliated, the foliation dipping steeply. In close proximity is a sheared contact between the talc-carbonate rocks and the footwall greenstone, dipping in the range 60° to 80°. From the results of laboratory strength testing and field observations, the talc-carbonate rock mass may be described as competent though relatively weak, and strongly foliated. The shear strength of the intact material is low in the direction of the foliation. The greenstone rock consists of high strength rock material, but the mass is well jointed.
Problem definition and analysis Medium-scale operation This medium sized case study was designed to operate simultaneously as an open pit with a block cave. The open pit would mine the up dip ore while underground would mine the down dip sections. The cave would however impinge on the open pit workings during pit operations, and draw control was paramount to successful completion. The pit design catered for the caving operation with the slope angle approximately 5° flatter than it would have been for a slope with no cave. Some disruption of the pit operations due to slope failure and the developments of sinkholes from the cave through to operating benches was experienced. The sinkholes posed a risk to equipment operating on the benches in the pit, but with improved draw control and management of the underground production blocks, the problems were alleviated.
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The problem is complex since it is three-dimensional in geometry and rock type, contains the major geological contact, and the rock mass structural characteristics are significantly different in the two rock types. Two important items of data available for the evaluation of the problems were the geometry of the collapsed area and the geometry of the mining layout at the time of the crown pillar collapse. The approach adopted was to carry out several stress analyses so that the collapse could be ‘back-analysed’. Four methods of analysis were used in all in an attempt to provide cross checks. These were three-dimensional boundary element stress analysis, finite element axisymmetric stress analysis, and three-dimensional and two-dimensional displacement discontinuity element analysis. The main analyses were carried out with the first two approaches.
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The geometry of the stopes was simplified for the purpose of analysis, but it was considered that this would not detract significantly from the realism of the simulations.
Results of the analyses The three-dimensional and axisymmetric simulations both indicated the existence of tensile and shear stresses in the weathered talc-carbonate crown pillar (which collapsed) of the major orebody. The vertical displacements at the surface calculated from the three-dimensional model indicated that regional uplift occurs as a result of the excavation of the stopes. In contrast, the results of the analyses also indicated that the surface above the major orebody would subside. The initial crater formed when the depth of mining was about 125 m and it was at this time that the back analyses were carried out, and predictions made of the extent of the crater for greater mining depth. In Figure 5 the actual crater geometry, for a mining depth of 280 m, is compared with the extent of the crater predicted from the
three-dimensional stress analysis, for a mining depth of 255 m. The talc-carbonate/greenstone contact was not modelled in the three-dimensional analysis. It can be seen that, notwithstanding the coarseness of the analyses, the agreement between the predicted and observed crater outlines is remarkably good. Resistance to shear in the tensile zone would be provided solely by cohesion of the rock mass. Frictional resistance could not occur in the tensile zone as there are no compressive stresses normal to potential failure planes. Maximum predicted shear stresses in this region were of the order of 2 MPa. The average ‘indicative’ cohesive strength measured on the samples of unweathered talc-carbonate taken from a depth of about 200 m below surface was 2.6 MPa. The cohesive strength of the weathered rock would be less than this value, and it is therefore expected that the shear strengths and calculated shear stresses would be of very similar magnitude. It was concluded from the results of the above analyses that
FIG 5 - Actual and predicted crater geometries resulting from collapsed crown pillar.
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the crown pillar failure was due to:
CONCLUSIONS
• the presence of a large tensile zone above the major orebody; • the low shear strength of the talc-carbonate rock mass parallel to the direction of foliation;
• the shear stresses induced above the stopes, which probably exceeded the shear strength of the talc-carbonate.
• It was concluded that the approach would be satisfactory for the further predictions required. The three-dimensional stress analysis indicated that a tensile zone occurred in the crown pillar of the secondary orebody. However, the magnitudes of both tensile and shear stresses calculated were considerably lower than those which resulted in failure of the major orebody crown pillar. No significant ‘subsidence’ of the surface was indicated by the analysis. It was predicted that the secondary orebody crown pillar would be stable. The close proximity of surface installations and shafts to the greenstone contact emphasised the importance of the stability of the footwall. It is often the reduction, rather than the increase, in stresses that causes the problems in mass mining operations. This is due to the fact that relaxation of the rock mass, and a reduction in confinement occurs, resulting in the loosening of blocks and in potential instability. The changes in normal and shear stress acting on potential failure planes in the footwall were determined from the analyses. The results showed that there was a large margin of safety against sliding on any of these potential shear planes in the footwall. In all cases the angle of friction required for stability was less than 10°. The analyses also showed that there was little likelihood of any toppling failure in the footwall. It was therefore concluded that the footwall would be stable, and that the safety of the shafts and surface structures was not in jeopardy. The stress analyses enabled the magnitudes and directions of the potential footwall displacements to be predicted at several locations relative to the stopes and shafts. The results indicated very small displacements of the footwall towards the stopes, reducing with distance from the stopes as would be expected. It was concluded therefore, that shaft movements would not represent a problem with future mining.
Comments The case study demonstrated that the use of coarse stress analysis models was adequate for the purpose. The agreement between the observed crown pillar failure and the failure predicted from the analyses was very good. This allowed the same models to be used with confidence for the evaluation of the effects of future mining. From a practical point of view, the predicted stability of the footwall rocks, the surface installations and the shafts, has subsequently been proved.
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This paper has highlighted aspects which must be considered in the transition from open pit to underground mining. Aspects considered to be of particular significance include:
• An economically designed pit will have slopes that are close to their stability limits. There is little scope for extending the open pit mining to greater depths, other than with a pushback.
• The planning and implementation period for transition from surface to underground mining can take as long as 20 years. Planning must therefore commence at an early stage, particularly if conditions are as in the first point above.
• Surface and underground infrastructure is often at risk due to deepening of pits, underground mining below pits, and deepening of underground mining beyond planned depths.
• Transition from open pit to underground mining often introduces the risk of mudrushes from within the rock mass (if mud-forming minerals are present), and from sumps and surface dams. Airblasts can occur as a result of underground collapses or in association with mudrushes.
• The presence of an abandoned pit above underground workings can lead to greater risks of dilution and mudrushes.
• The choice of underground mining method has a major effect on the stability of the surface. Stability requirements may therefore dictate the choice of mining method.
ACKNOWLEDGEMENTS The permission of the Management of Epoch Mine, Bindura Nickel Corporation, to publish the information contained in one of the case studies has been previously granted and is again gratefully acknowledged. The method of analysis of air flow resulting from an airblast was originally developed by Dr J A C Diering and updated by Mr W Joughin. Drs Diering and Rigby participated in some of the case study work, and their contribution is acknowledged.
REFERENCES Diering, J A C and Stacey, T R, 1987. Three-dimensional stress analysis: a practical tool for mining problems, in Proceedings APCOM 87: 20th International Symposium on Application of Computers and Mathematics in the Mineral Industries, Vol 1: Mining, Johannesburg, pp 33-42 (South African Institute of Mining and Metallurgy). Kay, D R, McNabb, K E and Carter, J P, 1991. Numerical modelling of mine subsidence at Angus Place Colliery, in Proceedings Symposium Computer Methods and Advances in Gemechanics, (Eds: Beer, Booker and Carter), pp 999-1004. Balkema. Stacey, T R, Diering, J A C and Rigby, N R, 1991. Stability predictions based on back analysis of collapsed crown pillar, Epoch Mine, Zimbabwe, in Proceedings Symp African Mining ‘91, Harare, Zimbabwe, pp 55-60.
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Quantifying Geotechnical Risk in the Mine Planning Process T P Horsley1 and T P Medhurst2 ABSTRACT In many cases mine design is the most important driver of the profitability of a mining project, however to reach an ‘optimum’ mine plan many compromises have to be made. Some of the fundamental profit (and loss) drivers relate to stoping methods and stope design parameters. Inevitably trade-offs have to be made between the stability of underground openings, ground support costs and production efficiency. This paper looks at the role of geomechanics in the mine planning and design process. The concepts of ‘controllable losses’ and ‘uncontrollable losses’ are developed as a means of partitioning ground control costs. A stope design example is given which provides a method of examining mining costs, and includes a decision making element in order to quantify geotechnical impacts and select stope dimensions. The method also provides a general framework as a means of balancing ground control costs against maximising returns from an underground mining project.
INTRODUCTION The mine planning process needs to take account of many, often competing, factors in order to achieve the ‘optimum’ mine plan. This invariably involves compromise in balancing costs, revenues, risks and operational flexibility. 1.
MAusIMM, Principal Mining Engineer, Australian Mining Consultants Pty Ltd, Level 11, 135 Wickham Terrace, Brisbane Qld 4000. E-mail: [email protected]
2.
Senior Geotechnical Consultant, Australian Mining Consultants Pty Ltd, Level 11, 135 Wickham Terrace, Brisbane Qld 4000.
For a given scenario, it is convenient to assume that the value of the ore in the ground is fixed and to treat all downstream activities and costs as losses. The focus on the mine planning and operations can then be directed at minimising these losses. Figure 1 shows the magnitude of the various loss components for a small underground gold mine. Mine planning and design, with an associated understanding of rock mass response plays a fundamental role in determining the magnitude of this cash margin. Relatively small changes to many of the loss components can have a dramatic effect on the bottom line. Some of the key operating parameters that impact on losses are set in during initial development of the mine or new ore block. These parameters (mining method, sublevel interval, cut-off, etc) will set the limit to which the operators can perform. Getting these parameters wrong or ‘not quite right’ can make a considerable impact on the viability of a project.
MINE DESIGN AND STABILITY ISSUES Underground mine design requires an iterative process which is often approached in different ways. Initially, possible mining methods are selected, likely production capacities determined and indicative cut-offs derived. Preferred mining methods are selected followed by more refined ore definition. Primary infrastructure options are then established and the analysis taken through to global extraction sequencing. Much evaluation has to be done before design commences. In many cases, the proposed mine will have been ‘regionalised’ into stoping blocks, and the primary infrastructure layout largely fixed.
FIG 1 - Breakdown from ‘in situ ore’ value to cash margin for a small underground gold mine. (Refer to the CD-Rom for colour explanation).
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When stoping blocks are to be designed, one of the first questions often put to the geotechnical engineer is ‘what is the maximum stable opening?’ However, it is often too early and there is rarely anything like enough data available on which to base an informed decision. In this case, a conservative design is likely to be forthcoming which may well add many millions of dollars to the operating costs over the life of the mine.
STOPE DESIGN EXAMPLE An example of a benching block in a narrow orebody is shown in Figure 2. Typically, the key design decisions to be made are:
• sublevel interval, and • maximum strike span. Sublevel benching is an inherently simple mining method but many factors have to be taken into account when selecting appropriate stope geometries. There are also many options with regard to access, filling, pillars and ground support. For the purposes of this example it has been assumed that the block dimensions have been fixed at 60 m high by 300 m along strike. There will be a single central access on each level, which precludes concurrent backfilling with production. The
methodology used to select the sublevel interval and strike span is a process of elimination.
Base operating costs and revenues The first requirement is to develop a spreadsheet model to estimate the operating costs and revenues for various operating parameters. This is essential to enable both planning and geotechnical engineers to develop an understanding of the commercial impact of planning decisions. Various combinations of lift height and strike length can then be analysed in terms of cash margin for the entire block. This includes capital and fixed cost components. To simplify the presentation, the results have been shown as shaded blocks in long-section format, Figure 3. The ranges are shown as a percentage of the maximum value within the block. The results here exclude the effect of both cemented fill and dilution, and clearly show how maximum value is obtained with larger stopes. Increasing the lift height translates directly into development cost-savings. Additional costs-savings can also be achieved, although significantly smaller, by increasing the strike dimension. Any decisions made on this information will however be premature, as there are many other factors to be considered.
FIG 2 - Sublevel bench stoping.
FIG 3 - Percentage of maximum value for various benching dimensions. (Refer to the CD-Rom for colour explanation).
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Backfill
Controllable losses
The mining method model assumes a continuous mining sequence that requires each stope to be filled prior to the mining of adjacent stopes. A feature of this method is that certain geometries will require a cemented fill to prevent the backfill from running into the next stope when exposed. The point at which cemented fill is required will depend on many factors, such as the rill angle of the uncemented fill, whether stopes are tight filled and how much fill dilution is acceptable. Hence there is a transition zone, rather than a clearly defined boundary, between cemented and uncemented fill (Figure 4).
This term has been used to describe minor localised overbreak and mining quality control issues. As the heading suggests these losses can be managed by the mining parameters selected and operational practices (ground support, etc). Some overbreak and loss of ore is considered inevitable and is a trade-off between operating costs and revenues. The losses associated with dilution and recovery are generally straight-forward to estimate and should require no further explanation. When these losses are factored into the value diagram a very different picture emerges (Figure 6).
FIG 4 - Backfill requirements for various benching dimensions.
When cemented backfill is taken into consideration the value distribution chart changes (Figure 5). In this example the use of cemented fill is not considered economically justifiable.
FIG 5 - Percentage of maximum value with cemented fill. (Refer to the CD-Rom for colour explanation).
Some investigations can now be made regarding the cost of failure, but what defines ‘failure’? Rock mass failure and/or economic failure - a hangingwall might fail back to a controlling structure and become stable but the diluted ore is no longer economically viable to extract. The short answer is that ‘failure’ must be treated in the same way as all other losses (Figure 1) and translated into the economic effect on the cash margin. To do this requires some partitioning of risk. Therefore, for the purposes of this analysis, these ‘losses’ can be categorised either as, ‘controllable’ or ‘uncontrollable’.
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FIG 6 - Percentage of maximum value with dilution and cemented fill. (Refer to the CD-Rom for colour explanation).
In this example, dilution has been assumed to increase with lift height, primarily as a function of geological control and quality control of development and drilling and blasting. The dilution assumed in this model is by no means excessive (max of 1.8 m combined h/w and f/w overbreak over 60 m height) and does not take account of high-grade ore that would otherwise be displaced through the mill. Dilution comes at a high price and is often the single most important factor governing the selection of mining parameters. Various tools can be used to evaluate the macro stability issues. Many will be familiar with the empirically derived stability chart method for predicting stope stability. Originally derived by Mathews et al (1981) and refined by many since (Potvin et al, 1989), the stability chart is essentially a tool to determine the relative risk of failure occurring. It can’t be used to determine with any reliability, when failure will occur, but will allow different options to be ranked against each other. The stability number, N, is calculated from an empirical formula taking account of various factors influencing stability. This is completed for the various stope walls and the least stable (lowest number) line drawn on the stability graph (Figure 7). The least stable bench face in this example has been determined to be the hangingwall. Four points have been plotted along this line (A to D), from expected stability to expected failure. The corresponding hydraulic radius values are taken and curves plotted to show the corresponding lift height vs strike length diagram (Figure 8). The points on the stability graph now become the curves in Figure 9, which can then be superimposed over the value diagram. The diagram now shows risk (of significant failure) and return. The number of different mining geometries for evaluation has also been substantially reduced. When the low value options and
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those in the caving zone are removed not many remain. Also the fact that the block height is fixed in the example (60 m) means that only the 15 m and 20 m sublevels remain in the running. Given also that there is very little difference in value between options of the same colour there are probably only two or three geometries worth considering (Figure 10).
FIG 7 - Stability graph method.
FIG 10 - Maximum value options and stability categories. (Refer to the CD-Rom for colour explanation.)
Uncontrollable losses
FIG 8 - Lift height vs strike span showing stability categories.
FIG 9 - Percentage of maximum value showing stability categories. (Refer to the CD-Rom for colour explanation.)
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In most cases, the preceding analyses would be sufficient to determine the preferred mining option. However, the impact of ‘uncontrollable losses’ is becoming increasingly important in the determination of optimal mining methods, and as such, warrants some comment. In our example, the term ‘uncontrollable losses’ is used to describe stope failures that have a major impact on the success of the proposed method. This refers to stopes that have gone beyond the range of ‘allowable’ overbreak. In order to evaluate the impact of uncontrolled rock mass failure for any given mining geometry requires an appreciation of possible failure modes, the respective economic impact on the operation, and some means of assigning a weighting of that perceived impact. It is clearly a challenging task to assign a cost to such failures, which may or may not occur. Past attempts at addressing this problem usually rely upon estimating the probability of failure, p(f), and using this value to proportion costs (Brummer and Kaiser, 1995). However, this method suffers from a number of shortcomings in that: a suitably well-defined geotechnical environment and stability analysis problem (eg a tunnel or dam embankment) is required; a probabilistic stability analysis usually does not accommodate operational or other external factors that may influence stability; and the resulting probabilistic outcome takes no account of its relationship to mining cost impacts or otherwise. A more appropriate approach would be to evaluate a term such as probability of cost impact, p(i). To arrive at a decision of likely cost impact, a variety of ‘risk analysis’ methods have been trialled over recent years, however, most suffer from inconsistency in judgements or ‘ratings’, or fail to arrive at any sort of quantitative solution. What is required is a tool that proportions the profit and recovery costs, which might be lost by choosing a particular mining option over another (as defined by the options considered). In essence, a tool that can use as inputs, the various factors that may affect stability and mining related costs, and as outputs, simply weigh up the odds for various mining geometries.
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A review of the literature reveals that such a tool does exist. Originally developed for the purposes of military strategy in the late-1960s, the Analytical Hierarchy Process (AHP) is a simple procedure, which allows comparison of any two given options, and then systematically works through all possible combinations to arrive at a relative weighting for each option (Saaty, 1988). A major advantage of the AHP procedure (over all other probabilistic or risk type analyses) is that it contains an analytical framework in which to pinpoint and hence, avoid inconsistencies in judgements or evaluations. In fact, once a consistent set of judgements is obtained, a rigorous analysis can be performed that quantitatively determines the relative importance of the inputs in relation to the decision under consideration. The AHP procedure uses a scoring model that allows comparison of qualitative and quantitative aspects subjectively by simple pairwise comparisons. Breaking down of the problem into pairwise comparisons also facilitates a more comprehensive listing of factors to be considered. The method is ideal for spreadsheets and has been used in fields such as sociology, economics, manufacturing and more recently for mining capital investment appraisal (Dessureault and Scoble, 2000). The process is best demonstrated by example. The four main steps in the AHP methodology are: (1) model the decision problem and list the hierarchy of decision criteria and elements; (2) complete the pairwise comparisons; (3) calculate the relative priorities or weights of the elements; and (4) determine the ranking of the decision alternatives. Using the preceding example, the decision problem is proposed, which comprises three possible objectives or failure modes. 1.
Large wedge failure - block fall restricts access to broken ore.
2.
Progressive collapse - excessive overbreak that leads to abandonment of working bench, but does not impact on other lifts.
3.
Caving rock mass - extreme overbreak which impacts on subsequent lifts.
We may now wish to examine each mining option in view of these failure modes. In this case, and following the analysis on controllable losses, the three most likely options that are left (Figure 10) can be evaluated. 1.
Lift Height = 15 m, Strike Span = 25 m.
2.
Lift Height = 20 m, Strike Span = 25 m.
3.
Lift Height = 20 m, Strike Span = 30 m.
The decision hierarchy would then be structured as shown in Figure 11. To arrive at a ranking of decision alternatives, a series pairwise comparisons must be undertaken. This allows each failure mode (or objective) to be compared with the others, and how each failure mode impacts on each mining option can also be evaluated. The options weightings and objectives weighting are then determined via matrix manipulation. For the interested reader, a worked analysis is provided in Appendix 1. The final weightings are calculated by multiplying the options weightings (mining geometries) by the objectives weightings (failure modes). The options weightings describe the relative importance of each failure mode for each mining option. In this example, there are three failure modes and three mining options under consideration. Therefore, there are nine possible weighting combinations (Appendix 1). The objectives weightings describe the relative importance of each failure mode in relation to its impact on loss of ore (three possible weighting combinations). The final weightings represent a relative measure of the relative importance of ore loss due to geotechnical factors for each mining option. Options Weightings x Objectives Weightings = Final Weightings . 013 0.42 0.46
013 . 0.42 0.46
0.07 0.07 0.09 L = 15, S = 25 0.28 × 0.25 = 0.33 ⇒ L = 20, S = 25 0.64 0.68 0.58 L = 20, S = 30
The final weightings (FW) show that geotechnical factors do not have a large influence on the L = 15, S = 25 option (ie 0.09) when compared to the other options. In contrast, the weightings show an increase from 0.09 to 0.33 as a measure of the impact of increasing the lift height, or then increasing strike span, from 0.33 to 0.58. The final weightings can be used to proportion the revenue that might be lost by choosing a particular mining option over another. For example, for each mining option and each failure mode there exists a unique cost associated with rehabilitating and TABLE 1 Recovery costs for each mining option Recovery costs ($) L Wedge
Prog Coll
Caving
L=15, S=15
20 000
60 000
1 000 000
L=20, S=25
25 000
75 000
1 000 000
L=20, S=30
25 000
75 000
1 000 000
re-entering the mining area. An estimate of recovery costs for each mining option is shown in Table 1. To be useful, these recovery costs must be normalised according to the relative importance (objectives weightings) of each failure mode. The relative recovery costs for each mining option are shown below. L = 15, S = 25 $20 k $60 k $1000 k 0.07 $697k $25k $75k $1000 k × 0.25 = $701k ⇒ L = 20, S = 25 $25k $75k $1000 k 0.68 $701k L = 20, S = 30 FIG 11 - Appraisal of impact of uncontrollable failures on stope design.
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The analysis shows in this case, that relative recovery costs are similar for each mining option.
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A relative risk cost (RRC) can now be evaluated in order to determine the ranking of mining options. When evaluated against a ‘zero’ risk option the relative risk cost will converge towards the true ‘cost of failure’. The relative risk cost for option 2 (L = 20, S = 25) compared with option 1 (L = 25, S = 25) would be: RRC2 = (Profit2 - Profit1) x (FW2 - FW1) + (REC2 - REC1) x (FW2 - FW1),
where FW = final weighting (ie 0.09, 0.33 or 0.58) for each mining option and REC = relative recovery cost (ie $697k, $701k or $701k) for each mining option. The relative risk cost is then subtracted from the net profit to determine the order of ranking of mining options. In this example, the L=15, S=25 option is rated at having low geotechnical impact (FW = 0.09) and therefore will be used as the basis for comparison. The summary calculation is shown in Table 2. From Table 2, the L=20, S=25 option would be selected as the preferred option as it represents the maximum relative profit when factoring in the relative risk cost. It is important to note that the AHP is a decision-making tool, which implies that results obtained (ie relative costs) can only be used as a basis for comparison among the options considered. It cannot be used for comparison with options not included in the AHP analysis. For example, if a total of six mining geometries were analysed, then the final weightings would change relative to the others considered. The total sum of final weightings will always be equal to one. Whilst this simple example serves only to confirm what might be intuitively already known, it is noteworthy to show how such intuition can be captured in a systematic decision making process. The power of AHP analysis is that it simply maps the mind process, ie reducing the problem to a series of pairwise comparisons. Owing to this unique feature, judgements can also be evaluated for consistency, which becomes very important when more than one or two options need to be considered and possible impacts are many.
CONCLUSIONS Mine planning and design have a fundamental impact on mine profitability. Analysis tools and methodologies are essential in order to assess the commercial impacts of mine design and
planning options. The overall process here uses established mine design and costing methods in most part. The difference is that most of the information is used in a more efficient way, allowing a combined focus on both mining and ground control issues. Only the uncontrolled failure component is new, and has been suggested as a possible approach to overcome the limiting scenario of ‘will it stand up or fall down’. It attempts to answer the real question underlying the planning philosophy, which is ‘what impacts do geotechnical factors have on any proposed mining option’. This impact must be measured in terms of cost. The example shown here uses the stope stability chart. However, any stability analysis or design method may be used depending on the preference of the designer and mining method chosen. The important feature is that the geotechnical engineer can use any design chart of their making (ie it could be based on numerical modelling, a mine specific empirical design model or published guidelines), which delineates geometry vs factor of safety (or probability of failure). Once the design chart is obtained, mining costs can then be overlayed. Risk management is becoming increasingly important, not least from the safety perspective. Uncertainty has always been a factor in mining and many decisions traditionally rely on the experience and judgement of operators. In the current environment is this acceptable? In the civil engineering industry once off infrastructure projects can support a very high factor of safety. If similar standards were applied in mining the industry would not be viable.
REFERENCES Brummer, R K and Kaiser, P K, 1995. Risk-cost benefit analysis for support design in underground mines, Trans Instn Min Metall (Sect A Min Industry), 104:A71-75. Dessureault, S and Scoble, M J, 2000. Capital investment appraisal for the integration of new technology into mining systems, Trans Instn Min Metall (Sect A Min Industry), 109:A30-40. Mathews, K E, Hoek, E, Wyllie, D C and Stewart, S B V, 1980. Prediction of stable excavation spans for mining below 1000m in hard rock, CANMET Report No 802-1571. Potvin, Y, Hudyma, M R and Miller, H D S, 1989. Design guidelines for open stope support, CIM Bulletin, 82:53-62. Saaty, T, 1988. The Analytical Hierarchy Process, (McGraw-Hill: New York).
TABLE 2 Ranking of mining options with consideration of uncontrolled failure
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Mining Option
Net Profit ($M)
Recovery Cost ($M)
Final Weighting
Relative Risk Cost ($M)
Relative Profit ($M)
L=15, S=25
4.09
0.697
0.09
0.00*
4.09
L=20, S=25
4.37
0.701
0.33
0.07
4.30
L=20, S=30
4.42
0.701
0.58
0.16
4.26
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APPENDIX 1 Example of mining options analysis using the Analytical Hierarchy Process (AHP) method The execution of the AHP is based on evaluating the pairwise comparison matrix (PCM) by comparison of the decision elements (a). Pairwise comparisons can be made using a simple rating system, as shown in Table A1. In this example, there are three failure modes and three mining geometries. Therefore, the PCM will have three rows (i = 1,2,3) and three columns (j = 1,2,3). The single objective matrix is completed in Table A2 to show how such judgements might be made. For example, a caving rock mass is absolutely more important to its impact on losses than a TABLE A1 Value of decision elements. Value of aij 1
Interpretation Objectives i and j are equally important
3
Objective i is slightly more important than j
5
Objective i is strongly more important than j
7
Objective i is very strongly more important than j
9
Objective i is absolutely more important than j
large wedge failure, but only slightly to strongly more important than progressive collapse. The next step in the procedure is to normalise each column by division of each value within a column by the sum of the column. The score of each objective is the average of the rows. Only a single PCM must be completed for objectives (where they are rated against each other), but three more PCM’s must be generated to evaluate how each failure mode impacts on the chosen mining options. Table A3 shows how each mining option might be compared for each of the failure modes. In this case, the large wedge and progressive collapse objectives were considered to have the same impact for each mining option, whereas caving was rated with a higher impact. It is evident from Table A3, that increasing lift height from 15 to 20 m raises the impact of losses for the first two options, and increasing strike span has only a slightly greater effect. For caving, the final scores show the relative impacts on both lift height and strike span. For a 3 x 3 matrix it is relatively easy to have consistency in the judgements - it would be inconsistent, for example, if it was judged that the impact of progressive failure and its likelihood would be reduced by increasing lift height. An important feature of the AHP is the ability to determine a consistency index (CI) that can evaluate the consistency of the judgements. The first step in calculating the CI is to calculate the w matrix from the objective matrix and final scores. The calculation for wedge failure is shown below.
TABLE A2 Comparison and weighting of objectives (failure modes). Pairwise comparison L Wedge
Normalised PCM
Prog Coll
Caving
L Wedge
Prog Coll
Caving
Average of normalised
L Wedge
1
1/4
1/9
0.07
0.06
0.08
0.07
Prog Coll
4
1
1/3
0.29
0.24
0.23
0.25
Caving
9
3
1
0.64
0.70
0.69
0.68
TABLE A3 Options weightings for each failure mode. Large Wedge PCM L=15, S=25
L=20, S=25
L=20, S=30
Normalised
L=15, S=25
1
1/3
1/4
0.125
0.14
0.12
Final Score 0.13
L=20, S=25
3
1
1
0.375
0.43
0.44
0.42
L=20, S=30
4
1
1
0.50
0.43
0.44
0.46
Progressive Collapse PCM L=15, S=25
L=20, S=25
L=20, S=30
Normalised
L=15, S=25
1
1/3
1/4
0.125
0.14
0.12
Final Score 0.13
L=20, S=25
3
1
1
0.375
0.43
0.44
0.42
L=20, S=30
4
1
1
0.50
0.43
0.44
0.46
Caving Rock Mass PCM L=15, S=25
L=20, S=25
L=20, S=30
Normalised
L=15, S=25
1
1/5
1/7
0.08
0.05
0.10
0.07
L=20, S=25
5
1
1/3
0.38
0.24
0.23
0.28
L=20, S=30
7
3
1
0.54
0.71
0.67
0.64
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. 0.39 1 1 / 3 1 / 4 013 w = ( PCM )3 x 3 × ( score )3 x 1 = 3 1 1 × 0.42 = 1.27 1 0.46 1.40 4 1 A single number, λ, is then calculated from the values found in the score matrix and the w matrix. λ=
w11 S11
+
w 21 S 21
k
+
w 31 S 31
=
0 . 39 0 . 13
+
1 . 27 0 . 42
+
1 . 40 0 . 46
3
112
λ − k 3.01 − 3 = = 0.005 k −1 3 −1
TABLE A4 Random indices for consistency check. k
2
3
4
5
6
7
8
9
10
RI
0
0.58
0.9
1.12
1.24
1.32
1.41
1.45
1.51
= 3.01
where k is the number of objectives. The λ value is then used to find the consistency index. CI =
CI can then be compared with a random index (RI) developed by Saaty (1988) and is shown in Table A4.
The consistency ratio, ie CI/RI, is used for determination of the consistency of the judgements: CI = 0, perfectly consistent; CI/RI < 0.10, consistency is satisfactory; CI/RI > 0.10, serious inconsistencies exist. In our example, CI/RI = 0.005/0.58 = 0.008. A consistent set of judgements has therefore been obtained for evaluating the relative importance of wedge failure to each mining option. The consistency check should be completed for each of the options (mining geometries) PCM’s and the objective (failure modes) PCM.
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Dilution Control in Southern African Mines R J Butcher1 ABSTRACT The economic viability of many Southern African massive mining operations is threatened by excessive dilution. In many operations currently in progress, some stopes are experiencing as much as 115 per cent waste ingress reporting to drawpoints. This paper examines the magnitude, causes and types of dilution and presents a set of general principles for the control of dilution in massive stoping operations.
INTRODUCTION All mining operations experience dilution at some time or another and the elimination of all waste ingress in most cases is impossible. However, experience has shown that dilution can be controlled to acceptable levels by the implementation of correct mining engineering principles (Butcher, 1997). This paper examines the magnitudes, causes and classification of dilution and proposes the reduction of excessive waste ingress through the application of the ‘define, design, draw’ principle. Based on Canadian experience (Pakalnis et al, 1995), dilution greater than 20 per cent is defined as excessive dilution.
MAGNITUDES AND CAUSES OF DILUTION A survey of massive mining operations in Southern African (Butcher, 1999a) has provided the information on dilution magnitudes and trends given in Table 1 and Figure 1.
1.
Senior Mining Engineer, SRK Consulting, 265 Oxford Road, Illovo, Johannesburg, South Africa.
TABLE 1 Dilution magnitudes associated with different mining methods. Mining method
Dilution %
Remarks
18 - 115
Unsupported
Sill and bench
5 - 48
Unsupported
Continuous undip retreat stoping
27 - 48
Unsupported
Sub level open stoping (SLOS)
Open benching
< 20
Unsupported
Fissure/vein mining
±40
Unsupported
Creeping cone
< 10
Unsupported/ Artificially supported
Cut and fill (CAF)
5 - 15
Artificially supported
Vertical Crater retreat (VCR)
10 - 38
Artificially supported
From the data in the table and figure the following can be concluded:
• From Figure 1 the average dilution is in the region of 40 per cent. If it is assumed that the average massive mine produces 500 000 tons per annum at a cost of $30/ton then the impact of this 40 per cent dilution rate would be about $6 million per annum. An approximate estimate of dilution costs per annum for three commodities is given in Figure 2. The cost per ton
FIG 1 - Dilution variation with time.
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information used in these estimates is based on real data obtained from international mining operations. This simple calculation illustrates the potential cost savings that can be made through the implementation of a dilution control strategy.
• Table 1 indicates that mining methods with unsupported stopes have higher dilution rates than artificially supported methods. This could be due to the lack of support afforded to incompetent stope rockwalls. It could be concluded that more orebodies should be extracted using mining methods involving artificially supported stopes (assuming that dilution control is the main design rationale).
• The data from Figure 1 (Butcher, 1999a) indicate that dilution levels vary with time. In this respect there could be two possible explanations: • variable stope rock wall competencies resulting in variable dilution levels; • poor mining practices due to: - poor blasting, resulting in the overbreak of the stope boundaries; - dilution due to a regular stope boundary profile being maintained when the orebody width fluctuates; and - poor mining discipline associated with a lack of draw control procedures. In summary, the main causes of excessive dilution are:
• incompetent ground conditions; • inappropriate mining methods; • poor mining practices.
THE ORIGIN AND CLASSIFICATION OF DILUTION In addition to the planned and unplanned types of dilution described by Dominy et al (1998), dilution can be classified according to its origin (Butcher, 1999a). Three types of dilution tend to affect massive stoping operations (see Figure 3).
Top dilution This can be defined as waste rock or ore which is of uneconomical value. This type of dilution normally occurs when crown pillars are wrecked or if sloughing of the back occurs during stoping.
Internal dilution This is the waste rock or low-grade ore that occurs within defined economic orebodies at the stope boundary (for example, shale floaters in a kimberlite orebody). This type of dilution can be thought of in a similar manner to the internal waste between reef bands. Internal dilution is the most difficult type of waste to control due to its close proximity to the ore. In certain cases, internal dilution can be as high as 40 per cent. This type of dilution is sometimes referred to as planned dilution (Scoble and Moss, 1994).
Side dilution
The first two points are important in understanding the origin of dilution for the purpose of formulating a prevention strategy. The last point focuses on the control and reduction of dilution.
This is the dilution that occurs due to the sloughing of the stope hangingwall and/or footwall in a steeply inclined orebody (or from the sidewalls in a massive deposit).
FIG 2 - Cost of dilution for three commodities based on an average mine production capacity of 500 000 tons per annum.
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FIG 3 - Classification of dilution.
DILUTION REDUCTION AND CONTROL STRATEGIES The magnitudes, causes and classes of dilution which affect Southern African massive stopes have been discussed above. The focus now changes to the prevention or control of the dilution using the define, design, draw principle. This principle enhances
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existing knowledge of dilution control for Southern African conditions. The method is currently being implemented at Rosh Pinah Mine, Namibia as part of the mine re-engineering process (Butcher, 1999b). There are two cornerstones to this principle:
• prevent dilution rather than control it, and • if dilution cannot be prevented, then control it.
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In essence. the define and design aspects act as the prevention component, and the draw aspect is the control portion of the principle. It may be considered that the control aspect is irrelevant since all mining methods should be designed to eliminate dilution. However, in many cases, the most dilution friendly mining methods do not fulfil the necessary tonnage requirements for operational viability. In these cases experience has shown that dilution can be reduced to tolerable levels by implementing a draw control strategy.
Dilution reduction using the define principle A major component in the prevention of stope dilution is good geological and geotechnical definition of the orebody and the surrounding country rock. The geological definition is important so that the amount of internal dilution can be determined and that the boundaries of stopes can correspond with the limits of the orebody. A common pitfall is to reduce or underestimate the amount of diamond drilling required for orebody definition. In the case of an irregular orebody, it may be necessary that drilling is conducted at intervals of less than 10 m. A study conducted in Canada (Puhakka, 1990) has shown that planned dilution decreased by ten per cent when the drilling definition interval was reduced from 25 m to 7.5 m. The geotechnical definition of the orebody and the country rock is important to:
• define weak country rock zones which could lead to dilution influx;
• define stable stope dimensions to prevent the failure of the rockmass surrounding the orebody; and
• determine the in situ rock mass strength, so that rib pillars can be correctly designed and stope spans limited, thus preventing dilution. Geotechnical definition can be accomplished with the use of a rock mass classification system incorporating the effects of blasting (Laubscher, 1990). It is important that geotechnical classification is not only carried out in the project planning stages, but also on an on-going basis during the mining stages, with the compilation of geotechnical plans being an essential part of the program. A useful exercise is the back analysis of dilution levels from old stopes and correlation with the classification values (Butcher, 1997). The purpose of this exercise is to identify high dilution geotechnical areas. Cavity monitoring systems can also aid this purpose (Gilbertson, 1995).
(1994) deals with the dilution associated with different types of mining methods. Taking cognizance of Elbrond’s observations and experience in Southern Africa (Butcher, 1997 and 1999a), the most suitable mining methods for dilution reduction are summarised in Table 2. The table indicates that cut and fill mining is the most dilution friendly mining method. However, this method tends to have the highest dollars per ton mining cost and the lowest production capacities. Figure 4 shows the potential cost-savings associated with changing the mining method from unsupported to artificially supported methods. With regard to the choice of mining method there is a dilemma in that a high tonnage/low mining cost method may be required due to the grade of the orebody, but in order to control dilution the preferred method may have the opposite characteristics. A compromise can be attained by using smaller open stope dimensions or a shrinkage method such as a creeping cone (Aplin, 1997). In these cases dilution can be reduced further by accurately setting out the stope boundary to the orebody contact. The non mining of the ore blocks with extremely complex and weak geologies can also assist. These measures will reduce the quantities of side and internal dilution (Butcher, 1997). The levels of side dilution can further be reduced by the implementation of good drilling and blasting practices (Dominy et al, 1998). One of the most common causes of dilution is poor drilling and blasting. The problem is essentially a design issue, although there are also control issues associated with it. Correction of the problem is difficult and time consuming. From experience with blasting projects on Zimbabwean mines in the late-1990s, a drilling and blasting improvement programme normally takes at least 24 months to show substantial results (Butcher, 1997). Although drilling and blasting aspects are beyond the scope of this paper, the following factors should be considered during the project design stage:
• the design of blasting rings/fans with blast hole lengths not exceeding 15 m, thus reducing potential blast hole deflection and stope rockwall damage;
• the design of stope fans/rings with small blast hole diameters rather than large diameters, hence reducing the charge per delay and the blasting damage;
• the determination of the correct powder factor; • the initiation of fans/rings using Nonel or Electrodet systems instead of detonating cord;
• the design of stope fans/rings using computer models to determine the effects of different charge lengths and timing systems;
Dilution reduction using the design principle The next stage in the dilution prevention strategy is the selection of the most dilution friendly mining method (taking cognizance of stable stope spans and pillar rockmass competencies). Elbrond
• the setting of realistic drilling and blasting targets, thus ensuring quality blasting;
TABLE 2 Dilution associated with different mining methods (modified after Brady and Brown, 1993). Orebody geometry
Rockmass competency
Dilution hazard
Massive/regular
Good
Little to none
All methods
Massive/regular
Poor/medium
Considerable
CAF/creeping cone, SLOS (with small stopes and post fill), VCR
Irregular changes from massive to narrow (eg vein) for example, over small strike distances
Good/medium
Internal and side dilution (due to stope boundary)
CAF, creeping cone, VCR
Irregular changes from massive to narrow, for example over short strike distances
Poor
Considerable side, internal and top dilution
CAF
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FIG 4 - Cost of dilution attributed to mining method.
• the provision for blast hole redrilling in the production planning stages, to avoid the charging of rings/fans with closed holes. This will eliminate overburdening and reduce rockwall damage.
Dilution control using the draw principle (draw control) The use of a draw control system in a stoping scenario differs from that which is used in block caving, the main focus being on grade control through the reduction of waste mining. The essential part of the draw management program is the establishment of drawpoint tonnage calls and acceptable dilution levels. In many mines these have not been determined and as a result it is difficult to ascertain the dilution level at which a drawpoint should be closed. The determination of drawpoint tonnage calls can be achieved with geological orebody modelling packages and production benchmarking exercises. The implementation of a dilution control program focuses on the prevention of excessive waste draw. In this, it is essential to have a draw control officer who regularly inspects the drawpoints, passes and tips for waste. The selection of the draw control officer is of the utmost importance and experience has shown that a very competent shiftboss is usually the most suitable person for this position. Such a person has sound knowledge of the production process. In addition, a draw control clerk is required to assist the draw control officer in compiling the relevant draw control statistics and preparing the monthly drawpoint calls. The setting of realistic monthly production tonnage calls is vital to prevent waste drawing. Experience has shown that if calls are set too high, underground production crews will draw waste to attain call. Unrealistic calls normally occur when over-optimistic forecasts of the production capability of particular mining methods are made, or when the mineral prices fall to such an extent that excessive production is required for mine viability.
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One of the main causes of excessive dilution is ignorance, and it is surprising that very few mines have dilution awareness campaigns which highlight the dangers of dilution on mine viability. An awareness program could be implemented at little cost and would involve posters at waiting places, lectures and regular reminders. A need which is sometimes overlooked is the requirement for dilution monitoring in mines which do not have excessive dilution. The main reasons for this are:
• to determine the correct level of dilution; • to ascertain whether dilution levels increase with the mining of different geotechnical areas. Even in mines which only suffer from a minor dilution problem, some form of draw monitoring is normally required to validate grades (Butcher, 1999a). Mines in this category normally overestimate the dilution level and lower stope grades accordingly. The correct mine dilution level can be controlled by a mine geologist conducting a dilution drawpoint audit on a quarterly basis. Figure 5 shows the potential cost-savings associated with dilution control.
CONCLUSIONS Excessive dilution can threaten the viability of most mining operations. In this respect it has been estimated that dilution could be costing some Southern African massive mining operations in the region of $6 million per year. However, with the correct definition of the orebody and the geotechnical environment, the most dilution friendly mining method can be selected. The implementation of a draw control system is fundamental to exercising effective control over the drawing of waste from stopes. These aspects can be summarised as the define, design, draw principle of dilution control. The application of these principles can result in major cost-savings at little cost to mining operations.
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FIG 5 - Potential cost saving associated with dilution control for mines with different capacity.
REFERENCES
BIBLIOGRAPHY
Aplin, P, 1997. Reducing dilution by the creeping cone, Mining magazine, 176:22-26. Brady, B H G and Brown, E T, 1993. Rock mechanics for underground mining, Second Edition (Chapman and Hall). Butcher, R J, 1997. SRK Consulting Report, No 248118. Butcher, R J, 1999a. SRK Consulting Report, No 259132. Butcher, R J, 1999b. Dilution control boosts viability at Rosh Pinah, African Mining, Vol 83, pp 13. Dominy, S G, Sangster, C G S, Camm, G S and Phelps, R F G, 1998. Narrow-vein stoping practice – a United Kingdom perspective, Trans Inst Min Metall, (Sect A: Min industry), Vol 107, September, pp A122. Elbrond, J, 1994. Economic effects of ore losses and rock dilution, CIM Bulletin, 87(978):131-134. Gilbertson, R J, 1995. The applications of cavity measurement systems at Olympic Dam operations, in Proceedings Underground Operators Conference, pp 13-14 (The Australasian Institute of Mining and Metallurgy: Melbourne). Laubscher, D H, 1990. A geomechanics classification system for the rating of rock masses in mine design, J S Afr Inst Min Metall, 90(10):257-273. Pakalnis, R C, Poulin, R and Hadjigeorgiou, J, 1995. Quantifying cost of dilution in underground mines, Mining Engineering, pp 1136-1141. Pilula, E M and Banda, J Z, 1994. The development of backfill mining methods at Nkana, in Proceedings XVth CMMI Congress, Vol 1, pp 177-188 (South African Institute of Mining and Metallurgy: Johannesburg). Puhakka, R, 1990. Geological waste rock dilution, Finnish Association of Mining and Metallurgical Engineers, Research Report No A94. Scoble, M J and Moss, A, 1992. Dilution in underground bulk mining: implications for production management, Mineral Resource Evaluation II: Methods and Case Histories, (Eds: M K G Whateley and P K Harvey) Geological Society Special Publication No 79, pp 95-108.
Chitombo, G and Scott, A, 1990. An approach to the evaluation and control of blast induced damage, in Proceedings Third International Symposium on Rock Fragmentation by Blasting, (The Australasian Institute of Mining and Metallurgy: Melbourne). Forsyth, W W, 1993. A discussion of blast-induced overbreak around underground openings, in Proceedings Fourth International Symposium on Rock Fragmentation by Blasting, Vienna, 5-8 July. Ouchterlony, F, 1995. Review of rock blasting and explosives engineering research at SveBeFo, in Proceedings Explo ’95, pp 133-146 (The Australasian Institute of Mining and Metallurgy: Melbourne). Persson, P-A, Holmberg, R and Lee, J, 1993. Rock Blasting and Explosives Engineering, Boca Raton, (CRC Press). Planeta, S and Szymanski, J, 1995. Ore dilution sources in underground mines interpretation and evaluation methods, in Proceedings Underground Operators Conference, pp 87-92 (The Australasian Institute of Mining and Metallurgy: Melbourne). Scoble M J, Lizotte, Y C, Paventi, M and Mohanty, B, 1997. The Measurement of Blast Damage, for publication in Mining Engineering, American Institute of Mining Engineering, Littleton, Colorado. Toper A Z, Kabongo K K, Stewart R D and Daehnke, A, 1999. The mechanism, optimization and effects of preconditioning, in Proceedings Fragblast 1999, pp 1-8 (South African Institute of Mining and Metallurgy: Johannesburg). Tsoutrelis, C E, Kapnis, A P and Theophili, C N, 1995. Determination of blast induced damaged zones in pillars by seismic imaging, in Proceedings Explo ’95, pp 387-393 (The Australasian Institute of Mining and Metallurgy: Melbourne). Vink, D M, 1995. Minimising blast damage to the extraction level of Northparkes Mine’s E26 block cave, in Proceedings Explo ’95, pp 251-260 (The Australasian Institute of Mining and Metallurgy: Melbourne).
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Reflections of a Mine Scheduler J Luxford1 • • • • • • •
ABSTRACT Scheduling mine development and underground construction has been the bane of most mine managers’ lives at one time or another. Even though scheduling software has developed significantly over the last 20 years, mining projects continue to take far longer than expected to complete. It would appear that possession of a gleaming 1000 activity bar chart generated by the latest project management software and plotted in multiple colours offers no guarantee at all that any of those 1000 activities will be completed on time, let alone the project as a whole. The author has earned his ‘masters in disasters’ with more than one project running over time and budget. This paper describes some of his experiences in scheduling major mine development and construction projects and the lessons he has learnt along the way. In particular, the paper examines a range of key issues that mine schedulers must take into account when planning mining projects. The paper also examines the development of scheduling software and particular advance rates used in scheduling mine development projects in Australia.
2.
Initial mine development for bulk sampling and diamond drilling;
3.
Permanent mine development including:
• • • • • •
INTRODUCTION The Collins Concise Dictionary defines ‘schedule’ as:
• ‘a list of tasks to be performed, especially within a set period’; or as a
• ‘plan of procedure for a project’. At its most basic, scheduling involves defining activities, assigning resources, estimating durations and linking activities within a logical framework. While this looks quite simple and straightforward, our industry continues to get it so wrong so often with mining projects running over time and budget. In the author’s experience, most mine development schedulers concentrate on the first dot point above, but often fail to address the second. In particular, schedulers are frequently unaware of the interaction of the underground development and construction activities they are planning. Nor are they often aware of the appropriate logical framework for sequencing the activities. Perhaps most importantly, many schedulers do not fully appreciate the constraints that a mining development and construction schedule must accommodate. This paper examines these issues and endeavours to offer some practical advice on how best to address them. The development and other advance rates quoted in this paper reflect what can be achieved by a mine development project when the constraints are identified and allowed for during scheduling and then well managed during construction.
1.
Site establishment including:
• site access, • earthworks, 1.
a.
workshops,
b.
Crushers, and
c.
pump stations; and
Permanent infrastructure for:
• • • • • 5.
Mine planning for:
• • • • 6.
diamond drilling, assaying and geological modelling, level layouts, and stope design.
Stoping activities such as:
• • • • 7.
ore handling systems, dewatering systems, power and water reticulation, communication and control systems, and ventilation systems.
development, ground support, drilling, and blasting.
Production!!!!!!!!!!!!!
SCHEDULING EXPERIENCES This section describes the scheduling methods used by the author on a series of projects in the last 15 years and how his approach to scheduling has evolved in that time. It concludes with a preferred approach to scheduling mine development in this new millennium.
Arrival of the computer: Ulan 1987
Principal, Luxford Mine Management Services, PO Box 1190, Booragoon WA 6154. E-mail: [email protected]
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decline development, hoisting shaft or conveyor drifts, level development, boring, sinking or raising ventilation shafts and raises; boring or raising ore passes; major infrastructure excavations for:
• boreholes for drainage and power reticulation. 4.
SCOPE DEFINITION The starting point with all scheduling is proper scope definition, as per the first dot point in the introduction. While there is a wide range of activities that may be contemplated in the context of underground mine development and construction, they will rarely all be present in the one project. For the purposes of this paper, here is a broad list of the activities that a scheduler may include in the scope when scheduling the development of an underground mining project:
portals and shaft collars, buildings, power supply, water supply, site communications, rubbish tips and waste handling, and water treatment facilities.
Every project the author had been involved in was manually scheduled until Ulan Coal (where the author was the colliery manager) bought their first PC in 1987 along with a copy of Hornet to schedule the first longwall move at Ulan. Hornet was
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one of the early scheduling packages written for the PC. Based on very dim memory, that IBM PC cost in the order of $10 000 in 1987 with Hornet costing more again. This project involved dedicating a young mining engineer to first learn how to drive the PC then master Hornet. That process took two months, followed by another two months to set up all the activities in Hornet. The mining engineer worked on night shift for the duration of the longwall move inputting data each night from the preceding 24 hours and producing the following each morning:
• updated bar chart with all activities; and • activity plans for the supervisors with all resources listed for each activity. With this revolutionary technology at hand, for the first time in his career, the author actually completed a major project on time if not budget. The key lay in understanding the trends and taking timely corrective action to maintain progress. That corrective action mainly consisted of pouring more and more resources into the project to maintain progress.
Another computer: Beaconsfield 1988 The following year, in 1988, the author learnt to use a PC and used Microsoft (MS) Project for the first time while managing the Beaconsfield shaft project in Tasmania. MS Project was used for scheduling the following:
• mine development for feasibility studies; and • various construction activities for budgeting and project control.
Scuddles 1989 The author’s third exposure to scheduling on a PC came with the Scuddles Ag-Pb-Zn mine development project in 1989. One of the author’s first tasks as manager mining at Scuddles was to review the overall mine development program. Over the next three years, numerous schedules were prepared to assist in managing the development and construction of the new underground mine at Scuddles. The main problem encountered when using MS Project (or any other time based scheduling package for that matter) was that the activity durations had to be manually calculated and errors were hard to avoid in the time consuming transcription process. In addition, it was impossible to track the actual meters of development in each heading each week so that checks could be made to prevent over estimation of development rates.
Early and mid-1990s In the immediate years after Scuddles, the author worked on a series of decline developments that were fairly simple scheduling exercises that could be easily handled in MS Project. However, during the review of one overall mine development schedule that contained serious scheduling errors, the author realised that the best way to schedule a complex mine development stretching over several years was by using a spreadsheet with rows for development headings and columns for every week.
Cannington 1996
• • • • •
ventilation, mullock handling, drainage, conflicting activities, and blasting damage.
One of the keys to the success of the scheduling at Cannington was the review process that evolved. The author worked up the initial schedules as described above. The senior project controls manager then thoroughly reviewed every aspect of the schedule and plans immediately after joining the project. This review process overcame one of the weaknesses present in too many projects where the schedulers (often junior mining engineers with limited practical operating experience) are left to work in isolation.
Didipio 2000 Didipio is a planned block caving project in the mountains about 200 km north of Manila in the Philippines. The project will involve a 7 km dewatering tunnel, 2.8 km conveyor decline and 9 km of ancillary and level development. The author has overseen a number of enhancements to the mine layout to optimise the development schedule. These enhancements have revolved around rearranging the level access and mine ventilation layouts to optimise level development rates in critical areas. The studies have been scheduled using the massive spreadsheet approach. As noted above, the great advantage of this approach is that it allows the scheduler to see all the potential problems in the mine development. However, the disadvantage of this manual approach is that revisions to the base schedule are very time consuming. The answer to this problem is discussed in the next section.
MINE SCHEDULING SOFTWARE The previous section described the author’s experiences with scheduling through the 1980s and 1990s. This section looks more generally at the evolution of the scheduling software in that time.
Software in the 1970s Scheduling methods and software have evolved considerably over the last 30 years. In the 1970s, when most scheduling was still being done by hand, the larger mining companies were starting to use ‘Program evaluation and review techniques’ (PERT) on mainframe computers. Because of the cost of mainframe computing in the 1970s and early-1980s, computerised scheduling was restricted to the major projects.
Arrival of CPM for PCs in the 1980s
The author was the project manager for the underground mine construction at BHP Minerals’ Cannington Ag-Pb-Zn mine in NW Queensland from 1996 to 98. While MS Project was used at Cannington to co-ordinate the overall mine development project, a massive Excel spreadsheet was used to schedule the mine development. Although a cumbersome way to schedule the work, the spreadsheet approach enabled the author to develop the mine
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development schedule week by week on the computer screen, while at the same time colouring increments of weekly advance on 1:1000 scale mine plans. Coloured pencils may be a very ‘low tech’ tool in this new digital millennium. However, the process of thinking through weekly advance on coloured plans, while using a spreadsheet to add (and check) the quantities identifies all the interactions and potential problems that will arise in the course of mine development, including:
The cost of computerised scheduling fell dramatically in the mid-1980s with the arrival of the IBM personal computer (PC) and the application of the Critical Path Method (CPM) of scheduling. Some of the better known CPM packages for PCs that can be traced from the mid-1980s include Artemis, Open Plan, Primavera and Microsoft Project.
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The shortcoming of the CPM programs is that they are based on time units of work. While this feature suits most construction activity where the schedules are based on time units of work and the scheduler is mainly concerned with tracking time, it does not suit the mine development scheduler. This scheduler is mainly concerned with physical quantities, such as development or drilling meters and in particular, ensuring that the quantities scheduled at any time reflect the resources available and constraints applying. Another shortcoming in the mining context is that the CPM programs cannot convert development progress into tonnes and grade. A number of manipulations have evolved to overcome the inability of these CPM bar charting programs to schedule weekly quantities instead of just elapsed time on activities. One such solution is to export the activities table from the bar charting program into a spreadsheet and then, with formulae in the spreadsheet cells, convert the timelines into numbers showing advance rates per week or day. Another solution is to use meters as units of currency for each development activity and track the physical development by way of the cash flow reports. The drawback of these approaches is that they require the manual transcription of advance rates and durations into the bar charting program for each activity. The problem is that manual transcription always introduces the risk of errors and consequently requires exhaustive checking to eliminate such errors. If the rates are too optimistic, it is a major task to reschedule the development activities.
Recent developments Underground mine development scheduling requires a package that combines the flexibility of time based logic linked bar charting software (such as MS Project) and the detail of all the physical quantities and advance rates for every activity for all weeks in a spreadsheet format. The only software that completely integrates these two functions that the author has seen is Mine Works Planner. Since being released in Australia in 1998, it is now being used on 35 mine sites around Australia and South Africa. The software is also being used by a number of mining consultants and contractors.
Power of the pencil As noted previously, the latest software is unsurpassed for scheduling all manner of underground mine development projects. However, there is still no substitute for the humble coloured pencil when trying to visualise the complex interaction of a range of activities in one area of a mine. Perhaps the most difficult exercise the author has ever attempted to schedule is the following sequence for the proposed Didipio advance undercut block cave development:
• undercut development, drilling and blasting; • extraction level development; • drawpoint and drawbell excavation behind the undercut’s advancing edge;
• drawpoint support construction; and • roadway concreting from drawpoints to the crusher. Separate 1:1000 scale plans were coloured up and overlain to ensure that the activities were happening in the correct sequence, and in the case of the drawpoint and drawbell development, that these were being excavated at least 35 m behind the advancing undercut front. This was important to ensure that the development only took place within the stress shadow created by the advanced undercut above.
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SCHEDULING CONTINGENCY The previous sections have reviewed the evolution of scheduling software and a practical approach to scheduling. This and the following sections discuss a range of factors that schedulers must consider and ultimately, what unit rates of advance or other productivities should be allowed in mine development schedules. There are two basic laws of scheduling major mine development. Rule 1 is that ‘Murphy was a miner, and what can go wrong (at the worst possible time) will go wrong’. Rule 2 is: ‘when in doubt, refer to Rule 1’. With this in mind, the mine scheduler’s defence against Murphy depends on adequate contingency allowances and alternative plans to overcome the inevitable setbacks and keep the schedule on track.
Contingency allowances Adequate contingency allowances must be provided in the time estimates for the known and unknown problems, blunders and planning omissions that will inevitably arise. The contingency allowance required is inversely proportional to the quantity and quality of planning applied to the mine development. It is common to allow 30 per cent contingency in scoping studies, reducing to 15 per cent in the detailed feasibility stage. It is also very common for the schedulers to have hidden another five to ten per cent in the detailed activities. The author’s experience is that at least ten per cent contingency is required, even when the project has been completely planned, scheduled and engineered.
Contingency planning In addition to contingency allowances, it is important to incorporate adequate flexibility into the mine plans, budget and schedules to deal with the inevitable problems as they arise. The need for such plans is particularly important in the context of the constraints described in the following section. Ventilation is one area that often disrupts mine development schedules and requires alternative strategies when the original plan fails. The author has temporarily converted more than one ore pass into a ventilation raise equipped with a booster fan in order to maintain critical mine development at the bottom of the mine. Apart from things going wrong and forcing changes, the main challenge in mine development scheduling is that the activities never quite run on schedule and in the sequence anticipated at the time the plan was prepared. As a result, from time to time, the scheduler has to rearrange the sequence or resourcing of critical activities to keep the project on time.
Critical paths and the float Inexperienced schedulers need to be very aware of not only the obvious critical path that appears as a prominent red bar on the chart, but also those ‘near critical’ activities lying just off the critical path that can easily become critical if delayed. A complex mine development and construction program will often have three or four different activity streams in this category. The answer here is to strive to preserve the maximum amount of float as possible on these activities in order to minimise the risk of project delays and disruption. Experienced schedulers also allow for the variable impact that different activities may have if they slip. For instance, level development usually only slips incrementally whereas vertical development, especially raise boring, can slip by 100 per cent or more quite easily.
SCHEDULING CONSTRAINTS The previous section broadly discussed the significance of contingency considerations. This section looks at why that contingency is necessary. In other words, it looks at those things that can and will go wrong in the course of an underground project.
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Learning curve
Drainage and dewatering
No matter what the activity, allowance must be made for the ‘learning curve’ as the workforce executes repetitive tasks quicker as they gain experience. If experienced trainers and supervisors are training a completely inexperienced workforce, it will take between six and twelve months to reach maximum development rates. Lack of training resources and poor supervision will prolong this learning curve period.
Decline development in wet conditions can be far slower than expected, particularly if the decline pumping system lacks the capacity to quickly clear flooded faces after pumping stoppages. Mining contractors in Australia absorb the cost of pumping up to 5 L/s because they can manage this rate of inflow. However, once face water inflow rates exceed 5 L/s, water related face delays increase dramatically.
Commissioning new equipment
Climatic extremes
New equipment can take anywhere from a week to six months to bring up to full efficiency. New shaft mucking and drilling arrangements can take three to six months to ‘debug’. If the contractor or mining company does not commit the engineering and design resources necessary to overcome the inevitable problems in newly designed mining equipment, they may never reach the intended productivity rates upon which the project completion date was based.
Problems arising from climatic extremes can disrupt mine development. For instance, water inflows may increase significantly during the spring thaw on an arctic project. Alternatively, finding water is often a major challenge on an arid desert mine site. When scheduling mine development in such demanding environments, it is essential that the scheduler consult those with local experience.
Ventilation
Haulage
Inadequate mine ventilation frequently disrupts mine development programs. This subject is discussed at length in a later section. It is referred to here because it is the most common restraint on mine development. Ventilation difficulties may arise from any of the following:
• excessively long vent ducts; • complications in the vent circuit; • drives that are too small for the ventilation duct that is required; which leads to
• excessive leakage from damaged or torn ventilation ducting or through bulkheads;
• • • • •
excess airway resistance and pressures; contamination from reusing the air too often; hot ground water; dead spots in the vent circuit; or larger than anticipated diesel fleets when applying extra resources to a project to get it back on track if it has been delayed.
Lack of ventilation at critical times during mine development is one of the most common delays, particularly where large diesel powered mobile machinery is in use. The key to avoiding ventilation delays is to over design the ventilation system during the mine development phase. The author’s experience is that if this is done, there may be just enough ventilation to maintain development in critical areas when the inevitable problems arise.
Heat Hot working conditions will severely affect any workforce. Productivity drops off noticeably when the wet bulb temperature exceeds 30°C. Work normally ceases in areas when wet bulb temperatures exceed 32°C because of increasing heat stress problems. Maintaining high air velocity will potentially limit temperature rises down to about 600 to 700 m below surface. Below that depth, adiabatic compression becomes the main factor and there is no option but to refrigerate the air to contain working temperatures. Many of the mines across northern Australia have been severely disrupted in some years by hot and humid conditions in the ‘wet season’. These conditions mainly occur in January and February. Development faces in some of these mines have reached rock temperatures in excess of 50°C.
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Trucks are used for all decline development in Australia where most mining contractors and companies are using 40 to 50 t trucks for mine development. Two trucks normally provide sufficient trucking capacity for a single entry decline. However, once level development commences, extra trucks may be required. More than a few mine developments have experienced significant delays at times when their ability to break rock far exceeded their capacity to remove it or their maintenance was sub standard.
Camp accommodation Most of the new mines developed in Australia in the last decade have been in remote areas of Queensland and Western Australia. Consequently, these projects have had to provide camps to accommodate their construction workforces. More than a few of them have underestimated their peak accommodation requirements and as a result, have suffered construction delays because of labour shortages at critical times.
Surface logistics Following on from accommodation planning on remote sites, supply interruptions to remote sites can severely impact on construction schedules. There are at least two mine sites in northern Australia whose road access is regularly cut-off for one to three months during the wet season. Spare parts can be flown into remote sites in an emergency; but major construction plant items cannot be airlifted unless a heavy lift aircraft is available and the runway is large enough. On very remote sites, delays often occur when even normal express delivery from the nearest capital city takes two to three days. For projects in developing countries with limited infrastructure, logistics planning is critically important and is an area where inexperienced contractors often learn painfully expensive lessons.
Clearing customs in other countries In addition to the delays that can occur due to limited infrastructure in some countries, schedulers must allow for the time required to clear customs when bringing plant and materials into such countries.
Decline and shaft congestion If getting materials to site can be a challenge, getting those
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materials into the mine can be even more so when underground construction activities are peaking. Underground level development often peaks at the same time as the construction of ore handling and other systems. At the peak of the Cannington development, five to seven haul trucks were running in the decline at the same time as seven concrete agitator trucks were hauling concrete and shotcrete to construction areas and major fabricated steel items were being dragged down the decline. Unless the underground logistics are carefully planned, significant delays will occur on major underground mine development projects.
In some off-shore environments, security issues may introduce potential delays, such as permitting for the transport, storage and use of explosives. Alternatively, site theft may be a consideration, as might be the risk of site invasion by ignorant itinerants or artisinal miners. In former war zones, the unexpected presence of ordnance can be a significant problem.
SCHEDULING TRAPS TO AVOID Too many activities in one place
Timely provision of plans Mining contractors often suffer delays in obtaining approved-for-construction mine development plans from their mining company clients. The delay in issuing the drawings may be due to lack of resources or indecision in the client’s planning team and must be taken into account.
Diamond drilling Underground diamond drilling is discussed here because it can play havoc with either mine development schedules or budgets or both. Most mine development projects for open stoping methods involve three stages of diamond drilling: Stage 1: widely spaced holes from the surface to define inferred mineral resources to justify the cost of underground development for the feasibility study and to locate the decline or shaft access to the orebody; Stage 2: closer spaced holes from underground drilling sites to define indicated and measured resources for the bankable feasibility study and to locate capital development on each level; and Stage 3: very closely spaced holes from drilling sites on or adjacent to each level to define ore outlines for stope planning. Diamond drilling in the first two stages usually does not pose problems. Scheduling problems often arise in the third stage when:
• firstly, the footwall drives (where one would prefer to drill these holes from) are never available because of ongoing development on the levels; and
• secondly, the timing of the stope definition drilling is critical if initial stope development is not to be delayed. There are two solutions to the problem:
• defer stope development and production start-up by six months to provide time for the stope definition drilling from the footwall drives and subsequent geological and stope planning (the six month delay usually does not appeal financially); or
• complete the definition drilling from drilling sites along the decline until scheduling permits the use of the footwall drives (thus incurring the cost of extra drilling and the excavation of drill sites along the decline). An entire paper could easily be written on diamond drilling alone. It is an area that has caused the author many headaches on past projects; hence the length of this discussion. The risk here is that schedulers will assume that they can some how fit stope definition diamond drilling in with development, or they overlook the drilling and/or its ventilation completely (surprising, but it has happened). The results are either seriously disrupted development schedules, blown budgets or a combination of both.
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Security Considerations
Scheduling too many activities in one working area inevitably leads to delays. Many problems stem from commencing construction activities before excavation blasting is completed in the area. Where there is a sequence of rock excavation, concrete pouring and structural steel installation, possibly followed by mechanical, electrical and controls installations, it is important to separate the activities as much as possible or make realistic allowances for the delays.
Complexity Older readers will be familiar with the KISS principal of ‘Keep It Simple Stupid’. This adage was never more apt than when applied to the underground environment. When young engineers are heard complaining that the construction manager and foremen are ignoring their development or construction schedules, it is quite likely that the schedule in question is complicated and difficult to implement. The author has learnt from bitter experience over the years that complicated schedules almost always unravel with significant delays.
Too many options Another rule of scheduling mine development is that ‘the rate of development of any face is inversely proportional to its priority’. Schedulers must not assume that they can maintain development rates in a decline, particularly if it is hot and wet, and carry out level development with the same crew. The crew will invariably favour the more comfortable level development and ignore the more important but uncomfortable decline development.
Too many priorities A mining schedule must have only one number one priority. Young schedulers have been known to produce schedules with six equal number one priority activities. As with too many options, a schedule with too many priorities will be ignored by the supervisors. Because of the complexities of working underground, particularly with fleets of trackless diesel equipment, supervisors need their work organised into as few priorities as possible. If a schedule is not simple and focussed on one priority at a time, it is of little use to the supervisors managing the underground work.
SCHEDULING RATES This section outlines the typical advance rates used by the author when evaluating or preparing mine development schedules. They are based on rates achieved either on projects managed by the author or observed on other projects. The rates quoted in this section apply to continuous operations working 2 × 12-hour shifts, seven days per week, as is now the norm in Australia.
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about 2500 tonne kilometres per day. The new Atlas Copco MT5000 truck, when it goes, is achieving twice this at two mines in Western Australia.
Mine development excavation rates Decline development Most major mine developments in Australia will involve at least one to two kilometres of single entry decline development to reach the first level at the top of the orebody. The actual advance rates achieved will depend upon:
• • • • • •
quality of the workforce, supervision and management; availability of the mobile and fixed plant;
This section considers the drilling rates that form an important part of many major mine development schedules.
Raise boring Raise boring is a critical part of most mine development schedules. It is normally used to excavate ventilation airways, escapeways and ore passes. Raise boring reaming rates depend on the:
geotechnical conditions; groundwater inflows; ventilating air temperatures; and spacing and depth of the stockpiles and truck loading bays along the decline.
The author has observed all these factors severely impact on decline development at one time or another. The best decline development rates will exceed 70 m/wk when all the factors are favourable. However most declines, with reasonable ground conditions and less than 5 L/s of water inflow at the face, will average 40 m/wk actual decline face advance. This figure includes allowance for 17 m deep stockpiles at 120 m spacing along the decline.
Level development The factors listed above all impact to some extent on level development. However, the greatest factor is the number of independent faces available for each jumbo to work at any time. Australian experience is that with at least three faces available in reasonable conditions, a jumbo (drilling 3.2 m long blast holes) will average 75 m/wk.
Shaft sinking Shaft sinking rates vary from country to country. Canadian blind sinking rates in large diameter deep shafts are averaging 30 to 35 m/wk. In comparison, Australian blind sinking rates are only 15 to 20 m/wk. However, Australian rates for strip and line sinking have averaged 35 to 40 m/wk on the last two major strip and lines at Mt Charlotte and Cannington.
Trucking The rate at which ore and waste rock can be transported out of the mine will limit overall development rates when the rate at which rock can be blasted exceeds that at which it can be removed from the mine. This observation borders on the trite, yet haulage restrictions have disrupted one major mine development that the author was involved in. Although not strictly an excavation rate, the trucking rate is an important component to consider when evaluating resources. If under resourced, trucking will significantly restrict development rates. It is important to consider all sources of broken rock to be transported when evaluating trucking capacity. Such sources may include:
• development ore and waste excavations; • major chamber stripping for crushers, pump stations and workshops;
• excavations for ore bins, dams and sumps; • strip and line shaft excavations; and • raise boring. In addition, the traffic density in the decline must be taken into account. When all things are considered, a 40 t truck will average
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• hardness of the rock to be reamed; • diameter of the drill string and reaming head used; and the • rig used. The size of raise bored ore passes and ventilation airways has steadily increased in Australia to the point where most internal bored raises are now in the range of 3.0 to 3.5 m diameter. Similarly, most ventilation raises bored from the surface are now in the range of 4.0 to 4.5 m diameter. Assuming that the appropriately sized drill string and raise boring rig are used, reaming rates for 200 MPa rock will average about 12 m/day. In quartzite that is harder than say 250 MPa, the reaming rate will drop to around 8 m/day. In most cases, pilot hole drilling rates are limited to twice the reaming rate to keep hole deviation under 1.5 per cent.
Long hole stope drilling Long hole stope drilling is included in this discussion because, although not strictly a development activity, it is a critical component in the sequence of activities needed to bring an underground mine into production. In the case of caving mines, long hole drilling is a critical part of the undercutting process. In the past, Australian mine owners have used contractors to complete the development excavations and then used their own workforce to undertake the stope drilling and subsequent production activities. The latest trend, as can be seen at Ridgeway in NSW, is to use the contractor to complete all the initial production activities as well as mine development. The contractor brings the mine up to full production prior to the mine owner taking over day to day operation. The advantage of this approach is that the mine owner is not faced with all the challenges of recruiting and training a new workforce at the same time as the challenges of starting a new mine are being dealt with. Typical long hole drilling rates for 76 mm diameter upholes (up to 25 m long) with T38 rods average around 2000 m/wk. This is based on a well maintained purpose built long hole drilling jumbo with a rod handling carousel.
Diamond drilling Diamond drilling is also an important part of the mine development process and is a critical step in the design phase needed to plan the initial stope development. Drilling rates vary enormously and depend on the ground conditions. In hard, abrasive and broken ground that is often associated with Ag-Pb-Zn orebodies, drillers will average around 250 to 300 m/wk when drilling BQ sized holes. In the more competent and easier drilling ground often found around gold orebodies, diamond drillers will average up to 500 m/wk.
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Mine development planning Mine planning is a critical and on-going step on the path to bringing a new mine into production. The sequence of activities involved includes diamond drilling, assaying, ore body modelling and stope development and drilling planning. An accurate orebody model is needed firstly; to plan the stope development, support and extraction and then secondly; to reconcile the actual produced tonnes and grade of ore against what was designed. The original orebody model is prepared during the feasibility study and is then updated from the stope definition drilling during mine development in order to plan the initial stope development and drilling. The author makes the following allowances for mine planning during the development phase:
• assaying after a drilling campaign is completed – one month; • orebody model updating and review – four months; and • final stope development planning – one month. VENTILATION Ventilation was mentioned earlier in the section dealing with scheduling constraints. It is discussed further in this section as it has caused the author more problems than any other factor when developing major new mines. The problems usually only occur in the lower areas of a deeper mine development where there is significant activity on the levels above. The problems are discussed in more detail here so that others may be forewarned and appreciate the need for rigorous ventilation planning in order to avoid these all to common troubles that beset too many new mine development projects.
Ventilation dead spots in the decline Because of the particular arrangement and connections between the decline and haulage shaft that were providing fresh air sources and the two exhausting ventilation raises, a ‘dead spot’ occurred in the decline at Scuddles during the mine development. A dead spot is a length of underground airway where there is no airflow. Unfortunately, this dead spot coincided with the access point where development was to commence into the main 325 production level at the time (with the surface being the datum 0 m level and the 325 level being 325 m below surface). Ventilation conditions were very poor and important level development was disrupted until sufficient development was completed and the exhaust airway raises deepened to eliminate the dead spot.
Excessive airway resistances Another problem encountered during the Scuddles mine development stemmed from the 2.4 m diameter raise bores being used for return airways. By the time these 2.4 m diameter raises had been extended to the 550 m level, the airway resistance in the raises exceeded the capacity of the surface fans and booster fan on the 325 level. Increased development was delayed on the level until another raise could be extended to the level and equipped with a booster fan in order to draw sufficient airflow through the level.
Ventilation modelling and planning The key to avoiding these ventilation problems during mine development lies in detailed ventilation planning. This involves laying out every step of the mine development sequence and then ensuring that the proposed layout can be ventilated. If ventilation planning is not carried out meticulously, the types of problems described above may disrupt critical stages of main development.
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When Scuddles was developed at the start of the 1990s, PC based ventilation software was cumbersome and under utilised until ventilation difficulties beset the project. The advent of easily used and powerful Windows based ventilation modelling software such as Ventsim changed all that. At Cannington in 1996, nine discrete steps in the mine layout were modelled to ensure that ventilation could be maintained during the development, especially with the main fans starting up half way through the development. Much to the relief of the author, the main fans at Cannington operated at their predicted duty point when started up for the first time. As an aside, at one stage during the Cannington development, prior to connecting the down casting haulage shaft to the decline at 450 m below surface, air velocities in the decline were up to 10 m/s without causing undue problems. This velocity dropped to about 5 m/s after the shaft holed into the decline.
Calibrating the ventilation model The initial ventilation modelling should be based on industry averages for various airway resistances. It is important to measure the actual airway resistances as the mine development proceeds so that the ventilation model can be calibrated to reflect the actual rather than theoretical resistances.
Adapting mine plans to ventilation needs Just as constructability is such an important consideration in construction work, ventilation is equally important in mine development work. As already discussed, many different ventilation problems can disrupt a mine development schedule. In addition, failure to consider practical ventilation issues when planning and scheduling level development may severely impact on development rates. In particular, the following issues must be considered:
• fresh air sources governing where auxiliary fans will be located;
• • • •
length of ventilation ducting from each fan; number of bifurcations in the ventilation ducting; number of times the air will be reused; heat and humidity picked up by the air before reaching the working faces; and
• ventilation duct condition after it has been in service for some time. These issues appear to be self evident, but too many mine planners and schedulers ignore them at their peril.
OTHER VITAL USES OF MINE SCHEDULES Mine development and construction schedules have two main purposes: firstly to underpin the contracts and contractors building the mine; and secondly, to provide a project control reference document that can be used to monitor progress.
Scheduling for contracts A complete and coherent mine development and construction schedule will underpin and enhance any major mining contract. In order to prepare a thorough and detailed schedule, the planner must first consider all the issues previously identified in this paper and resolve them in a logical manner. The first benefit arising from a sound schedule is that it minimises the risks that contractors have to price when bidding for major mining work. When the risks due to poor planning and scheduling are minimised and the contractors tendering for the
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project are able to adequately assess the geotechnical and other risks, their prices will often be very competitive with little separating them. On the other hand, shoddy schedules will concern contractors and lead them to load their rates to cover ill defined risks and unrealistic expectations by the project owner. As a result, prices may well vary by over 100 per cent as opposed to say 20 per cent when the risks are well defined.
RECOMMENDATIONS The following recommendations are offered as ways to improve scheduling methods and outcomes in the mining industry:
• thoroughly define the project scope as the first step; • allow appropriate contingencies at every stage of the schedule;
• use conservative industry average rates when compiling mine
Project controls
development schedules;
The second benefit arising from a sound schedule is that it provides the most important project control tool of all. In any project, and particularly in a complex underground mining project, time and control of that time is critical. The majority of the costs involved are time dependent. Therefore, if the project manager can control the time, the costs will follow. The key to using a schedule as an effective project control tool is to keep it up-to-date and to reassess it as required to address the changing requirements of the underground situation. In a dynamic and high pressure environment such as a longwall move in a coal mine, this may entail updating the schedule every day. At Cannington, where the requirements were not quite so intense, the schedule was updated weekly and reviewed monthly.
• use a quantity capable CPM scheduling package such as
CONCLUSIONS
The author wishes to thank fellow mining engineers John Dunlop, Alan Dickson, George Wakenshaw, Doug Syme, Heath Sandercock, Ian MacLeod-Carey and Tony Lennox for reading this paper and offering many useful insights and comments.
Scheduling is one of the most important project management tools. This is because; firstly, a sound schedule assists in obtaining the most competitive prices from contractors to build the mine. Secondly, it is a tool to assist the project manager in completing a project on time and budget. Whether it is mine development, construction or both, if the project manager can control time, more often than not, the dollars, pesos, rands or whatever will follow. Scheduling is too important to be left to junior engineers without significant input from above. Unless the senior project engineers and the project manager dedicate the necessary time to reviewing and guiding the planning and scheduling process, the resulting schedules may well be less than adequate. Poorly prepared and managed schedules are often at the root of project delays and cost overruns. Remember the cliché ‘If you fail to plan, you plan to fail’.
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MineWorks Planner in conjunction with plans and coloured pencils to assess all activity interactions and conflicts and resolve all the project execution problems;
• keep mine development schedules simple so that supervisors can understand and follow them;
• thoroughly evaluate all ventilation options and plan the ventilation for every step of the project; and most importantly
• senior engineers and the project manager must maintain a continuing interest in the planning and scheduling process.
ACKNOWLEDGEMENTS
REFERENCES Bertram, R (Raise Bore Australia) 2000. Personal communication. Kellerman, W and Clark, E, 1993. Underground Hard Rock Mining Capital in Cost Estimation Handbook for the Australian Mining Industry (The Australasian Institute of Mining and Metallurgy: Melbourne). Luxford, J, 1997. Surface to Underground – Making the Transition, in Proceedings Mindev 97 — The International Conference on Mine Project Development (Ed: E Barnes) pp 79-88 (The Australasian Institute of Mining and Metallurgy: Melbourne). Stacey, G S and Pak, D L, 1980. Some applications of the computer in project management in the mineral industry, in Proceedings Management in the Mining Industry, pp 69-79 (The Australasian Institute of Mining and Metallurgy: Melbourne). Thompson, A, (BHP Minerals – Cannington) 2000. Personal communication.
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Optimising Mobile Equipment Resources in Massive Mining T Puhakka1 ABSTRACT Mobile mining equipment presents only a small portion (ten per cent mining project costs but the selected technology will determine to a large extent the mining performance and mine profitability. The actual performance of the selected mobile equipment technology will rely on the support functions, infrastructure, user interface and user acceptance. Optimising the technology to match the circumstances is an increasingly important phase in a mine feasibility study. Smartly done feasibility studies include the possible future needs of changing or building-up of the selected technology. In massive mining applications the machinery faces quite different challenges than in selective or small-scale mining applications. In massive mining applications the efficient and continuous move of large amounts of material has led to the use of increasing degree of automation. This higher degree of automation in a mining application, on the other hand, will be best applied to a smooth and streamlined mining process and operation. Mining methods like block caving support automated operation. Automated mobile equipment needs supportive infrastructure, monitoring and information transfer systems as well as well established working procedures. Too often the user interface and human factor is disregarded during the consideration of automation. Through the available supplier programming tools it is possible to see how much the selected technology provides flexibility in changing production or economical situations, when major rebuilds or replacements are necessary, and view the effects of all these on the mine economy. When adequate technology is available from various sources, the difference becomes from preparing for the future, managing costs and combining the best layout with the most suitable technological solutions.
• physical (drift size, curves, inclinations); • environmental (humidity, temperature, water quality, rock mechanics, geology);
• ergonomic (machine layout, noise, vibration); • safety (fire protection, covers, breaks, system shutdowns); • legal (international, regional or mine based legal notes in any subject);
• processes (main processes and sub-processes); • operation targets (schedules, unit and fleet capacities, operation cost, investment);
• people (level of skills, compensation, training, hierarchy, decision making);
• way of working (monitoring, support systems and logistics, measuring, reporting);
• communication; and • infrastructure. These can be combined in a few main categories such as: meeting process requirements, meeting operational requirements and meeting unit and fleet functional requirements (Figure 1). The approach is the same for all technology levels from manual to automated. Technology contribution to total process
INTRODUCTION TECHNOLOGY SUITABILITY TO TOTAL PROCESS
PROCESS REQUIREMENTS
OPERATIONAL REQUIREMENTS
FUNCTIONAL REQUIREMENTS
PROCESS CONTRIBUTION LEVEL
Different operation targets and mine layouts support different technology solutions. Massive mine layouts like block and sublevel caving set different demands for applied mobile mining technology than narrow mining operations. Existing mines are bound by their infrastructure during expansions and renovations, whereas the new projects can review the available technology much more openly. In a change the resistance to change increases exponentially with time, one more reason to overview carefully the technological alternatives at an early stage. Smartly done feasibility studies include the possible future needs of changing or building-up of the selected technology.
FLEET APPLICABILITY TO LONG-TERM OPERATIONAL CONDITIONS AND NEEDS
EQUIPMENT ABILITY TO PERFORM SPECIFIED TASKS PHYSICAL DEMANDS
OPTIMISING TECHNOLOGY In a typical mine several levels of technologies are applied at the same time. The importance of thorough technology analyse in a mine project at an early stage is better acknowledged today. The mobile machinery may only be a fraction of the total mine investment, but it will largely determine the effectiveness and economy of the operation. The higher the technology applied the more important it is also to plan for its support and long-term economy in advance. Mine layouts are determined by geology, rock mechanics, and market situation and production requirements. The mine layout, technology applied and production form the optimisation playground. The following factors in the mine impact the selection of the underground mobile technology: 1.
Sandvik Tamrock Corp, PO Box 100, FIN-33311 Tampere, Finland. E-mail: [email protected]
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FIG 1 - Optimising technology to meet process requirements.
Identifying processes Process thinking has today reached also the underground development and production of a mine. There is a desire to integrate the underground production tighter with mineral processing. When this becomes a full chain, an underground rock factory has been created. The identification of all underground mine processes (eg development), subprocesses (eg development loading) and support processes (eg ventilation or service material flow) is the basis for applying optimised technology and support functions for them. This provides information to critical path thinking and simulation models. The identification of various mine processes will promote cross-functional thinking and lead to optimised technology solution and minimised technology risk. Most often there is a
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direct performance improvement to be gained directly through existing process improvement. The traditional approach optimises the unit performance and physical size in relation to the given mine layout. The unit performance gives limited information for the fleet performance without understanding of how the total process works.
Meeting fleet requirements The mobile equipment fleet and suitable technology, for example in a development stage, needs to match the production needs. Even if the rock support capacity or loading capacity is sufficient for one or two face attack, it may become the critical link in multiple face development like block cave development typically is. One individual unit utilisation may rise to critically high value (close 100 per cent) in a multiple face excavation case even if the fleet utilisation stays as a considerably lower figure (70 per cent). This can be an indication of missing resources, but also, bad machine-match, wrong development cycling or fleet distribution. The way of working, accesses between levels, daily shift and service arrangements for example affect the fleet utilisation. All resources should be well utilised in an optimised process with built-in flexibility for unexpected disturbances. Looking at the full process makes it possible to see the size and seriousness of changing parameters in relation to development time. Noticing this only at the actual development stage leads typically to throwing in more resources without the necessary cause and effect thinking; what are the consequences in other parts of the fleet performance and support functions. Delays in development
stage can seriously affect timing in production phase and the mine rate of return. Modern technology review shows the level of flexibility and how the fleet performance and cost will develop in time.
Meeting layout and infrastructure needs In less than 30 years, underground loader development has evolved from the first simple LHD models to integrated, automated loading systems. Performance, reliability and safety have been the three main development focuses in the LHD-development. During the 1990s, the 25-ton carrying capacity electric LHD was developed for massive mining purposes. Each generation reflects the equal mining method and mine process development and operational requirements. Applying a new technology to a mine process affects the design, infrastructure, production and way of working in many ways. Finding the best matching layout requires discussions between the mine designers and machine supplier. For example, in block caves the angles and lengths of the draw point depend on the caving layout, rock mechanical factors and draw configuration. Draw points also have the tendency to decrease and shorten in time due to draw point maintenance. For best material moving performance, the optimal herring bone angle and length needs to be reviewed also from the loading machinery’s point of view. Energy and underground ventilation requirements have re-boosted the use of electric loaders underground. When electric machinery is considered, the full loading process needs to be thought through for the largest
FIG 2 - Example of three-face development resource use with Sandvik Tamrock study program.
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FIG 3 - Development of LHD-loading machinery.
FIG 5 - Use of simulation as a technology and operation optimisation tool. FIG 4 - Meeting operational requirements.
Benchmarking possible machine coverage and to avoid for example the machine electric cables crossing. Plans need to be in place for maintenance and service bays. Full benefits with the lowest operational costs and highest production can only be achieved by reviewing technological alternatives well already in the feasibility stage. The use of simulation models has made it possible to test the suitability of the various technological alternatives in given mining conditions. Simulation is also the only way in identifying possible congestion in complex development or production situations. The more detailed simulation models require more information on equipment operation and long-term cost behaviour. Available study and simulation tools show how much the selected technology provides flexibility in changing production or economical situations, when the major rebuilds or replacements are necessary, and view the effects of all these on the mine economy. When adequate technology is available from various sources, the difference becomes from preparing for the future, managing costs and combining the best layout with the most suitable technological solution.
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Benchmarking of planned performance factors and layout to existing mines is a part of the feasibility stage. Both mines and suppliers have recorded information on machinery performance and costing in corresponding operations. They should never, though, be viewed without their operational environment information. Benchmarking best practises is a good way of determining the best use and utilisation of technology and people skills. Some of the values to compare are general machine performance, unit and fleet availability, utilisation, capacities, cost and reliability. Beyond lay the issues under cover like inventory volumes and costs, response times, turn- over, human competence, compensation, working arrangements.
Automation Automation has become an attractive technological choice in massive mining applications. Mining methods like block caving, where large amount of material is moved during steady state production support well automated operation. The higher degree of automation in a mining application on the other hand will be
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best applied to a smooth and streamlined mining process and operation. The challenge is then to master the effect of process disturbances or discontinuities like secondary breaking and build that into an automated model.
Automation is a higher form of technology, but it too needs supportive functions, service organisation, parts, diagnostics, logistics and a good supportive infrastructure. Automated systems build from ‘directed products’ that are integrated into one functional entity. From the owner’s point of view the operational reliability of this system becomes a different challenge since the boundaries between various suppliers are not straight lines. The system integrator, whether the mine or supplier, becomes the system supporter responsible for the functionality and system operation. Eventually also the automated systems will grow old and need upgrading; new software and hardware. If the mine lifetime is 20 years, several upgrades will be made to the machinery and automation system, some of which will also need infrastructure changes.
LIFETIME OPTIMISATION
FIG 6 - Infra-free system used in position measurement in an automated LHD-system. View through an automated LHD front camera.
The obvious expected result of an automated system is the increase in utilisation and decrease in operative unit cost. To obtain this task, several optimisation steps must be followed. The operational alternatives both for manual and automated solutions must be visualised and analysed. This helps to identify the benefits, challenges and possible risks. The automated system must be functional in the given mining process and the mine layout must be adapted to automated operation. There has to be a clear understanding of what type and level of decisions and how frequently need to be made at any given stage of the process. Only when the full system with machinery, process control, communication and operator interface has been identified, the final economical justification can be made. The automation solution must have an organisation acceptance and management commitment behind it. Only too often the new aspects of way of working and performance measuring are forgotten. The operator’s role changes to a decision-maker demanding training and different motivational targets. Elements like amount of light, colour and changing images in a monitor have in experiments been found important for concentration. Communication in the control room or at the operator’s desk does not become less but more important.
The initial investment cost is about one-quarter of the drilling machine lifetime cost. Drilling consumables and parts and service combined both are about one quarter each of the unit lifetime cost. The actual numbers depend on machine manufacturer, operational conditions and machine type and technology level. The rest of the costs are energy, water and operator. In loading the cost distribution is similar so that the initial investment presents some 30 per cent and operating costs some 70 per cent of the total costs. Out of the operating costs the fuel, tyres and bucket form most of the operating cost. Interestingly enough, if the operation on an LHD is reviewed, the bucket filling time describes a large amount of the operating costs (affecting bucket, tyres, fuel, operator cost service and maintenance). How the unit performs is a combination of design and manufacturing, optimised mine layout and operator know-how. Those machines having long bucket filling time will also have increased operating costs. In optimisation configuration a high-capacity unit typically has higher initial investment cost, but better reliability expectancy and thus longer lifetime and lower operational cost. The unit cost per hour is only interesting in relation to its performance and thus the over all high capacity unit cost per loaded ton will be lower than that of the low investment and low capacity unit. To maintain lifetime profitability, the mine needs to view its technological competitiveness. These can be built in also in early stages through various service contracts, performance contracts and technological enhancement programs. They commit to a regular continuous improvement procedure where the best practises, performance, technology and operator knowledge are reviewed and brought up-to-date. Mines are moving towards real time process control and resource monitoring systems. They combine the use of simulation tools for production forecasting and evaluating changing situations. This in turn means that processes are well defined and the resources carefully thought of. The resources will be increasingly more dedicated in autonomous system where moving a unit from one area to the other may not be possible at all. Using optimised number of resources and advanced working practices a building up a ‘transparent’ mine is possible. The information provided through individual unit monitoring will then be combined to process information in a larger scale. The chain effect of optimised technology reaches from mining project profitability evaluation to operation and from initial investment to mine logistics. Optimised machinery provides flexibility and increases competitiveness in changing market situations and has a high user acceptance.
FUTURE FIG 7 - LHD-operator in a control room.
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Massive mining methods like block caving and sublevel caving provide an excellent base for long-time optimisation where the technology optimisation is a part of the equation. It may be
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difficult to optimise a full mine for its lifetime due to changing internal and external conditions. When the number of variables in the lifetime optimisation becomes so massive that the forecasting reliability decreases below acceptable, the real-time resource and operation control becomes the key in maintaining the desired profitability. Here the individual decisions (related to machines or people) have an increasing effect. Increasing desire for lower cost and more massive operations moving to deeper deposits and decreasing human involvement will eventually lead to developing mining methods and processes
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more suitable for automation. Need of rock support may decrease and the importance of time and dimensions in development and excavation will increase. The future layout can be better fitted to suit the automated mining needs. The new mining machinery can look quite different with no cabin and no drivers seat, but will be more reliable and transmit more information than ever. The mining layouts and mobile mining technology will evolve tightly together for better over all mine rate or return.
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Intelligent Mine Technology Program and its Implementation J Pukkila1 and P Särkkä1 ABSTRACT Mining is commonly seen as a typical basic industry with rough and even dangerous working conditions, heavy environmental load, and a low level of high-technology and automation with much manual work and operations. A mining operation (open pit or underground) is a very complex one consisting of many manual, physical, mechanical and logistical operations with different human interfaces and decisions. For this reason it is a demanding and potential area for all applications of automation and information technology: controlling difficult non-linear, time varying multivariable processes and machines; solving of ergonomic safety and environmental problems; automation of logistical systems; applying robotics, mechatronics, and information handling, etc. In addition to that, big economic sums are involved with these operations. On the other hand, automation technology has developed towards intelligent and adaptive systems including some from the above mentioned fields of automation. Connecting them with modern information and communication technologies makes the idea of an intelligent and unmanned mine more and more feasible in the near future. This requires intensive co-operation and co-ordination between the mining industry and manufacturers of mining machinery. This has been successfully accomplished in the five-year Intelligent Mine research and development program as well as the three years of subsequent implementation of its results in Finland. The basic elements of an Intelligent Mine concept are: • mine-wide information and data acquisition system; • high-speed two-directional mine-wide communication network for real-time monitoring and control; • computerised information management, mine planning, control and maintenance systems; • automated and tele-operated machinery and equipment which are connected to the information system; and • links to the public information networks. The degree of automation of an intelligent mine depends on many technical and economical factors. The basic precondition for such an approach is that it will improve the total economy of the mine. This paper will describe the Intelligent Mine Technology Program with the results. Also the implementation which took place during starting of a new underground mine (Outokumpu Chrome Oy Kemi Mine) will be described.
INTRODUCTION The productivity of mines can be significantly improved and the costs of the final product decreased by applying new mining methods, advanced mining technology, automation and good organisational quality (Pukkila and Matikainen, 1994). The new mining methods in most cases are directed towards selective mining. This means minimising of waste rock dilution. This is done by controlled mining within the planned ore boundaries, and delineation of the orebody during production. The excavation methods in hard rock are still today based on blasting but in the near future non-explosive methods will be applied. Today the emphasis is given to automated or tele-operated drilling, charging, loading and transporting. The mining cycle will be made process-like as much as possible and with the economical coming of continuous miners to hard rock mining, this can be reasonably easily implemented (Elbrond, Matikainen and Pathak, 1993).
The automation process in mining started about ten years ago with automation of hoisting, pumping, ventilation, and other minor activities together with the process automation of the concentrators. Today emphasis is given to automation of production machinery and mine-wide information and communication including maintenance and production control systems. We could call this an era of high technology transfer to open pit and underground mining. The objectives of the Intelligent Mine and Intelligent Mine Implementation Technology Programs were to increase the productivity and profitability as well as improve the working conditions of the miners of Finnish mines by developing working methods, equipment, communication and data transfer systems, and automation. The future vision has been Intelligent Mine, which is a mining process controlled and managed in real-time most economically according to internal and outside conditions. Machinery is autonomous or remote controlled. All machines and activities are integrated by bi-directional high-speed mine-wide communication and data networks to enable the real-time communication, monitoring and equipment control. Mine planning, production planning and equipment maintenance planning systems are integrated by this network allowing all control and decision-making of the mine to be centralised above ground. The mine could be managed and controlled from anywhere through extended network or by satellite.
INTELLIGENT MINE TECHNOLOGY PROGRAM In September 1992, the Mine Automation Group of Finland, comprising members of the mining industry, mining machinery manufacturers, the Laboratory of Rock Engineering of the Helsinki University of Technology, and the Technology Development Centre of Finland (TEKES), started a five-year technology development program named ‘Intelligent Mine’ (Seppänen and Pukkila, 1993). The purpose of this program was to introduce automation into the most critical areas of hard rock mining to increase productivity and lower mining costs over a relatively short time and at the same time improve the safety and working environment. The critical areas are in the real-time management of the mine linking it to the total process including the concentration and metallurgical factories, maintaining machinery at high utility levels and automation of hazardous activities. The development program was formed in close co-operation between the mining industry and mining machine manufacturers and the emphasis is on the requirements of the end user, the mines. The practical development work was done in a real mining environment with the assistance of the mine personnel. The total budget of approximately USD 15 million was provided equally by participating industry and the Technology Development Centre (TEKES). The program management was as follows:
• Outokumpu Oy, represented by Outokumpu Mining Services Oy and later by Outokumpu Metals and Resources Oy and presently Outokumpu Mining Oy, MTG;
• Tamrock Oy, presently Sandvik Tamrock Oy, represented by Tampere and Turku factories;
1.
Helsinki University of Technology, Laboratory of Rock Engineering, PO Box 6200, FIN-02015 Hut, Finland.
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• Orion Corporation, Normet, presently Normet Oy; • Lokomo Oy, later Nordberg-Lokomo Oy;
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• Helsinki University of Technology, Laboratory of Rock Engineering; and
• Technology Development Centre of Finland (TEKES). The program management has been kept at minimum. It included only the representatives of the firms and organisations that had responsibility of the projects in the program. This was proved to be one of the reasons for the successful realisation of the program and keeping it in the planned schedule. In addition to these, many research centres, consultants and component manufacturers were working as contractors in the R&D projects of the program. The existing new technology from other industries was applied and developed to serve the needs of mines. This five-year technology development program is a beginning for future development which will lead to a concept of the Intelligent Mine - an automated high-technology mine with automated processes and autonomous or tele-operated machinery which are controlled in real-time to provide the best possible economical production according to the internal and external conditions (Lappalainen and Pukkila, 1993). The basic elements of an Intelligent Mine are:
• mine-wide information and data acquisition systems; • a high-speed, two-directional, mine-wide communication and information systems network for real-time monitoring and control;
• computerised information management, mine planning, control and maintenance systems;
• autonomous and tele-operated machinery and equipment connected to the mine-wide communication networks; and
• communication and monitoring systems to other mines within a company, machine manufacturers and public networks.
All these elements could not be implemented in five years. Therefore the development was done gradually according to the priority requirements of the mines and step by step to avoid ‘bottlenecks’ in the total production process (Figure 1). The automation process continues in such a manner that each automated function can easily be integrated into the next in the chain, and into the whole mining system. The research and development work was divided into four main areas: 1.
the real-time management of resources and production;
2.
machine automation;
3.
automation of production maintenance; and
4.
safety, training and motivation.
methods
and
production
Each of these areas contained R&D projects which together formed the most important development needs of this particular area in mine automation; taken together, the areas formed the essential backbone for the further development of the intelligent mine (Figure 2). Under the real-time management of resources and production were projects which constituted information acquisition, communication and information transfer as well as the processing and utilisation of this information for management purposes. This area of development had the most important role in mine automation. It also contained the major part of the application of computers in mine-wide production planning and control (Figure 3). The backbone of the real-time management is the mine-wide information network. This is composed of a two-directional, high-speed main network into which local networks are connected (Figure 4).
FIG 1 - Development steps toward the Intelligent Mineä.
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FIG 2 - Intelligent Mine Technology Program: R&D projects.
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MINE MANAGEMENT
DEVELOPMENT CONTROL OPTIMISATION
STRATEGIC PLANNING
OPERATING CONTROL
ACTIONS
MAINTENANCE
DECISION-MAKING PROGRAMS
CONNECTIONS TO OTHER CONTROL SYSTEMS
ANALYSING
PUMPING
AUTONOMOUS MACHINES
VENTILATION
VEHICLE LOCALISATION
MAINTENANCE MANAGEMENT
PERSONNEL IDENTIFICATION
SAFETY / SECURITY
ORE HANDLING
NAVIGATION
SURVEYING
The machine automation development area contained R&D projects of the mine machinery manufacturers. These projects included automation of production machines in such a way that it resulted in increased machine-working hours, dependability and higher productivity. Included were fault diagnostics and data acquisition systems incorporated within the machines, as well as the automation of machine functions and operation. Navigation systems, modular type of production monitoring and fault diagnostics were developed in co-operation with mines, research centres and other manufacturers within the program. The production and condition control, fault diagnostics and monitoring systems were tested in the mine-wide intelligent condition control system, developed within the Intelligent Mine Technology Program by Outokumpu Mining Services Oy, at Pyhäsalmi Mine (Figure 5).
PRODUCTION PLANNING RESOURCES GEOLOGY
WORK ORDERS SPARE PARTS RESOURCES HISTORY
PRODUCTION CONTROL
MAINTENANCE CONTROL
LOCAL AREA NETWORK
GRADE CONTROL
REMOTE CONTROL FAULT ALARM OPERATION
DAILY SERVICE FAULT ALARM OPERATION
FIG 3 - Real-time management of resources and production (Courtesy of Outokumpu Mining).
To Public ATM-Networks Video Conference
AREA MAINTENANCE
PRODUCTION REPORTS SERVICE REQUESTS STATUS INFORMATION
OPERATOR
MAINTENANCE INSTRUCTIONS SPARE PARTS FAULT DIAGNOSTIC
Concentrator Plant Offices Mine Control Control Room Room PRECALCULATION DATA SAMPLING
SDH / SONET
ATM-switch
Crusher Station
Duplicated Fault-Tolerant Folded Bus
Mine
ATM - WAN
PLC Optical Fiber Pair
ATM-Switch
FIG 5 - Intelligent production and condition control system (Courtesy of Outokumpu Mining).
155 Mbit/s
Colour Videos Control Room
Digital Radiotelephone Base Station
U.G. Office and Workshop Colour Videos
FIG 4 - Intelligent Mine integrated service communication network (IMIS-Net) (Courtesy of Outokumpu Mining).
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The development of drilling equipment was done by Tamrock Oy Central to the design of Tamrock’s control system has been modularity, as was the case with Tamrock Loaders (Figure 6). The development of underground drilling has already gone through mechanisation and automation during the last ten years. Machines have become more reliable in operation and today automated long-hole drilling rigs are more or less standard equipment for modern hard rock mines. The main goal for automation in long-hole production drilling is to increase effective drilling time and thus achieve more drilling metres per shift. Automation also results in accurate control of hole length. The increase in shift drill metres can exceed 25 per cent through automation.
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FIG 6 - Modular automation system (Courtesy of Sandvik Tamrock).
As in underground drilling applications the significance of accurate drilling in quarrying operations is remarkable. Drilling results can be maximised through quality drilling by eliminating errors in collaring, aligning, deviation and hole depth. These errors influence the actual drilling pattern by increase or decrease of spacing and burden from the planned pattern. Field tests have proved that the correlation between drilling accuracy and drilling quality affects the costs of blasting remarkably. Tamrock’s contribution to the automation program included also the development of the systems for quality surface drilling, including localisation of drill holes, and navigation system for autonomous rigs for open pit mines (European Community’s ESPRIT II program). In the future the level of automation will increase and drilling equipment will become increasingly integrated into mine planning and control systems with an overall goal of a more optimised total mining process (Tuunanen, 1993). Features of automated drilling machines in future’s automated mine will include:
• • • • • •
FIG 7 - The effect of mine automation on work safety and motivation (Pukkila, 1999).
increased level of automation and autonomy, remote operation (tele-operation), data communication,
Results
maintenance support,
The five-year period to develop all systems needed for Intelligent Mine was not realistic. Therefore the approach was made in steps: first to investigate what was available and then develop the most urgently needed systems and machines and make them functional, keeping in mind the final vision possibly realisable after 10 to 50 years. Unfortunately, during the five years, there was not a single mine where all results of the projects could have been tested together. Therefore testing was done in several mines in Finland and abroad. In any case thanks to the good co-operation of the participants and the common coal, the program was a success and a good start for implementing the technology when such mine appears.
navigation, and task planning and reporting.
The automation of production methods and production maintenance covered a large area in the automation program. It constituted such fields as machine maintenance, new hoisting methods and mine profitability analysis. New machine constructions such as a mobile underground crusher, an automatic LHD machine, an automatic haulage truck, charging equipment and shotcreting equipment were also included in this area of development. Another important area of research and development in the mine automation program was human engineering (Pukkila, 1999). Automation will bring new type hazards of accidents and have an effect on work motivation. Although safety has been taken into account within each development project, the human role in the automated environment, especially in a mine, has not yet been extensively studied. In this part of the program, the methods and tools for introducing automation successfully into the mine environment, were being developed (Figure 7).
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INTELLIGENT MINE IMPLEMENTATION Intelligent Mine Implementation Technology Program (IMI) was a research and development program in which the major part of the machinery and systems developed in the Intelligent Mine Technology Program were being further developed and tested to the stage in which they can be implemented in the planned underground mine of Outokumpu Chrome Oy at Kemi.
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The results of this program were: readiness to realise the Intelligent Mine vision, have the machinery and systems developed ready for commercialisation, and show internationally the know-how of Finnish mining industry in concrete examples at the Kemi Mine. The technology program lasted three years (1997 - 1999) and its estimated budget was USD eight million. It contained 12 development projects, which were divided into three main areas: implementation of advanced technology, data utilisation, and training and adaptation of mine personnel to the new technology, systems and environment (Figure 8). In addition the program contains two supporting projects: co-ordination of the program and a project concerning with layout of advanced mining. The objectives of the program were set according to the needs and requirements of the Kemi mine. The contents comprises of the following:
• fast mine-wide communication and data networks (Figure 4); • systems integrated to the networks (production management
INTELLIGENT MINE TECHNOLOGY PROGRAM 1992 - 1997 Real-time control of production and resources
Automation of production and production maintenance
Machine automation
IMPLEMENTATION 1997 - 1999 TRAINING OF
AUTOMATION AND REMOTE OPERATION
PERSONNEL
PERSONNEL
MINE 3DPLANNING
system, Figure 9);
• machines and equipment integrated to the networks; and • mine infrastructure and personnel training (Figure 10).
REAL-TIME INFORMATION
TECHNOLOGY
INFORMATION UTILISATION
PRODUCTION MANAGEMENT
REAL-TIME INFORMATION TRANSFER
The program was realised by the management of the Intelligent Mine Technology Program in association with the management of the Kemi Mine. The program consisted of:
INTELLIGENT MINE ~2000
• further development and testing of the machinery, equipment and systems developed during the IM-program;
• developing and testing necessary additional systems and computer programs;
Mine planning
Production management
FIG 8 - Intelligent Mine Technology Program and its implementation.
Maintenance management Remote-control tele-operation
Production monitoring
monitoring and fault diagnostics
Smelter
Process automation
FIG 9 - Production management system of the Intelligent Mineä (Courtesy of Outokumpu Mining).
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FIG 10 - Training program for the personnel at the Intelligent Mineä.
• further developing and testing the compatibility of systems in order to form one integrated system suitable for the underground mine of Kemi; and
• planning and carrying out a training program for the operating personnel of the mine. During the implementation, the mine was made ready to apply the Intelligent Mine concept. The Kemi Mine acted as the testing ground for integration of all partial components into one functioning system and will be, in the future, the first mine to fully utilise the concept. The IMI-program supported the needs of the Kemi Mine, but at the same time provided readiness for commercialisation of developed machinery, equipment and systems individually or as total integrated Intelligent Mine concept.
Research and development projects All of the R&D projects were started in the beginning of 1997 and have been concluded at the end of 1999 (Table 1). During this implementation stage the new technology was tested. Possible unknown faults or new requirements in systems and applications were recognised and corrected. For the training and adaptation of mine’s personnel were given much attention. The changes in work and new environment will create need for preparation. This is extremely important in the early stage. The tasks, character of work, safety matters and
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organisation will change. This might cause opposition and if not properly dealt with, can make the implementation of new technology difficult. The training program (Figure 10) was developed during the IM-program and was partly realised during the implementation stage. This program can be used wherever the new technology will be implemented. The mine’s organisation has also been planned to suit the applied new technology and best utilisation of the systems created (Figure 11). All monitoring and control will be done from an operation centre build on the surface within a new office building. This operating centre will have the necessary computer systems for real-time control, simulation and visualisation of mining process, monitoring of the mine and concentration, and tele-operation of the automatic machines (Figure 12). The organisation structure of the mine as described cannot be very complex. The information is available to everybody and the decision-making of day-to-day production decisions is given to lower levels of organisation. The management of the mine will have more of a supporting and advising role. The lower management will deal more in the safety issues and be ready to give advice whenever necessary. The operators of the different sections work as teams and take responsibility of their own section. Different sections assist each other and sell their services when needed. For example, the maintenance teams give services to the production teams. This way the responsibility will be shared and can be rewarded to the teams as extra pay depending on the result measured by productivity increase or other means.
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TABLE 1 Intelligent Mine Implementation Technology Program: projects and implementation organisations. Research and development areas
Projects
Implementation
1. Projects supporting the program
1.1 1.2
Layout of advanced mine
HUT, Lab of Rock Engineering
2. Data utilisation
2.1
Real-time information networks
Outokumpu Mining Oy, MTG
2.2
Economical optimisation
Outokumpu Chrome Oy, Kemi Mine
2.3
Management systemsfor mining systems - 3D-mine planning - Maintenance management
Outokumpu Mining Oy, MTG
3.1
Remote controlled/ automated drifting process
Tamrock Oy
3.2
Remote controlled/ automated production drilling
Tamrock Oy
3.3
Unmanned loading
Tamrock Oy
(3.4
Automatic haulage)
3.5
Underground emulsion charging
Orion-Corporation Normet
3.6
Automatic shotcreting
Orion-Corporation Normet
3. Application of new technology
4. Personnel training
Co-ordination
HUT, Lab of Rock Engineering
3.7
Intelligent crushing
Nordberg-Lokomo Oy
4.1
Personnel training for ‘Intelligent Mine’
HUT, Lab of Rock Engineering Outokumpu Chrome Oy, Kemi Mine
MANAGEMENT
RESEARCH GROUP
PLANNING GROUP
OPERATIONAL GROUP
SUPPORT FOR OPERATIONAL ORGANISATION
OPERATION CENTRE - Mining process control - Concentration process control - Maintenance and repair - Traffic and roads - Developments - Personnel - Machinery control and remote operation - Production infrastructure
Orders Production targets Communication
Development works
Stoping
Mucking and haulage
Crushing and hoisting
Milling
FIG 11 - Proposed plan of operational organisation for the Kemi Mine.
Results The final results of the implementation are not yet seen, but during 2000 many of the systems and machinery developed will be in practice. The Kemi Mine will most probably apply the technology that has been developed and has developed it further
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for its own use. Some problems are yet to be solved but as the mine development continues the solutions for these problems will be found. It is too early to say that all applications of the systems, machinery and organisational changes will be a success but so far there have been no impossible situations.
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FIG 12 - Proposed production visualisation and simulation systems.
CONCLUSION As a conclusion, it could be said that both of these Technology Development Programs have given much to the Finnish mining and mining machinery manufacturing industries and also to the Helsinki University of Technology. The mining industry has had a chance to develop its mine management systems, mining methods and methods of increasing productivity and safety. The mining machinery manufacturers have had a good opportunity to develop their machinery, automation and systems integrated to the machines and so to increase their competitiveness in the market. The university has had a good opportunity to increase and develop its know-how and so to be able to produce more confident new mining engineers.
REFERENCES Elbrond, J, Matikainen, R and Pathak, J, 1993. Robotisability of operations in hard rock mining, in Proceedings Second International Symposium on Mine Mechanization and Automation 1993, (Eds: G Almgren, U Kumar and N Vagenas) pp 411-414 (Balkema: Rotterdam).
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Lappalainen, P and Pukkila, J, 1993. The mine automation complex, in Proceedings APCOM XXIV International Symposium on the Application of Computers and Operations Research in the Mineral Industry, Montreal, Canada, 1993, (Eds: J Elbrond and X Tang) Volume 1, pp 134-144 (Ecole Polytechnique de Montreal: Montreal) Pukkila, J, 1999. Implementation of mine automation: The importance of work safety and motivation. Acta Polytechnica Scandinavica, Civil Engineering and Building Construction Series No 118, 153 p (The Finnish Academy of Technology: Espoo). Pukkila, J and Matikainen, R, 1994. Mine automation, the key to profitability, in Proceedings 16th World Mining Congress, Sofia, 1994, Volume 2, pp 258-267 (Bulgarian National Organizing Committee of WMC: Sofia). Seppänen, P and Pukkila, J, 1993. Intelligent mine technology program, in Proceedings Second International Symposium on Mine Mechanization and Automation 1993, (Eds: G Almgren, U Kumar and N Vagenas) pp 79-82 (Balkema: Rotterdam). Tuunanen, A, 1993. Automation of hard rock drilling machines, in Proceedings Second International Symposium on Mine Mechanization and Automation 1993, (Eds: G Almgren, U Kumar and N Vagenas) pp 177-184 (Balkema: Rotterdam).
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Use of OPTIMINE™ Simulation Tool for Mobile Fleet Selection T Puhakka1 and V Kainulainen1 ABSTRACT Sandvik Tamrock Corp has developed an underground rock flow simulation tool OPTIMINE™ which is a direct simulation tool for matching suitable technology in rock material flow for underground mine layouts and production requirements. The tool gives the possibility of determining intersections, limitations and critical points in the layout and working procedures at an early stage of a new mine/mining area planning process. The tool combines the supplier machine knowledge with operational know-how and mine environment. It can be used for identifying the effect of an individual incident such as the effect of one additional turn in a continuous loading cycle to the daily capacity. This paper discusses the way the OPTIMINE™ simulation tool operates, the benefits and restrictions, and reviews how it simulates some example cases and how the mines benefit from the tool in maximising the performance and optimising the mobile fleet, scheduling work and fine-tuning the layout.
INTRODUCTION The development of computer hardware and software technology during the 1990s has made the development of more demanding simulation programs possible. Simulation programs have become a widely used tool in massive mining operations to analyse the mining process and system functionality already in the feasibility stage. Large mine-wide simulation programs concentrate on describing the system and giving information on the progress of operations through the mine lifetime. The challenge is then to make a simulation model in which the smaller operational details can be identified, combined with a large system without sacrificing the speed of a run and system flexibility in mine lifetime simulations. The time to update for example layout changes can be fairly long. Process visualisation and animation are an important part of the simulation tools and often the mental ‘proof’ of process functionality. The latest development focuses on applications that can be used during mining for operational and process guiding purposes. Operationally more important detailed information on the layout, way of working and resources in use is then added to the simulation model. The development of the Sandvik Tamrock Optimine™ simulation tool has been developed to match suitable technology to the operational layout.
• operates on laptops, Windows 95/NT environment; • functional user interface; • use of accurate machine performance and behavioural information;
• identify effect of fairly small operational details (short and long runs);
• flexible reporting and data collecting and inputting systems; and
• has expansion capabilities for further simulation purposes. The simulation model was developed in co-operation with Cybercube, a Finnish company specialising in various types of simulation application tool development. One of the software components used in the OPTIMINE™ simulation is a traffic-controlling tool TRAM™, a program developed by Cybercube.
Building a mine simulation model The simulation model utilises a mine’s 3D model already planned or real process information, such as physical size and dimensions of the simulated production area. The mine model can be created from existing mine design models (Surpack, Datamine), but even hardcopy mine level layout can be used as a basis for the simulation modelling (example .dxf files). Another option is to create the model by using AutoCAD and to import the layout to OPTIMINE™. The actual simulation model is then created on top of this mine layout model. The user enters basic running parameters and variables into an input table. Typically, information on equipment, mine conditions and working arrangements, and for example limiting factors and prioritisation information for conflict solving can be added. The simulation can be started after the 3D information has been imported and all input data entered. Typically, the mine view (Figure 1) is for the whole simulated 3D model which the user then can zoom in/out and turn to view the model from any desired angle by using mouse controls.
OPTIMINE™ AS A MINE SIMULATION TOOL The focus on the OPTIMINE™ simulation tool development has been to combine the available technology with mine layout and operation for optimised end-result and to tie up the actual operational parameters of mining machinery into a mine operational simulation model. This first-stage simulation model looks into the mine rock material flow and identifies the effect of detailed layout arrangements on every day operational parameters and working efficiency. Important design factors:
• flexible and fast mine layout model building system for simulation base;
• easy and quick layout, machinery and other updates; • easy and flexible simulation model building capability; 1.
Sandvik Tamrock Corp, PO Box 100, FIN-33311 Tampere, Finland.
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FIG 1 – An example of a 3D model of an underground mine.
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OPTIMINE™ SIMULATION AS A FLEET OPTIMISATION TOOL A large block caving mine is a typical fleet optimisation simulation case. An operation with 100 or 200 drawpoints needs to have just the right number of units to have a continuous, trouble-free operation and to avoid traffic problems. A mine with various mining areas and/or truck hoisting is another typical example well suited for simulation. OPTIMINE™ gives a way of determining intersections, limitations and critical points in the layout and working procedures at an early stage of a mine planning process. The tool concentrates on describing and simulating, but it can be used as a basis for large-scale simulation programs.
Suitability for underground production optimisation The OPTIMINE™ simulation describes well and fast how the rock material flows for the given production layout. For example, a truck being loaded by an LHD can be looked at in detail and the user can see the machines moving and performing in the given environment. Traffic conflicts with identified reason codes can be analysed. The user can vary the input parameters to simulate various situations and determine the optimum solution. The capacities/machine(s), operating information (production, trammed distances, utilisation) of individual machines or the total fleet can be seen, for example, graphically. The material flow, like volume of ore in different locations (drawpoint, crusher, and storage silo) can be followed in the tables or from the graphs. With the OPTIMINE™ simulation, the real non-productive time and reasons for this can be easily identified. Through simulation, evaluation of the issues behind low fleet utilisation can often be identified as wrong resource selection and constraints in process establishment.
Examples of simulation The two technological scenarios examined have been comparing manually operated and automated loading and trucking systems. In large and/or complicated production systems it would be impossible to evaluate the changing situations in a reasonable time. The effects of small incidents or combined incidents would be impossible to forecast without a simulation tool. The picture describes a detail of a simulation where a more detailed routing is described for a loading machine during operation. Here the loading route has been divided into smaller sections for machine behaviour identification. Each section has its own parameters (acceleration, speed, etc) that can then be changed, if necessary. Although simulation does not solve any problem or optimise the system itself, it does give valuable information as how to proceed in solving the problem. The information output from run simulations is as accurate as the input, and hence the simulation should always be done in close co-operation with the mine and the supplier. This helps identifying the actual existing and desired processes in the mine and increases understanding of the simulation results. The mere operational data of machinery or mine is just part of the information. What the processes are and how they ‘flow’ give the operational rules to act in ‘what if’ situations. The simulation identifies behaviour of each LHD in a block cave simulation (Figure 3). LHD-unit speed curves give information on how the simulated machinery behave in the given conditions and help to find the best possible layout and unit routing. The related capacities of each unit can be viewed on line and recorded in out-put table (Figure 4). Occurred instances and user delays are available for inspection at any given time and give on-line view of how the simulated system works.
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FIG 2 - Example LHD hauling route to crusher or dumping point.
Optimine™ simulation reports how the individual LHD or truck total time is divided and what is the actual unit utilisation to its dedicated work (Figure 5).
Benefits of OPTIMINE™ simulations The simplicity of the tool makes layout changes quick and relatively easy to do compared to some other existing simulation programs. A new model can be added, and the existing models expanded fairly easily. Depending on the complexity of a model, making a new layout model takes approximately two to four days. The simulation is a virtual reality 3D-model built on top of the mine layout database. The equipment selection data can be acquired from the user maintained database. The simulation program works in Windows 95 environment and laptops. The program gives basic information on ‘what if’ situations and layout design in a small-scale and how they will affect the total material flow. Simulations can help to maximise an individual unit and fleet performance and identify critical paths during operations. It gives information as to how the process reacts to various disturbances and how the technology meets the process requirements. The fact that generating different scenarios is easy enables the user to make several runs and generate large amount of information for the decision making process. Simulation, due to the easy layout making capability, is a good tool for fine-tuning layouts. Thus, it can be used as a basis and an as expansion to larger scale mine simulation programs. The simulation tool can be used in existing mines for fleet renovations and expansion programs where the typical challenge is the limitation of the existing infrastructure to a new equipment fleet. OPTIMINE™ can be used to detect and forecast disturbances in the rock material flow, identify the effect of individual incidents on the total process and maximise the resource utilisation. The effect of drawpoint hang-ups on the machine forecasted service scheduling can be put into the model. By adding cost configuration to the model output reporting data, the economical effectiveness of the process can be reviewed as well.
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USE OF OPTIMINE™ SIMULATION TOOL FOR MOBILE FLEET SELECTION
FIG 3 - An example of a large-scale simulation model showing the layout, machinery and operational data during a simulation run.
Scenario
Oper. System
Dispat ching System
1 2 3
Manual Manual Manual
4 5 6 7 8 9
Auto Auto Auto Auto Auto Auto
No No Yes (Truck s) Yes Yes Yes Yes Yes Yes
ID
Number of LHD’s From From U/Cut draworepass point 2 1 1
8 8 8
6 6 6
Extract. Level Producti on [ton/day 12 192 12 904 13 528
1 1 1 1 1 1
7 6 5 5 5 4
7 7 6 5 4 4
19 714 17 710 14 984 14 530 13 359 11 418
Numbe r of Trucks
Production Criteria (Yes/No)
Production ton/day
Production Criteria (Yes/No)
Total Production
Yes Yes Yes
1 622 1 088 1 243
No No No
13 814 13 992 14 771
Producti on Criteria (Yes/ No) No No Yes
Yes Yes Yes Yes Yes No
2 617 2 608 2 801 2 807 2 548 2 585
No No No No No No
22 331 20 318 17 785 17 337 15 907 14 003
Yes Yes Yes Yes Yes No
FIG 4 - Example of some output data from simulation (not integrated into other figures).
FUTURE It has been said that an underground mine is mainly a material handling and scheduling exercise. What makes this material handling exercise interesting is the nature of various materials challenging the environment and infrastructure and changing conditional constraints underground. The new optimisation tool has been used to optimise rock material flow and machine fleet and help design and verify
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crushing and hoisting system capacities. The trend towards further mine automation demands first a well identified process that can be operated with confidence before the benefits of full automation can be added. Simulations provide information for feasibility studies that evaluate various operational and technological solutions and open a window for selecting the best options. Typical mining method applications today are block caves and massive sublevel stopes. The use of simulation tools is also typical for applications with ramp hoisting systems.
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Total LHD Fleet
Service 15 %
Tramming time 43 %
User delay 14 %
The future will undoubtedly bring in large room and pillar operations where specially worked scheduling and traffic arrangements play a big role. Simulations of rock material flow also give a review of the daily decision making process and needed level of material flow management based on selected technology. One year after the start-up of the OPTIMINE™ simulation tool development and testing, most of the above targets have been met and the tool has proven to be both flexible and user friendly and the models are fairly quick to build. The model boundaries have yet to be found and expansion possibilities as to development and scheduling features kept realistic.
Wait for dump truck time 1%
Congestion 15 % Tipping time 6%
Load time 6%
FIG 5 - Example of fleet/unit utilisation graphics as an output from a simulation (not integrated into other figures).
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Drilling and Blasting Designing and Delivering Explosives Systems and Solutions for the Underground Massive Mining Industry: The Next Five Years with Dyno Nobel
W R Adamson
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Digital Blasting — An Opportunity to Revolutionise Mass Underground Mining
D Kay
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Designing and Delivering Explosives Systems and Solutions for the Underground Massive Mining Industry: The Next Five Years with Dyno Nobel W R Adamson1 INTRODUCTION The extraction of world-class mineral deposits using massive underground mining techniques is not an entirely new theme to world mining companies and experts. Previous MassMin conferences have helped to gather and disseminate the valuable experience and knowledge so important for the proliferation of safe and efficient mining of large underground deposits. With the passing years since the first MassMin conference in Denver in 1981, a number of factors have contributed to an increase in massive underground mining around the world, including environmental pressures, gradual exhaustion of shallower deposits and the enabling effects of improving technology in disciplines related to underground mining. Massive underground mining is heading deeper, embracing challenges which include working with elevated temperatures, increasing in situ and mining induced stress levels and longer ore transport routes. Improvements in mineral processing technology will tend to lead operators to contemplate the extraction of lower grade orebodies in the event that the ore can be efficiently transported to the surface and hence the plant. No doubt this tendency, and its associated challenges is encouraging a rich growth in innovative engineering in such fields as mine planning and materials handling as well as geomechanical issues such as support design and the simulation of extraction sequence effects on stress re-distribution. MassMin 2000, we are sure, will showcase a wide variety of new contributions in thought and actions in these fields. There is a definite place for academic institutions and mining related suppliers in the process of developing new technologies and practices in support of present and future endeavour in this field. Dyno Nobel has its own part to play, certainly as an enthusiastic and proud major sponsor of MassMin 2000, but also by means of identifying common cause between our Vision and Values and the needs, present and future, of the underground massive mining industry. The critical success factors, which Dyno Nobel pursues are: 1.
health, safety and environment,
2.
customer relationships and market development,
3.
cost optimisation (while maintaining quality),
4.
technology and innovation,
5.
organisational development, and
6.
leveraging capabilities globally.
Through the present paper we would like to share some of these values and offer some practical insights as to how their realisation will contribute to the achievement of MassMin goals and needs.
1.
Senior Technical Consultant, DynoConsult, Dyno Nobel Asia Pacific Ltd, Perth WA 6000. E-mail: [email protected]
MassMin 2000
THE NEEDS OF MASSIVE MINING: HEALTH, SAFETY AND ENVIRONMENT The viability of massive underground mining depends on a number of factors and the satisfaction of some basic needs. An important need is that of producing ores and delivering them to the site of further processing at a cost which is competitive with alternative sources of production. Another such need is that of minimising the ‘non-valuable’ or waste component of broken rock which is extracted along with the ore. These needs will be addressed in due course, however the most important need that must be satisfied for massive mining to be considered successful is that it is carried out in a safe manner. A unique feature that sets the underground mining world apart from other parts of the industry is that we simultaneously extract the resource we produce while constructing the environment (the tunnels and galleries) in which we work. This means that our daily working environment may expose us to increased risk. What allows us to survive these risks and carry out a productive working methodology is a diligent and systematic attitude to Safety. Safety is the foundation upon which all other Dyno Nobel Values, the pillars of our success are based. Our approach to safety depends upon a combination of proven systems, recognised and accredited audit tools and the fostering of a safety culture. Embedded in this culture is a duty of care to ourselves and to our customers, a promise so to speak, to decline unsafe practices, products and systems. The Right of Refusal to Work protects all Dyno Nobel employees in the event of doubt regarding the existence of potential hazards to resources, human or otherwise. This protection is extended through our service to the industry in cooperation with, and for the benefit of, our customers. However, safety is also a matter of removing or reducing risk through control of the process. Two aspects of the explosive supply role are considered here; that of reducing human contact with the hazardous environment and that of reducing the hazard itself. the key to both is innovative technology and training.
THE NEEDS OF MASSIVE MINING: TECHNOLOGY AND INNOVATION Innovation is one of the pillars of our corporate strategy. Our record for innovation already includes the development of water based explosives (that started with watergels and later continued with emulsions), as well as Nonel® initiation systems, and the mining industry can be assured that we have now and will continue to maintain a strong focus on innovation. Within Dyno Nobel there is a multi-level approach to innovation involving each stage in our productive processes through to and including those points of contact between supplier and customer, between explosive and rock. While in a commercial sense Dyno Nobel is a vendor of explosives to the mining industry we are, at the most basic level a supplier of energy to rock masses.
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The focus of our innovative work and one of the keys to safety, is to be found in the enhanced control over the delivery of energy to the rock mass you wish to extract. We pursue this control through a number of means:
• control of the energy at its source through innovative products;
• control over the release of the energy through timing; • control over the placement of the energy through new innovative delivery systems; as well as taking a holistic approach to optimising the blast outcomes in such a way as to maximise the economic benefits to our customers and their shareholders. As a global company, and in accordance with the Critical Success Factor of Leveraging Capabilities Globally, Dyno Nobel dedicates significant resources to the facilitation and support of cross-functional international technology teams. Through these teams, we co-ordinate the skills and thinking of specialists from all parts of the globe in order to create vital synergies. Recognising that the creative process often begins with dreams, our teams are encouraged to think beyond our current boundaries in the same way that you - the massive underground mining pioneers - decide to challenge the apparent limits of depth, temperature and stress in extracting needed ores from ever more demanding underground environments. These teams unite chemists with systems engineers, marketing specialists with blasting applications engineers, dreams with reality, imagination with realisation. All with one focus; to be the best supplier of explosive systems and solutions to the mining industry in general, and the underground massive mining industry in particular. How does this take place?
Product engineering We first seek to understand our own products; our explosives and blasting agents. Constant research at our West Jordan (Utah, USA) laboratories supported by similar activity at Gyttorp (Sweden), Gullaug (Norway) and Mt Thorley (NSW, Australia) is targetted at exploring new and more efficient means of unlocking the chemical energy contained within commercial explosives. Close co-operative contact with our suppliers stimulates the search for water based explosives with:
• enhanced stability and shelf life; • improved safety and performance
under elevated temperatures and special ground conditions such as presence of reactive sulphides;
• better control over rheology and handling characteristics; and • finely controlled variable density and energy characteristics permitting intra- and inter-hole variation in loading specifications. This is largely proprietary research and it is difficult to dwell in excessive detail on the intimacies of these activities. Suffice to say, however, that Dyno Nobel is well staffed, around the world, with specialist blasting applications engineers. These engineers, with their constant demand for greater control over blasting outcomes, provide constant feedback to the core research teams. The intimate contact between these blasting applications engineers and you, our customers, ensures that this feedback is closely aligned with your needs, ambitions, indeed your dreams.
Delivery systems From our products themselves we turn our attention once again to delivery of the energy contained in the explosive to the rock mass. We share with the industry the fundamental task of
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extraction of value from in situ rock. For us, this begins with the application of explosive energy to the rock mass. An apparently simple task, the distribution of energy throughout the rock is in fact a multi-faceted process involving the application of modern drilling equipment and the ever more sophisticated skills of blast design engineers. With the blast holes drilled it falls to ourselves, the explosive supply member of the team, to apply the final adjustments to this process, once again in co-operation with the mining company’s blast design engineers. Naturally we are relying on you, our customers and our partners in the drilling equipment field to place the blast holes as close to their design position as possible. The importance of efficient and effective explosive delivery systems cannot be overstated in this discussion. Transforming design into actuality requires the ability to deliver consistently and cost-effectively the required amount of energy to specified volumes of rock. This fine control will be achieved through the application of bulk explosive delivery systems for all configurations of blast hole; up and down for production blasting, as well as horizontal tunnel development. Dyno Nobel formulation design and emulsion gassing technology will allow the delivery of ample energy where most needed for production fragmentation. Variable density control in up and down holes, and horizontal string loading will allow underground mass mining engineers to ‘dial back’ the explosive energy intensity where necessary to reduce back break damage to walls, backs, pillars, and draw point brows. More will be said of this shortly. The other advantage of modern delivery systems from which mass miners will benefit increasingly during the next five years is the increased productivity associated with the use of dedicated systems. Manpower requirements will be reduced through outsourcing the loading of blast patterns to specialist crews whose combination of technological support and experience means a shorter loading cycle time with less product waste. Combining shorter loading times with fewer more experienced operators and a growing tendency to more remote and automated operating systems means removal of personnel from hazardous positions and further enhanced safety. Use of PLC based systems to control the delivery of the specified product allows us to move towards automation of much of the explosive loading process and will soon permit a more complete and efficient interaction with blast design models and software systems. The end result will be a much enhanced degree of control and explosive/rock matching on a hole by hole, ring by ring basis. Such a capability to finely control the explosive delivery system returns the focus of blast outcome engineering to the blast design engineer. Once this intrepid individual (or team) has established where and how much energy is to be placed in contact with the rock, the final task is to determine the when.
Blasting timing systems There is little doubt that the mining industry is keenly interested in the latest developments in the Electronic Initiation Systems, which are constantly reported at blasting conferences and workshops around the world. While perhaps apparently less flamboyantly visible than the majority of the manufacturers and representatives of this product type, Dyno Nobel is no less committed to the concept of electronic initiation of blasts, particularly in the underground massive mining industry. First generation electronic detonators have been used by us since 1989 on a number of important projects in Europe and Canada. Some of these projects have involved objective
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DESIGNING AND DELIVERING EXPLOSIVES SYSTEMS AND SOLUTIONS FOR THE UNDERGROUND MASSIVE MINING INDUSTRY
independent research centres such as SveBeFo, and the consistent focus of the work has been on the quantitative verification of theories surrounding electronic initiation of explosive charges and tangible benefits to the cause of managing the rock mass (damage, fragmentation, displacement). An ample selection of references describing this work is supplied at the conclusion of this paper. Bergqvist (1999), for example, reported on results achieved in conjunction with SveBeFo which indicate that some benefits for tunnel wall control are clearly achievable using the kind of timing only available from Electronic Delay Systems. Even now, the Generation 1 system is being employed in Canada as part of the INCO sponsored Mine Automation Project in Sudbury. Recently a fully electronic shot was due to be initiated by the Prime Minister of the country employing the Dyno Nobel DynoREM remote initiation system, from the surface. In this product/system configuration, we believe we are promoting the way forward, both in terms of control (EDS/DynoREM) as well as safety (by removing the miner from the zone of maximum hazard). Our dedication to safety and quality remains a constant part of the development of the Generation II system, which is planned to undergo extensive field trialling during the first half of CY2002.
Blasting applications Within Dyno Nobel, we have developed a team of specialist blasting applications engineers in order to assist the underground mass mining industry in optimising blast outcomes, leveraging a combination of technology, experience and product knowledge. Strong emphasis is placed on field measurement techniques in order to quantify the interaction between explosive, rock and mining systems. The technological toolkit routinely applied by Dyno Nobel in underground mines includes in-hole continuous VoD measurements, blast vibration monitoring and fragmentation estimation through photographic techniques. A number of predictive blast models are also used to explore the consequences of suggested blast design variations. These blasting applications may be used on an ad hoc basis in order to satisfy intermittent needs and demands, however more productive gains are available if the technological resources of blast engineering are more completely integrated into the full stream production process of the mining company. In order to better pursue these gains, Dyno Nobel has developed a novel consultative approach for working with selected customers in order to optimise the downstream processes which are affected, for better or worse, by blasting. The name of this initiative is DynoConsult.
THE NEEDS OF MASSIVE MINING: DYNOCONSULT Through our products and systems, Dyno Nobel has an effect on a multitude of mining process outcomes subsequent to actual blasting. DynoConsult was formed in 1998 to work more closely with our industry customers to optimise these down stream processes that we affect. In order to achieve success in this quest we will continue to work closely with customers to understand these processes; exploring points of contact with our own activities and tailoring the outcomes of these contacts to meet or exceed customers’ expectations. This work enhances our own understanding of mining processes and often that of our customers themselves as we learn together. The path to this learning might lie through a mine to mill optimisation study, concentrating on the ‘product’ which is delivered to the mill. As discussed previously, successful massive underground mining will depend on efficient productive
MassMin 2000
movement of broken ore from production stopes through the mine to the surface. The productivity of this process will be directly affected by a number of aspects of the broken rock such as size distribution and control of drawpoint productivity. It will also be linked to an understanding of the materials handling system within the mine. The future for DynoConsult lies in working with the mines and specialist third party experts such as the JKMRC in order to achieve both this understanding and improved predictive control over the process efficiency. Alternatively, the DynoConsult approach may be configured to examine, quantify and control the negative aspect of blasting; namely damage. Such undesired and unplanned alteration of the surrounding rock mass will have a number of negative consequences for underground mass mines such as:
• • • • •
reduced operator safety and increased ground support costs; increased cost of backfill where applicable; reduced regional stability due to damage to pillars; increased fall off of oversized material; and subsequent reduced drawpoint/excavator performance.
Through close consultative contact with mine geotechnical engineers, we will seek to obtain a better understanding of the host rock in which we work to improve the selection of explosives and applications to achieve our joint objectives. Once we have ascertained the needs and limitations of the rock mass we will be able to apply fragmentation, vibration, and damage models and measurement techniques in order to bring about the best blasting outcome according to the needs of the mine production process in question. These outcomes will be achieved by integrating the rock condition with known performance characteristics of our products and systems so that objectives such as perimeter control (tunnelling and stoping) and dilution will be incorporated with optimised fragmentation. There is a strong tendency to couple physical blast outcomes to financial results by means of process cost analysis and a proof of hypothesis approach to demonstrate achieved savings. A key feature of DynoConsult is that we judge our success by that of our customers. For this reason, and recognising the constant need to reduce production costs, we are willing to link our financial reward exclusively to your success. This means that our underground customers will be able to access state of the art blast engineering consulting, confident in the knowledge that payment can be a negotiated share in quantified, agreed savings or increases in mine revenue. Alternatively DynoConsult can explore new innovative forms of benefit sharing that minimise the impact on customer cash flow. These new methods of remuneration for consultants will become an increasingly important consideration in the selection of partners for mining companies into the future. The major thrust of this partnering is enhancing the quality and level of service from Dyno Nobel to the underground massive mining industry to more fully integrate our role with the activities and processes of our customers. A tangible example of how DynoConsult can work with an underground massive mining operation was showcased in a technical paper published at the 1999 Fragblast 6 Conference in South Africa. This joint publication between DynoConsult and the management of the Olympic Dam Operations described the steps taken in order to meet a technical challenge of significant economic consequence. In simplistic terms the regular incidence of ‘bridging’ or ‘freezing’ of slot blasts was provoking a number of problems for mine staff, including;
• deterioration in the safety of the local working environment; • delays in bringing stope production on line, often for many days;
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• production of oversize due to damage to pillar walls and subsequently blasted rings; and
• the necessity for expensive (up to $A100 000 - per ‘bridged’ stope in some cases) re-drilling campaigns. Those interested in a further explanation are directed to the Fragblast 6 paper for details of how this problem was solved. In broad terms however, DynoConsult and the DNAP Underground Industry Team, working closely with ODO Operations and Planning staff, explored and quantified the effects of rock properties, drilling pattern design, bit size, explosive selection and timing, on the breakage requirements of each stope in a very complex orebody. Today this association continues with DynoConsult working closely with ODO to continuously improve and extend blast design theory and application at the mine. The conversion of technical solutions into the realm of mine to mill scoped projects will soon follow. This is an important part of our future over the next five years with Dyno Nobel (and DynoConsult). We offer the industry our undertaking that you will not stand alone, that our specialists will bring additional perspective, resources and energy so that the challenges of mass mining become opportunities for us all. We believe that synergistic, creative cooperation will become the norm and that mutual enrichment, both professional and financial will become commonplace wherever and whenever our paths cross.
EXTERNAL CO-OPERATION The core business of Dyno Nobel is the design, manufacture and delivery of explosive systems and solutions. There are significant stores of internal knowledge and expertise within the company, which are applied to the benefit of our customers. Nevertheless it is imperative to recognise the fact that few, if any, organisations enjoy a monopolistic knowledge of mining matters, including rock breakage. For this reason we work in open cooperation with such independent sources of research and development and learning as the Julius Kruttschnitt Mineral Research Centre (one of the organising partners of this conference). Dyno Nobel and DynoConsult have adopted a strategy of actively pursuing a Strategic Alliance with the JKMRC in an effort to mutually enhance our knowledge and extend our mutual learning opportunities, all to the benefit of our common customers. Already actively in a number of mine to mill studies in the open pit mining arena, we hope to be able to commence the same analytical and creative approach to whole process optimisation in the underground environment before very much longer. Each one of us is aware that the most fertile ground for the application of continuous improvement programs in process efficiencies are those where production rates or cycles of activity are high. Underground massive mining certainly lends itself readily to this scenario and we will be pursuing this philosophy most energetically with customers, current and future. Reflecting our global nature, and our desire to leverage it, we also sponsor and interact with the activities of the Swedish entity SveBeFo, from whom there is much to learn regarding rock fragmentation and damage. In conjunction with these two world renowned research centres, Dyno Nobel will be working to facilitate and catalyse
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new moves forward in order that the industry in general and the mass mining industry in particular is well placed to reap the benefits that correctly applied, appropriate technology and training can bring about.
CONCLUSIONS We stand today on the threshold of a new millennium. Few of us would expect to enjoy the necessity of long-term planning for that full-term however we look forward, at least, to the next five years. We envisage five years replete with challenges, advances and finally, success. The nature of the company here assembled dictates that no other outcome will be acceptable. The tasks associated with wresting mineral wealth from massive underground orebodies which are ever deeper, ever hotter and ever more hostile, are manifold. The tools and techniques with which all of us will carry out these tasks will be developed and improved through a process of consultation, communication and cooperation. As such, Dyno Nobel and DynoConsult see our technology growing, maturing and adapting to the demands we will face together. During the next five years we will welcome each opportunity to overcome the technological and human challenges, both those which you, the industry, will bring to us, as well as those which we proactively seek to identify on your behalf. We feel confident that our global company stands ready to offer you, the global mass mining industry, the level of support that you need and deserve. This support will embrace Safety, Technology and Innovation as well as a passionate desire to exceed your expectations at every opportunity. It has been our profound pleasure and privilege to work and plan with you this day. It will continue to be so throughout the next five years, and well beyond.
REFERENCES Adamson, W R and Bailey, J J, 1999. Use of blast engineering technology to optimise drilling and blasting outcomes at the Olympic Dam underground mine, in Proceedings Sixth International Symposium on Rock Fragmentation by Blasting (South African Institute of Mining and Metallurgy). Bergqvist, I, 1999. Test of SSE String Loading for Stockholm Ringway Tunnels, Internal report, Dyno Nobel Europe, December. Niklasson, B and Keisu, M, 1992. New methods for contour blasting using electronic detonators and water-notched boreholes, including longer drift rounds without large cut holes, in Proceedings Fourth High-tech Sem Blasting Techn, Instr and Explosives Appl BAI, Allentown, PA. Olsson, M and Bergqvist, I, 1993. Crack lengths from explosives in small diameter boreholes, in Proceedings Fourth International Symposium on Rock Fragmentation by Blasting (Ed: H P Rossmanith) pp 193-196. Olsson, M and Bergqvist, I, 1996. Crack lengths from exposives in multiple hole blasting, in Proceedings Fifth International Symposium on Rock Fragmentation by Blasting (Ed: B Mohanty) pp 193-196. Ouchterlony, F, Olsson, M and Båvik, S-O, 2000. Perimeter blasting in granite with holes with axial notches and radial bottom slots, FRAGBLAST – Int J Blasting and Fragmentation, 4(2000):55-82.
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Digital Blasting — An Opportunity to Revolutionise Mass Underground Mining D Kay1
free face
ABSTRACT TM
This paper discusses how the i-kon Digital Energy Control System may be deployed by miners of massive underground orebodies to deliver significant advantage in terms of safety, cost and environmental compliance. Digital Energy Control offers an exciting opportunity for mine operators to review their blasting and mining practice to gain the most advantage from this new technology.
Underground mining is a complex affair. Not only do mine operators have to produce from an environment as diverse as mother nature can devise, but they must meet increasing expectations opposite employee safety, protection of the community and of the environment. This, coupled with volatile commodity prices, and shareholders who rightly demand a reasonable return on their investment, make for interesting challenges into the new millennium. Mine operators, therefore, need to look to any advantage that can deliver improved safety, reduced cost or greater community compliance to secure their future. In turn, those who supply the mining industry, must also secure their future by enabling their customers to meet these goals. This paper discusses how the i-konTM Digital Energy Control System may be deployed by miners of massive underground orebodies to deliver significant advantage in terms of safety, cost and environmental compliance. Initially a review of some blasting theory and how this is applied to current blasting practice is followed by typical blasting applications, and issues faced by underground operators that dictate current practice. After introducing the i-konTM Energy Control System, the paper will discuss how the system capabilities can be used to benefit key areas of the mining process.
SOME THEORY Effective underground blasting requires three basic elements as shown in Figure 1:
• Free face - a surface for shock waves to reflect off which allows tension cracks to develop
• Void - sufficient space to accommodate the swell of blasted material without choking
• Relief - the time taken for blasted material to adequately move and provide space for subsequent firing blastholes. Too little relief can result in ‘freezing’ where blasted rock can reconsolidate upon the influence of later firing blastholes. The job of the underground blaster is to develop a free face and void. Figure 2 shows how the free face and void is developed. Initially the free face and void required for blasting is provided by holes drilled into the rock close to the blastholes. When a blasthole is fired the compressional wave generated by the explosive shock will reflect from the void hole wall allowing Electronic Blasting Specialist, Orica Explosives, Technical Centre, George Booth Drive, Kurri Kurri NSW 2327. E-mail: [email protected]
MassMin 2000
(transition)
crushed zone
INTRODUCTION
1.
out going compression waves
radial crack
reflected tension waves original blasthole expanded blasthole
seismic (elastic) zone
FIG 1 - Systematic blasting mechanisms.
cracks to propagate under tension. The explosives gasses generated by detonation then swell the cracked rock and eject pulverised material from the void. Once the first blasthole has fired and ejected, the area vacated by the first blasthole provides a free face and a larger void for subsequent firing blastholes If too much rock is blasted at one time into a small void, then the rock may reconsolidate or ‘freeze’. To avoid ‘freezing’, the delay detonator was developed to allow blastholes to fire in a sequence to develop successively larger voids as shown in Figure 3. Delay #1 would fire to a reamer (or void) and allow sufficient time for the blasted material to move and eject from the reamer before Delay #2 fires. The material blasted by Delay #2 would then use the void created by blasthole #1, then Delay #3 into #2 and so on. Delay detonators, therefore, generate successive free faces and voids by providing adequate time for blasted material to move and create a larger void for the next blasthole(s) in the sequence. The time required between firing blastholes to allow void creation is known as ‘relief’. Smaller voids need more relief than larger voids. It is by this process, that small holes become big holes. Reamers become box cuts; box cuts become rises; rises become slots and slots become stopes. A relationship also exists between the delay timing and blast performance to produce productive rock fragmentation for excavation and comminution (crushing and grinding). It is standard practice in open pit blasting, where development of void is not an issue, to use millisecond delay timings that provide better blast outcomes for the whole mining process. This is also used underground in stopes where sufficient free face and void is available. The difference in blast performance using different delay timing is significant, and has been well documented over the years.
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FIG 2 - Sequence of developing free face, void and relief using reamer holes and delay detonators.
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FIG 3 - Further development of void and free face using delay detonators.
Starting Signal
Pyrotechnic Composition
Delay Element
Base Charge
FIG 4 - Schematic of a pyrotechnic delay detonator.
To achieve underground delay blasting the accepted norm is the use of pyrotechnic delay detonators. The detonators, upon receipt of a signal (either electrically to a fusehead, or flame transmitted via non-electric signal tube) starts a delay element to burn for a set time before initiating the detonator. Pyrotechnic delay detonators as shown in Figure 4, are manufactured with set delay times at the factory, in two different delay series:
• Long period range – where the delay between detonators is 200 milliseconds (ms) or greater, and generally used in underground applications where ‘relief’ is critical to blast outcome especially in tunnels, rises and longhole winzes.
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• Millisecond range – being delay periods of 25 milliseconds and greater between delay numbers, and used in surface bench blasting and underground stoping where void and relief is not an issue. The manufacture and use of pyrotechnic delay detonators has been refined over the years such that they can be used in a variety of applications and have a high level of reliability and user acceptance. For complex blasts, combinations of in-hole and surface delays can be used to extend the range of available delays whilst still maintaining a total burning front.
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DIGITAL BLASTING — AN OPPORTUNITY TO REVOLUTIONISE MASS UNDERGROUND MINING
The total burning front ensures that all detonators are activated within their respective blastholes, before the first blasthole fires. This is critical to underground mining due to the confined space and the enormous airblast produced by early firing blastholes. Should initiation be reliant on surface initiators lying outside the blasthole then the airblast could disrupt them, resulting in catastrophic blast failure. Underground blasts require a ‘total burning front’. The limitation of pyrotechnic delay detonators is related to the inherent variation in the delay composition burning speeds and range of available delays. Pyrotechnic delay detonators are designed to fire at a nominal time, and manufactured to fire as close to that time as possible. Relatively small variations in burning rates are exaggerated in the higher delay numbers such that the absolute firing times may be different to their nominal. The pyrotechnic delay range is designed to ensure in-sequence firing at the longer delay times by lengthening the delay interval between numbers. Whilst this may be detrimental to blast performance, guaranteeing firing sequence is more important for blast applications where long delay times are needed.
Fully programmable digital detonator Figure 5 shows a photograph of the i-konTM detonator. It contains a microchip, energy storage capacitor, safety structures and conventional explosive components. The microchip circuitry includes an oscillator for timing, memory for retaining its programmed delay, and communication functions to receive and deliver digital messages to and from the control equipment.
SOME PRACTICES In terms of practical underground stope blasting, the primary consideration is one of ‘risk management’. A significant expenditure is undertaken to locate, define and develop orebodies. Only after the infrastructure is in place, and extraction commenced, do mines start returning revenue. Should full extraction of the projected mineral resource at the desired grade be jeopardised, then dire economic consequences are faced. The opportunity to recover stopes from failed blasts may be hampered by limited access and safety of personnel and equipment. The approach taken therefore is one of caution, to minimise the risk associated with blast failure and the associated cost of recovery. This is often reflected in conservative mine design and production scheduling. Conservative approaches include; multiple stope access drives, firing progressive numbers of small blasts, full opening of rise and slots before commencing ring firing and close sublevel intervals. Some of these designs are dictated by local ground conditions, ore body geometry, pillar requirements to support openings, equipment limitations and considerations to the safety of personnel. Some, however, are limited by the capability of the current blast initiation systems to deliver the flexibility and accuracy of delay sequencing to perform blasts that can optimise all aspects of the mining process. At some time in the life of an underground operation the risk equation can change. These circumstances can include changing stress regimes resulting in pillar failure, need to recover pillars, loss of access, frozen slots, failing hanging/foot walls, or deterioration of ground conditions such that safety of personnel cannot be assured. Under these circumstances the risk of firing a large blast is outweighed by the risk of losing the remaining ore. This style of blast is often referred to as a ‘mass blast’. These are usually complex and require specialised design and sequencing using combinations of in-hole and surface pyrotechnic delay detonators to provide a firing sequence with a total burning front. Often, the initiation design is a compromise to ensure a large number of blastholes can be fired using the available combination of delays, so that all the ore is blasted to rest in the stope draw points from where it can be dug. The expanding nature of the pyrotechnic delay range dictates the use of non-optimal blasthole timing, limited by available delay combinations, with multiple blastholes on the same delay and long intervals between blastholes higher in the delay range.
FIG 5 - Fully programmable i-konTM Detonator.
The capacitor can store sufficient energy to run the microchip independent of external power for 8 seconds with enough energy remaining after this time period has elapsed to fire the fusehead.
i-konTM Logger The i-konTM Logger, shown in Figure 6, is used to communicate with the Detonators during hookup. Operating at an inherently safe voltage, the Logger recognises and tests each Detonator as it is clipped onto the harness wire. Each Detonator ‘responds upon connect’, giving the operator reassurance of knowing the connection is good, and that the Detonator has responded. The required delay time for each Detonator is entered and written into the Logger memory. This information is stored in non-volatile memory (hard memory) of the Logger and used to program each Detonator only during the firing sequence. At any stage the Logger can be used to test the hook-up and response from every Detonator.
THE i-konTM ENERGY CONTROL SYSTEM The i-konTM Digital Energy Control System consists of the following components:
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FIG 6 - Inherently safe i-konTM Logger.
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i-konTM Blaster TM
The i-kon Blaster is used to fire the blast and only deployed from the firing position once the operation is clear of personnel. It is the only piece of equipment that contains the required voltages and codes capable of firing the detonators. The Blaster communicates to each Detonator in turn via the Logger. Figure 7 shows the full system consisting of a Blaster and 8 Loggers, giving a system capability of 1600 Detonators per blast.
DIGITAL ENERGY CONTROL IN MASSIVE UNDERGROUND MINING This section will discuss the applications of Digital Energy Control Systems in massive underground mines and the potential benefits these may bring.
Digital energy control applications Firing larger blasts Correct sequencing and accuracy can now be assured, thus larger blasts can be fired without compromising the delay timing effect on fragmentation and other key blast outcomes. Large blasts can be fired at very much reduced risk, and have their delay sequencing optimised to maximise fragmentation and minimise vibration. Provided adequate void is developed, mass blasts can be considered for the ‘norm’ as opposed to the ‘exception’, due to elimination of risk associated with firing sequence.
Develop relief anywhere within a blast
FIG 7 - Full i-konTM Digital Energy Control System showing an i-konTM Blaster and 8 i-konTM Loggers.
Blasts can start slowly and then speed up. Provided sufficient void is available to accept the blast swell, a blast can start from the rise or slot followed by production rings. The relief required to ensure clearance of material from the slot or rise can be afforded by slow timing initially, and speeded up to provide productive timing in production rings. Relief can be provided against delicate foot/hanging walls within a blast. Blasts can be started quick, slow down, then speed up again. Tight corners can be catered for by re-establishing relief at any stage.
i-konTM System Safety
Ensure each blasthole fires on a unique delay
The i-konTM Digital Energy Control System complies with the concept of ‘Inherent Safety’. This means the Logger, or other test devices, used over the blasthole are unable to produce enough energy to fire the Detonators even if the Logger develops faults. During manufacture, each electronic firing module undergoes a ‘cull’ test, whereby a firing signal is supplied to each device at a voltage higher than that available from the Logger, but well below that required to normally fire Detonators. Modules that fail this test are rejected and take no further part in manufacture. The Detonators also require complex digital signals to arm and fire. The programming required to generate these signals is not present in the logger. The chances of stray currents mimicking the firing codes has been calculated at one in 16 trillion. The Detonators are also tested for immunity to 600 volts AC, 30 000 volts static discharge and 50 volts DC, radio frequency interference and electromagnetic pulse.
It is possible to minimise blast vibrations and reduce damage to mine infrastructure, delicate stope structures and nuisance to community by limiting the maximum instantaneous charge by providing each blasthole with a unique delay.
Key features of the i-konTM Digital Energy Control System
Channel blast vibration frequencies By selecting appropriate delay timings, blast vibration energy can be channelled such that the predominant energy falls into frequencies outside the resonant frequency for structures. This limits the effect of blast vibrations on the community and mine structures.
Fine tune blast timing to maximise fragmentation Experience has shown that subtle changes in blast timing can make significant changes to rock fragmentation. Digital Energy Control Systems are fully programmable, offering the possibility of using novel timing sequences to optimise fragmentation in underground mines.
Key features are:
Simplify stock inventory
• inherently safe Loggers and test equipment; • fully programmable Detonators in one
As i-konTM Detonators are fully programmable, all blasting tasks can be performed with a single inventory item, thus eliminating the need to purchase and hold a large range of initiating products.
millisecond
increments from 1 to 8000 milliseconds;
• accuracy is less than one millisecond; • ‘respond on connect’ hook up system; • full two way communication between Detonators and control equipment;
• fully testable at any time; and • capacity of 1600 detonators per Blaster.
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Benefits to mass underground mines The following benefits are achievable:
Reduce stope development infrastructure Many operations form the rise and slot in the centre of the stope and fire rings from both sides into the slot. This method requires
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DIGITAL BLASTING — AN OPPORTUNITY TO REVOLUTIONISE MASS UNDERGROUND MINING
two access drives to allow charging crews access to both sides of a fully formed slot. The timing flexibility afforded by Digital Energy Control enables the top of the slot and/or rise to be left intact and only removed as the first part of the stope blast. This will allow access to both sides of the slot for charging. By blasting this way, the second access drive can be eliminated, representing a huge saving in development costs.
Decrease stope charging times The majority of time taken to charge stopes is generally in preparation of the blast site and blastholes after the previous firing. Backs and walls have to be scaled and made secure. Blastholes must be cleaned and prepared ready for charging. Those blastholes closest to the previous blast often suffer damage to the extent that significant remedial work is required to clear them. Back break beyond the last ring can render the first rings of the next blast inaccessible, or damaged beyond repair. Digital Energy Control allows larger blasts to be fired thus reducing the number of blast preparations required to complete a stope.
BLASTING EXPERIENCE WITH THE i-konTM DIGITAL ENERGY CONTROL SYSTEM Mt Isa Mines, George Fisher – 614D Trial Stope The concept for George Fisher was to modify the conventional open stoping system to provide maximum support to the opening and to minimise exposure of operators and equipment. The first stope, 714D, was designed to test the new mining method and geotechnical concepts. Figure 8 shows a schematic of the 714D stope at George Fisher 12 - 13 levels. A representation of a main ring of blastholes is shown between the 12 and 13C sublevels while a representation of the cut-off slot and winze is shown between 13C and 13 levels. The type and density of cable-bolt support is shown in the hanging wall, back and brows of the stope.
Improved safety for personnel Firing larger blasts as a routine will reduce the exposure of personnel to re-entry to stopes for preparation of the next blast. Small blasts that nibble away at the orebody can throw stresses back to the remaining section of the stope. This can significantly deteriorate ground conditions, thus jeopardising personnel required to charge the next section.
Improved productivity Bigger, better controlled blasts will produce a better overall fragmentation. Each small blast that is fired disrupts the ground for the next blast. The next blast will therefore be pre-conditioned (cracked), resulting in large, disjointed blocks, back break and dislocated blastholes. Consequently, charge distribution is poor and blastholes overburdened, resulting in poor blast results. This results in poor digging rates and costly blockages within the stope. A uniformly, well fragmented ore pays dividends both for haulage, hoisting and mill throughput. This can be achieved not only through bigger blasts, but also by applying delay timing effective for fragmentation, right through to the end of the blast. This can only be achieved with programmable delays.
Reduced dilution The ability to fire a stope starting from rise or slot, allows for support to hanging/foot walls up to the final firing. In delicate ground conditions, firing the whole stope in one blast, will put the blasted ore into the draw points awaiting excavation. Any failure of foot/hanging wall after firing will sit on top of the blasted ore. Given suitable drawpoint control and draw down of the ore, dilution will not be encountered until ore extraction is complete. Damage to foot and hanging walls can be reduced by engineering relief against these structures through manipulation of the delay sequence. Vibration can also be minimised to reduce disruption to delicate stope structures.
Less damage to mine infrastructure and greater environmental compliance Reduced vibration from blastholes firing on unique delays and the potential to channel frequencies into non-resonant bands.
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FIG 8 - Schematic view of the 714D test stope at the George Fisher Project.
The stoping concept was to develop the necessary void, to pull the winze to within four metres of breakthrough, the slot to within ten metres of the level, leaving a temporary crown pillar on the sublevel. This allows blasting personnel access to charge the remainder of the stope on each sublevel. It also provided a greater degree of support to a weak hanging wall. The 714D had a total height of 65 metres requiring it to be blasted in two separate parts with an intermediate sub level (13C Sublevel) for access to the first lift. To extract the total stope of 80 000 tonnes would require formation of two longhole winzes, two cut off slots and firing of two separate main blasts. The first lift between 13 level and 13C sublevel would be blasted in its entirety before extracting the second lift between 13C sublevel and 12 level. The blasting in the longhole winze and cut-off slot of the first lift was carried out using conventional pyrotechnic delays and run-of-mine timing in order to gain data to allow the definition of the programmed i-konTM delay period. Analysis of the vibration data showed that a delay period of 20 ms was optimal. This value was used to program the i-konTM detonators in subsequent blasts. Because of the necessity of leaving a temporary crown to allow access on the top sublevel of each blast, a complicated timing sequence was required to allow the final blast to take place. The remaining winze, slot and full rings would be fired in one main blast. The ‘direction of blasting’ needed to change three times during the blast and the required pyrotechnic delay pattern was unnecessarily complicated and exposed the blast to a significant
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risk of failure. Conventional firing would also require multiple blastholes to fire on the same delay, thus possibly increasing ground vibration. Since no access would be possible after the final blast, the ability to rectify problems resulting from blasting was extremely limited. Critical to the success of the blast is removal of the winze by use of slow delay timing. This would then be followed by quicker firing of the slot followed by fast firing of the rings either side of the slot. The i-konTM initiation sequence was designed to ensure all blastholes fired on a unique delay to minimise ground vibration. The first Lift Main Blast was fired using 200 ms in the winze and 20 ms in the slot and rings. In order to successfully remove the winze, its duration was taken to 2200 milliseconds, followed by the cut-off slot from 2240 milliseconds to 2760 milliseconds, and finally four full rings (two each side of the Cut.Off.Slot) with a blast duration of 1760 milliseconds. The duration for the total blast being 4540 milliseconds or 4.54 seconds. The second Lift Main Blast was fired with a similar timing design, with overall successful outcomes (Tucker and Kay, 2000).
WMC Junction Gold Mine, 611-2 South Stope Blast A schematic view of the 611-2 South Stope at the WMC Junction Gold Mine is shown in Figure 9. The open stope blast consisted of six main rings of seven or eight, 89 mm diameter. upholes per ring that were required to fire into the main stope void. Following this, and as part of the same blast, an adjacent tabular (narrow) stope was to be fired starting from its initial raise. Due to the proximity of the two ore bodies it was felt that, if fired separately, damage to the adjacent blastholes and drive would be excessive and jeopardise extraction. This type of blast would be very difficult to design using a limited delay range of fixed pyrotechnic delays. To achieve a successful outcome i-konTM detonators were used to design a blast which would start with the main rings using
Open Stope
Slot & Raise
611-2 Sth Footwall Drive
Main Rings
611-2 Hanging Wall Drive
Narrow Stope
FIG 9 - 611-2 South Stope, WMC Junction Gold Mine.
conventional fast timing, then slow down to accommodate opening of the slot around a box-hole rise, followed by a medium firing to allow relief in the narrow orebody. The blast consisted of 72 blastholes of varying length, double and triple primed for a total of 185 detonators. The blast was a success.
Normandy Mining Golden Grove, 460 Z21 STP2 A schematic view of the 460 ZS1 STP2 Stope at the Normandy Golden Grove Mine is shown in Figure 10. The mine had been left in a predicament due to ground conditions and a pillar that had failed beneath. Firings in the longhole rise caused slabs to fall from the stope and it was feared that any further firings of the orebody would result in catastrophic failure and complete loss of this particularly high grade zone.
Slot fired in 1.8 seconds
Remaining 21 rings fired in 2.5 seconds
Rise complete in 3.2 seconds
FIG 10 - 460 Z21 STP2 Stope, Normandy Golden Grove.
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66 000 tonnes of high grade zinc ore had to be fired as a single blast consisting of three to six metres of the top of a longhole rise, full slot and then 21 rings The number of delays required, the need to start slowly for relief in the rise and slot followed by faster firing to gain fragmentation and to complete the blast before stress induced failure took place ruled out conventional pyrotechnic delays. The blast fired with the i-konTM Digital Energy Control System consisted of 264 blastholes drilled as either down or side holes of 115mm diameter. A total of 526 detonators were used in a single blast of total duration 7.5 seconds consisting 200 millisecond interval within the rise, 50 ms interval in the slot and generically 10 ms within the ring and 100 ms between rings. The blast was a success.
The ability to reliably fire large, complex blasts can offer significant advantages to miners of massive orebodies who need to minimise cost, and maximise safety, productivity and environmental compliance
ACKNOWLEDGEMENTS The author would like to thank the staff at Mount Isa Mines (George Fisher Project), WMC Junction Gold Mine and Normandy Gold Mine for assistance during the firing of blasts discussed in this paper. The assistance of colleagues from the Orica Explosives Technical Centre and Initiating Explosives Pty Ltd who contributed to the firing of these blasts and review of this paper is acknowledged.
CONCLUSIONS REFERENCE
TM
The i-kon Digital Energy Control System has demonstrated it’s capability at a number of underground operations in large, complex blasts. The system can significantly reduce risks associated with firing large and complex blasts, so much so, that large complex blasts may now be considered for routine production
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Tucker, G B and Kay, D B, 2000. The use of Electronic Blasting Systems at the MIM George Fisher Mine, to be presented at the European Federation of Explosives Engineers, Symposium on Blasting Technique, Munich Germany, September 2000.
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Enabling and Potential Technologies Hydraulic Fracturing as a Cave Inducement Technique at Northparkes Mines
A van As and R G Jeffrey
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Tele-Operation at Freeport to Reduce Wet Muck Hazards
G Hubert, S Dirdjosuwondo, R Plaisance and L Thomas
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In Situ Stress Measurements Using Oriented Core
E Villaescusa, M Seto and G Baird
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Terratec Universal Raiseborer — Raisebore Machine Technology
C Paterson
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The Potential Use of Foam Technology in Underground Backfilling and Surface Tailings Disposal
A J S Spearing, D Millette and F Gay
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The Design, Testing and Application of Ground Support Membranes for Use in Underground Mines
A J S Spearing and J Champa
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Remote Monitoring of Rock Mass Deformation During Mining
M T Gladwin, R L Gwyther and M Mee
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Hydraulic Fracturing as a Cave Inducement Technique at Northparkes Mines A van As1 and R G Jeffrey2 ABSTRACT The Northparkes, Endeavour 26 orebody is a porphyry copper–gold deposit and comprises a moderately to well jointed rock mass characterised by gypsum veining, the intensity of which decreases with depth. The top 480 metres of the orebody is currently being mined by block caving, the undercut dimensions measuring 196 metres in length by 180 metres in width. Continuous caving was never achieved after the completion of the undercut thus cave inducement techniques were sought to maintain caving so as to safely sustain production. The use of hydraulic fracturing as a cave inducement tool was successfully trialled in December 1997 using existing exploration boreholes located midway up the current lift. Subsequent fracturing campaigns have enjoyed various degrees of success and have yielded in excess of seven million tonnes of ore at a significantly lower cost than conventional cave inducement techniques. Monitoring of both the hydraulic fracture system and the rock mass response has provided considerable insight into the geometry, growth and influence of hydraulic fracture networks above the Northparkes block cave, and has initiated additional research into the use of hydraulic fracturing in other mining methods.
INTRODUCTION Geology The Northparkes Endeavour 26 (E26) deposit is a pipe like orebody approximately 200 metres in diameter and extends to over 800 metres in depth. The orebody comprises trachyandesites (volcanics) and finger-like monzonite porphyry (MP) intrusions. Mineralisation is associated with potassic alteration and occurs predominately in stockwork quartz veins. The density and thickness of these veins diminish with distance from the MP intrusions, hence the copper and gold grades are concentrically zoned around the instrusives (Figure 1). The upper half of the E26 deposit is characterised by gypsum veining, the intensity of which decreases with depth. However above approximately 200 metres below surface the groundwater has leached out the gypsum yielding open fractures within the rock mass. Hence the rock mass strength, and ultimately the geotechnical zonation, of the E26 deposit is a function of the fracture frequency and the intensity of both the gypsum and quartz veining. In general the competence of the rock mass increases with depth and inwards toward the centre of the orebody.
Cavability The deposit has been divided into two mining blocks, the first block (Lift 1) extends to 480 metres below surface and comprises 27 million tonnes of ore. Block caving was selected as the most appropriate mining method based on the geomechanical and geometrical characteristics of the deposit. The Laubscher Mining Rock Mass Rating (MRMR) classification system (Laubscher, 1990) was adopted to classify the rock mass and assess its 1.
Geotechnical Engineer, Northparkes Mines, PO Box 995, Parkes NSW 2870.
2.
Project Leader Hydraulic Fracturing, CSIRO Petroleum, PO Box 3000, Glen Waverley Vic 3150.
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FIG 1 - Geological section and plan of the E26 Orebody. (Refer to the Cd-Rom for colour explanation.)
cavability (Figure 2). The MRMR’s for the various geotechnical zones in Lift 1 ranged from 33 to 54, hence continous caving was predicted to occur at a hydraulic radius of approximately 20 to 25. Initial caving commenced once the undercut had attained a hydraulic radius of approximately 23. Sporadic caving followed the advance of the undercut, however, once the undercut development was complete (196 metres long by 180 metres wide), caving ceased. On completion of the undercut, the cave back had propagated to a maximum height of 95 m above the top of the undercut, yielding approximately three million tonnes of caved ore. Increased production rates failed to induce further caving and various means of reinitiating caving were sought. A number of options for inducing further caving were considered, particularly from the underground exploration level development (1 Level) located approximately midway up the Lift 1 orebody which contains numerous drill holes into the cave back. The options included the extension of the undercut, blasting a boundary weakening slot (Zhang and Tong, 1991) blasting a central cave back slot, and hydraulic fracturing. It was initially percieved that little effort would be required to induce continuous caving. In view of the fact that the conventional methods required the implementation of a complex and costly drill and blast program an initial trial of hydraulic fracturing seemed prudent.
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Monitoring The drill holes on 1 Level provided an excellent means of monitoring the cave progression. The cave monitoring system on 1 Level comprised wire extensometers, time domain reflectometers (TDRs) and simple open holes. Both the TDRs and extensometers provided information regarding the response of the rockmass in the vicinity of the cave back, although the extensometers proved problematic and hence yielded little data of value. Open holes allowed for the direct measurement of the cave back and ore pile positions through the simple use of a weight lowered on a wire-line. In addition to the cave monitoring system an ISS microseismic system was initially installed with the intent of monitoring the seismicity associated with the progression of the cave.
HYDRAULIC FRACTURING — CASE HISTORY Initial hydraulic fracturing attempt
FIG 2 - Laubscher’s stability diagram (after Laubscher, 1990).
The deleterious affects of water on the stability of rock structures is well known and this background led to the idea of attempting to induce instability in the cave by pressurising existing exploration drill holes with water. The trial involved packing off the bottom 5 m of the hole and pumping water into this packed off section, using a high pressure (20 MPa), low volume pump.
FIG 3 - Lift 1 cave monitoring system. (Refer to the CD-Rom for colour explanation.)
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HYDRAULIC FRACTURING AS A CAVE INDUCEMENT TECHNIQUE AT NORTHPARKES MINES
Unfortunately no failures were recorded as the water leaked off into the cave through the highly fractured rockmass, thereby preventing any pressure build-up. From this work it became obvious that in order to propagate fractures or pressurise the rockmass near the cave back, a high pressure, high volume pump was required. Similarly such a pump is also necessary to overcome the higher induced stresses found above the cave so as to ensure both fracture initiation and propagation.
Hydraulic fracturing trial – second attempt Prior to resorting to a more conventional inducement option, expertise was sought to improve the hydraulic fracturing methodology, hence the engagement of CSIRO’s Division of Petroluem Resources. As with the earlier trial the hydraulic fracturing was conducted from 1 Level.
Equipment and methodology An inflatable straddle packer system and diesel powered triplex pump were used to conduct the hydraulic fracturing. The straddle packers were connected to AQ drill rods and lowered down a selected borehole using a diamond drill rig. Once the packers were in position they were inflated with water to a predetermined pressure, usually around 5 MPa above the anticipated injection pressure. Water was then pumped under high pressure, by means of the triplex pump, along an injection line and into the straddle section between the packers. The pressurisation of the rock between the straddle section induces tensile stresses along the walls of the hole and eventually fractures the rock or opens existing fractures. Continued injection forces the water into these fractures causing them to open and extend into the surrounding rock mass (Figure 4).
FIG 4 - Schematic illustrating hydraulic fracturing above the cave.
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Throughout each fracture treatment the packer pressure, injection pressure and injection flow-rate were monitored and recorded.
Results from the second trial The second trial comprised three separate fracturing campaigns, each of a one week duration. Over the course of the second campaign a total of 127 hydraulic fracture treatments were conducted in ten boreholes inducing between 2.5 and 3.0 million tonnes of ore to cave. The maximum cave height was increased from 130 to 165 metres resulting in a considerable increase in the average ore column height. The fracture propagation data exemplified classical growth behaviour of hydraulic fractures and compared well with earlier numerical fracture simulations. For most of the treatments the flowrate was kept constant at approximately 400 litres per minute with the gradual reduction in pressure over time attributed to the propagation of the fracture. Figure 5 presents a plot of a single fracture treatment in a hole at 30 m from the collar of the hole. At the end of each injection, the pump was stopped and a valve in the injection line closed, shutting in the packed off interval and fracture. The pressure recorded immediately after the valve is shut is called the instantaneous shut in pressure (ISIP) and represents the pressure in the fracture at the borehole with losses from fluid friction removed. The ISIP is an upper limit measure of the normal stress acting across the fracture plane and this normal stress is equal to the minimum principal stress for pure opening mode hydraulic fractures. The measured shut-in pressures generally increased linearly with distance from the cave back, as illustrated in Figure 6. This phenomenon can be explained as the transition between the low radial stresses at the cave boundary which increases with distance away from the cave. Of particular note was the fact that the ISIPs recorded in re-fractured holes showed a noticeable decrease when compared to the initial ISIPs recorded earlier in the program. With few exceptions, each hydraulic fracture treatment was characterised by an increase in seimicity both during and after the injection period. In several cases this increase in seismicity was followed by significant caving events, none of which proved violent. Analyses of the microseismic events, geotechnical monitoring data and the occasional intersection of a hydraulic fracture with a nearby borehole have provided insight into hydraulic fracture growth. These measurements of fracture growth (see Figure 7) and geometry suggest that the fractures range from single planar fractures (extending over 40 metres from the injection point) to more diffuse pressurisation and opening of existing fracture systems(extending up to 135 metres from the injection point). The seismic events depicted in Figure 8 are similar in distribution to those recorded during monitored conventional oil and gas fracturing work (Warpinski et al, 1996) and those reported during stimulation of hot dry rock geothermal reservoirs (Pine and Batchelor, 1984; Murphy, Keppler and Dash, 1983). Hydraulic fracture growth, combined with pore pressure induced shear movement around the main fracture channel, is consistent with the location and shape of the micro-seismic event cloud recorded during the fracture treatments. Another interesting obsevation was that in a large number of cases the ‘flow back’ fluid contained a visible amount of suspended fine gypsum, most probably eroded and flushed from the joint veins. Consequent to the success of the hydraulic fracturing trials, Northparkes Mines purchased and commissioned its own hydraulic fraturing system in June 1998. Over 1500 fracture treatments have since been conducted in the E26 orebody from 1
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FIG 5 - Pump injection pressure and rate plot (Open Hole #7).
FIG 6 - Bottom hole ISIP values increasing with distance from the cave back.
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HYDRAULIC FRACTURING AS A CAVE INDUCEMENT TECHNIQUE AT NORTHPARKES MINES
FIG 7 - Recorded fracture growth data from intersections with other boreholes compared to growth predicted by a numerical hydraulic fracture model.
Level. Various chemical additives have been tested, the use of a crosslinked gel proving particularly useful nearer the cave back as the gel has an apparent viscosity in the fracture of approximately 1200 times that of water and hence is not lost into cross-cutting natural fractures and into the cave as rapidly as water.
Hydraulic fracturing from surface Although hydraulic fracturing from 1 Level produced substantial failures along much of the cave back, the method enjoyed only limited success in the southern side of the cave. As a result of this uneven cave progression the southern cave face lagged behind the rest of the cave forming a steep overhang (Figure 9). Unfortunately the southern half of the orebody contains the highest grades and the risk of diluting or loosing this ore became a major concern. Hence all subsequent hydraulic fracturing treatments were focused solely on propagating the southern cave back. Drill hole access into the southern cave face was however fairly restricted from 1 Level thereby hampering the success of the hydraulic fracturing method from underground. Consequently a program of hydraulic fracturing from surface was attempted whereby 5 HQ diameter drill holes, drilled along the southern and western cave boundaries, were repeatedly fractured. The objective of the program was to induce instability into the southern face thereby causing it to propagate and ultimately lead to continuous caving. Hydraulic fracturing from surface proved considerably more difficult than from underground largely due to the increased hole depths. One of the main problems encountered involved the deflation of the packers. The hydrostatic head of water in the inflation line was sufficent to keep the packers partially inflated
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thereby hindering their movement. To overcome this problem nitrogen gas was used for packer inflation although this seemed to reduce the packer life. A downhole packer inflation valve was also employed. This valve inflates the packers by using the pressure in the injection line together with a flowrate and pressure sensitive valve. However, reliable ISIPs cannot be measured when this downhole valve is in use. In general the handling of the large injection string and the deployment of the packer inflation system proved awkward and costly. After several weeks of continuous fracturing without success and the loss of several packers and an entire 400 metre BQ rod string, the surface hydraulic fracturing program was abandoned.
BOUNDARY WEAKENING TRIAL The unsuccessful surface hydraulic fracturing campaign forced Northparkes to resort to the more conventional inducement option of boundary weakening. Essentially this option involved the development of a subvertical north - south striking slot located along the southwestern boundary of the cave (Figure 10). The objective of the boundary slot was four-fold: 1.
To cut-off any clamping stresses acting across the southern cave back and create a destressed or tensile stress environment which is known to result in caving.
2.
To increase the minimum span of the southern half of the cave thereby improving its cavability.
3.
To blast the southern cave back to the same elevation as the remainder of the cave. This would place the entire cave back into a seemingly weaker rock mass thereby increasing it caving potential.
4.
To induce natural continuous caving.
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FIG 9 - Schematic of the Lift 1 cave progression in response to the mining of the undercut, hydraulic fracturing inducement and boundary weakening inducement.
FIG 8 - Microseismic events generated during fracturing above the cave back.
In addition to creating the boundary slot, several large diameter ‘strip’ holes were drilled across the southern cave face. The entire blasting operation proved challenging with several holes in excess of 120 metres in length. Although costing an estimated $A 1.4 million, the blasting program proved highly successful in inducing 2.7 million tonnes from the southern half of the cave and achieved all of its objectives with the exception of inducing continous caving.
THE FINAL HYDRAULIC FRACTURING CAMPAIGN From the experience gained up to the time the slot was formed, it was evident that the most successful hydraulic fracturing treatments occurred in holes which covered a large area of the cave within 20 m of the back, ie the zone of discontinuous deformation (Duplancic and Brady, 1998). Based on stress conditions around the cave and microseismic monitoring of the fracture growth, it is believed that hydraulic fractures propagate roughly parrallel to the cave boundary, somewhat analogous to smoothwall blasting (sliping).
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FIG 10 - Boundary weakening slot.
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A decision was made by Northparkes Mines to drill and fracture several subhorizontal holes which paralleled the cave back (Figure 11). This fracturing program was seen as a final effort to induce caving through hydraulic fracturing prior to re-commencing a more costly drill and blast campaign. As with all previous cave inducement programs it was hoped that as the cave back propagated up into progressively weaker material it would eventually attain a state of continuous caving.
2.
The gypsum joint fill has been partially dissolved and eroded by the fracturing fluids. The fine gypsum was then partially flushed out by the water injected into the fracture system. This mechanical degredation of the gypsum has the effect of reducing the tensile and shear strengths along the joints.
3.
Once localised failures have occurred the change in the affected cave back geometry consequently produces larger zones of low stress or tensile zones, which in turn, encourage further instability and caving. The fracturing response and monitoring data show that both opening mode (or conventional hydraulic fracture) and shear mode fracture growth occurred. The plots of pressure versus time fit the classical growth behavior of hydraulic fractures, and the numerical simulations were able to match this growth behavior. The sharp breakdown behaviour that was observed and the very low water loss into straddled intervals before fracture initiation both support the initiation of conventional opening mode hydraulic fractures. The gypsum filling present in nearly all natural fractures undoubtedly limited fluid loss into these natural fractures and limited shear fracturing. Injections into the zone near the cave back did produce shear fracturing response which is consistent with the presence in this zone of numerous stress-induced fractures.
Hydraulic fracturing + caving = future mining Not only does hydraulic fracturing offer an extremely cost effective method of cave induction but it has the potential to improve both the cavability and fragmentation size of an orebody through preconditioning. In essence, mass fracturing of a mining block could reduce the rock mass strength and primary fragmentation size through the creation of new fractures or the opening of existing discontinuities.
CONCLUSIONS
The fracturing of the new holes proved problematic in that the hydraulic fractures initially propagated along the axes of the holes often resulting in substantial packer by-pass. Several hundred fracture treatments were performed occasionally inducing minor caving. Cave back monitoring revealed that the failures were highly localised and confined to the central region of the cave.
The use of hydraulic fracturing as a cave induction tool has proved highly successful at Northparkes Mines with over seven million tonnes of ore attributed to the method at a cost of around AUD$1.1 million. Monitoring information has confirmed that hydraulic fractures propagated in excess of 40 metres from the packed off hole. Pressurisation of the joints around the main fracture was measured to extend to over 130 m. This vast fracture growth has the ability to soften and weaken the rock mass whilst also altering the local stress environment. The fracture data suggests that most of the treatments conducted within 20 metres of the cave back produced shear mode failures and displayed a greater incidence of microseismic activity. Conversely the treatments performed further away from the cave back were characterised by opening mode fractures and served to pre-condition the rock for caving by the introduction of new fractures.
DISCUSSION
ACKNOWLEDGEMENTS
The success of hydraulic fracturing at Northparkes has been attributed to a number of factors, namely:
The authors would like to thank Northparkes Mines for permission to publish this paper. We would especially like to acknowledge both Ross Bodkin and Michael House for their support and encouragement of this new initiative.
FIG 11 - Plan and section showing the hydraulic fracture hole locations.
1.
The propagation of multiple fractures has reduced the overall strength of the rockmass by creating and opening new fractures or opening and producing shear displacement on existing fractures. In addition, the introduction of water under pressure into these fractures reduces the effective stress across them allowing them to slip ultimately promoting instability. Shear movement in the rock mass was more often associated with treatments near the cave back.
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REFERENCES Duplancic, P and Brady, B H G, 1997. An investigation into caving mechanics – review and report of investigation, Northparkes report. Jeffrey, R G and Mills, K W, 2000. Hydraulic fracturing applied to inducing longwall coal mine goaf falls, in Proceedings Fourth North American Rock Mechenanics Symposium, Seattle, 31 July – 3 Aug.
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Laubscher, D H, 1994. Cave mining the state of the art, J S Afr Inst Min Metall, pp 271-292. Murphy, H, Keppler, H and Dash, Z, 1983. Does hydraulic fracturing theory work in jointed rock masses?, Geothermal Resources Council, Transactions, 7(461). Pine, R J and Batchelor, A S, 1984. Downward migration of shearing in jointed rock during hydraulic injections, Int J Rock Mech Min Sci and Geomech Abstracts, 21(5):5-249. van As, A and Jeffrey, R G, 2000. Caving induced by hydraulic fracturing at Northparkes Mines, in Proceedings Fourth North American Rock Mechenanics Symposium, Seattle, 31 July – 3 Aug.
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Warpinski, N R, Wright, T B, Uhl, J E, Engler, B P, Drozda, P M, Peterson, R E and Branagan, P T, 1996. Microseismic monitoring of the B-Sand hydraulic fracture experiment at the DOE.GRI Multi-Site Project, in Proceedings Society of Petroleum Engineers 36450, SPE Annual Technical Conference and Exhibition, Denver, 6 - 8 October. Zhang, S and Tong, G, 1991. Influence of block boundary weakening on the caving process, Min Sci Techno, 13:157-166, (Elsevier Science Publishers).
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Tele-Operation at Freeport to Reduce Wet Muck Hazards G Hubert1, S Dirdjosuwondo2, R Plaisance3 and L Thomas4 ABSTRACT PT Freeport Indonesia, a subsidiary of New Orleans-based Freeport McMoran Copper and Gold of New Orleans, operates a large mining complex on the Western half of the Island of New Guinea in Tembagapura, Irian Jaya, Indonesia. The IOZ (Intermediate Ore Zone) mine, one of the underground orebodies, is mined by block caving. The area affected by the cave collects a large quantity of rain and ground waters mixing with fine rock particles. The mixture, commonly called ‘wet muck’, sometimes spills into drawpoints and makes the working environment hazardous to the loader operators. Wet muck is observed in an increasing number of drawpoints in the IOZ mine and remote LHD (Load-Haul-Dump) technology has been implemented to handle the ore. The driving force for this automation project has been the safety of the loader operators. The system has been developed mainly in-house and incorporates state-of-the-art features. Up to eight remote loaders and nine rock breakers are controlled from an underground control room located approximately 400 metres from the active extraction panels. The two underground crushers and nine conveyor feeders are also operated remotely from an adjacent control room. The development of Freeport’s remote system has taken time and considerable effort. Not only the success of the IOZ mine was at stake but also the fulfillment of environmental commitments pertaining to the mill tailings. Approximately eight per cent of the mill feed must come from underground skarn deposits to neutralise the acid-generating tailings generated by the huge Grasberg open pit ores.
INTRODUCTION The IOZ (Intermediate Ore Zone) Mine is located in the Ertsberg Mining District in the western highland ranges of Irian Jaya in Indonesia. This block cave mine, in production since 1994, produces 19 000 tpd. LHD (load-haul-dump) units with capacities of 6 - 7.5 cu yd handle the ore from drawpoints to grizzlies located on the North and South fringes of the deposits. Rock breakers are used to fragment the oversize rocks. The presence of a large cave surface expression coupled with a highly permeable geological environment allows the surface and ground water to migrate into the column of caved ore. The water mixes with fine particles to form what is commonly called ‘wet muck’. The consistency of wet muck resembles mud with a variable viscosity. Wet muck spills are frequent and generally occur while pulling ore from wet drawpoints. Having LHD operators exposed to such conditions is unacceptable and remote loader technologies have been implemented to minimise risks. The level of technological sophistication has increased with the magnitude of the wet muck problem. Conventional loaders have been replaced by extended line-of-sight remote control systems, followed by tele-remote loaders. Rock breakers have also been equipped with cameras and remote controls. The latest
1.
Superintendent – Underground Capital Projects, PT Freeport Indonesia, PO Box 209, Tembagapura, Irian Jaya 98663, Indonesia.
2.
Superintendent Engineering – IOZ Mine, PT Freeport Indonesia, PO Box 209, Tembagapura, Irian Jaya 98663, Indonesia.
3.
General Superintendent – Tele-communications, PT Freeport Indonesia, PO Box 209, Tembagapura, Irian Jaya 98663, Indonesia.
4.
Senior Manager – Underground Mines, PT Freeport Indonesia, PO Box 209, Tembagapura, Irian Jaya 98663, Indonesia.
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addition to the remote system is the relocation of the crusher rooms and conveyor feeder controls. Rock breakers, crushers and feeders likely to be exposed to wet muck flows are operated remotely. The challenges with the IOZ tele-remote systems continue. Freeport is looking at ways to improve the productivity of the remote loaders. Auto-guidance and auto-loading technologies are being investigated.
IOZ MINE The IOZ deposit was brought into production in 1994. The mining method is mechanised (LHD) block caving. Block caving is a non-selective mass mining method where the fragmentation of the ore largely depends on the action of gravity and pre-existing or induced weaknesses in the rock mass. By blasting a thin layer of ore under the base of the ore column, the vertical support of the ore column is removed and the ore caves by gravity. The cave propagates upwards as additional ore is pulled from the drawpoints.
Setting The IOZ mine is one of three block cave mines that exploit the Ertsberg Mining District (Figure 1). The district is located in the extremely rugged Sudirman Mountain Range of Irian Jaya, the easternmost province of Indonesia, which occupies the western half of the Island of New Guinea.
FIG 1 - Location of PT Freeport Indonesia mines.
The high-altitude tropical climate delivers rainfalls averaging five metres per year. Temperature at the mine site is cool from -2°C to +22°C. The GBT (Gunung Biji Timur or ‘East Ore Mountain’ in the Indonesian language) mine was in operation from 1980 through 1993. The lowest extraction level was developed at the 3625 m elevation. Approximately 60 million tonnes of ore grading 1.95 per cent copper, 0.67 ppm gold and 9.83 ppm silver were recovered. The GBT cave reached the surface topography in 1986 and mud spills were observed.
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The IOZ mine is located 4 degrees-6′ south of the equator and 137 degrees-7′ east longitude. It exploits the deposit between the 3474 m and 3626 m elevations. A total of 43.5 million tonnes grading 1.23 per cent copper, 0.49 ppm gold and 7.14 ppm silver will have been recovered when the mine is exhausted in 2003. There is now a vertical cave connection between IOZ and GBT. The DOZ (Deep Ore Zone) block cave mine, currently under development, will recover block cave reserves underneath IOZ, from 3135 m to 3474 m elevations. Undercutting will start in the second half of 2000 and production will ramp up to 25 000 tpd by 2003. A study is now underway to investigate the feasibility of increasing production up to 35 000 tpd. Wet muck is expected to be a concern in DOZ when there is a cave connection between IOZ and DOZ.
Production LHD units in the range of 6 - 7.5 cu yd are used to draw ore from drawpoints. The ore is dumped onto grizzlies with openings of 450 × 450 mm. Oversize blocks are broken with rock breakers. The ore flow system consists of gathering conveyors and two jaw crushers, ore passes and main conveyors to the surface mill stockpiles. The two jaw crushers are located 20 metres below the extraction level. Rock Breakers #1, 2, 3, 4 and 9 feed the North Crusher while Rock Breakers #5 through #8 feed the South Crusher. IOZ was brought into production at an initial rate of 10 000 tpd. Further delineation drilling added reserves and production is currently 19 000 tpd. Three new panels (#9 through #11) and additional drawpoints in existing panels have been developed to support the expansion program.
Mine layout The footprint of the IOZ orebody is 330 m long and 220 m wide with ore column heights of 150 to 220 m. Drawpoints are first driven at the base of the deposit (3456/L) and undercutting takes place at an elevation 18 m above (3474/L). The drawpoints in the first panels were laid out diagonally (El Teniente straight through pattern) while new drawpoints followed an offset herringbone layout. The change was implemented to facilitate the handling of wet muck: with the offset herringbone layout, loaders buried in wet muck are generally easier to recover. The drawpoint spacing is 17.3 m and the distance between panels 30 m. All drawpoints and panel drives are reinforced with poured concrete and steels sets.
WET MUCK A wet muck flow is a sudden collapse and run-out of wet granular material following some disturbance. Flows can be dynamically or statically triggered by the following:
• Vibrations • blasting, • moving equipment, and • earthquakes. • Change in stress conditions • consolidation within a drawbell, • higher head (water level) in the drawbell, and • mucking.
FIG 2 - Schematic long section of GBT, IOZ and DOZ.
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• Wet muck flows only occur under certain conditions in the IOZ mine
• more than 20 per cent sand-size material (50 mm Moisture Content 50 mm Moisture Content 50 mm were used in the past). Moisture Content 8.5 - 11 % Supervision required
RL
Muck Type
Remote Loader
Supervision required to set up loader.
D - Coarse Very Wet Size +70 % >50 mm Moisture Content >11 % E - Fine Wet Size +30 % >50 mm Moisture Content 8.5 - 11 % F - Fine Very Wet Size +30 % >50 mm Moisture Content >11 %
The class determines the loader type, the role of the supervisor and safety procedures to be used based on muck size and wetness. For instance, Class RL drawpoints require that remote loader be used. The supervisor is to be present to set up the loader. Remote loaders can only be used after the operator and the supervisor barricade all access ways to the wet muck area. Steel and electronic barriers, described in another section, are used for that purpose. The number of wet drawpoints has continuously increased since IOZ production began. In 1994, there were only two wet drawpoints. This number progressively increased to reach 73 in 1999.
MINE PROFITABILITY AND ENVIRONMENT The Grasberg porphyry deposit, lying northwest of the Erstsberg Mining District, contains the largest single gold reserve and is one of the three largest open-pit reserves in the world. More than 90 per cent of the mill feed or 210 000 tpd currently comes from the Grasberg Open Pit. Tailings generated from the Grasberg ores are acid generating and cannot be released in the environment. On the other hand, tailings from the IOZ Orebody contain large quantities of acid neutralising minerals. Approximately eight per cent of underground ore is required to make the mill tailings neutral. Maintaining a mill feed with a minimum of underground ore has been Freeport’s approach in recent years. The IOZ is a low-cost operation and a very profitable mine in itself. The requirement to produce neutral tailings makes IOZ an even more strategic asset. When faced with all the difficulties encountered with wet muck, Freeport management commissioned studies and introduced new procedures to assure the safety of the LHD operators.
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FIG 5 - Recovery of a remote LHD buried with wet muck.
WET MUCK HANDLING Several systems have been used and developed over the years to handle wet muck in the IOZ. Conventional remote loaders were replaced by extended-line-of-sight remote control systems with a small TV monitor mounted on the operator’s harness. As operators were still too close to wet drawpoints, a full tele-remote system was commissioned in February 1999. The system has been expanded since then to operate up to eight loaders from eight consoles located in an underground control room. A combination of steel gates and electronic barriers control the access to the wet muck area where remote mobile equipment is in use. The loaders operated in remote mode are deactivated when the safety perimeter is violated.
Steel cab loaders While Freeport was in the process of obtaining LHD remote systems incorporating video, some conventional loaders were equipped with heavy steel cabs, oxygen bottles and radios. These units, still in use, have been successful for pulling low risks drawpoints and to clean up spills in the panels.
Extended line-of-sight In 1996, RCT® (Remote Control Technologies) from Perth, Western Australia, supplied Freeport with an extended line-of-sight remote system for two loaders. The operator controlled the LHD from the intersection between the panel and
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FIG 6 - Steel cab loader.
FIG 8 - Temporary control room.
At the same time, Freeport staff continued to investigate more advanced technologies to control LHD units from a central control room.
Tele-operation Following several mine visits where tele-remote systems were in development or in operation, Freeport staff initiated discussions with various suppliers of LHD units and remote equipment. Remote systems supported by VHF leaky feeder systems showed limitations in the number of loaders that could operate simultaneously. Other off-the-shelf solutions required Freeport to purchase a whole new fleet of LHD units or did not match all requirements. Freeport staff gained a lot of experience in designing state-of-the-art tele-communication systems such as private phone and internet networks in several locations worldwide involving satellite links, fiber optics and complex telecom equipment. Freeport’s telecom personnel decided to tackle the challenge by developing an in-house tele-remote system. The main design criteria for the tele-remote infrastructure were: 1.
accommodate at least six remote loaders and four rock breakers;
2.
possibility to expand and eventually integrate digital video technology;
3.
good quality of picture with color video;
4.
sufficient video transmission capability (frames per second) to allow adequate control of remote LHD;
5.
ergonomic chairs;
6.
LHD consoles with three monitors (17 inch for LHD front view, 14 inch for rock breaker view and spare 14 inch monitor); and
7.
possibility to add Modular Mining System’s Dispatch® consoles.
FIG 7 - RCT®’s extended line-of-sight system.
the access to the rock breaker. A small video monitor mounted on the operator’s harness provided video feed back to better control the LHD over longer distances and around corners. Several limitations were observed with harness-mounted video monitors. Controlling the LHD unit with such a small monitor was difficult. Operators were still too close to wet muck areas. Using the same remote equipment, makeshift control rooms were established in the rock breaker accesses on the north fringe. Large TV monitors made the remote-control operation easier.
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The current system includes eight consoles for eight remote loaders. Operators generally prefer to use smaller monitors (14 inch) rather than 17 inch to control the loaders. A Modular Dispatch® system to control production on the extraction level was commissioned in early-2000. Remote and conventional loader operators receive instructions from the dispatcher and are prompted to respond using the touch screen menu.
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FIG 11 - Steel gate to block access to tele-remote area. FIG 9 - New tele-room.
protection. First, a sound and light alarm system warn workers when they go beyond the steel gate. If the workers go beyond, the second detection mechanism shuts down all remote loaders.
Rock breakers Large spills of wet muck reached the grizzlies in several occasions. This was a warning sign that rock breaker operators would be at risk and a hard-wired remote system was commissioned in 1999 for the rock breakers on the north fringe. The system was extended to the south fringe in 2000. Rock breaker operators work from three consoles located in the same control room than the LHD units. Two fixed cameras installed above the grizzly chamber provide two views to better control the rock breakers and see if an LHD is coming.
Crushers and feeders
FIG 10 - LHD operator in the new tele-room.
Electronic barriers Remote LHDs create hazards to employees working on the extraction level. Supervisors, secondary blasting and draw control employees, electricians, mechanics and engineering personnel travel mostly by foot and means of restricting access to remote equipment areas must be in place. Initially, Freeport attempted to design a selective system of electronic barriers that would automatically shut down only the remote loader in the panel where the employee violated the safety perimeter. Complex situations were identified due to the large number of accesses to any given panel. The concept for the electronic barriers was simplified by dividing the wet muck operating area in two: panels 1 to 4 and panels 5 to 8. For instance, a violation in panel 1 would automatically shut down all remote loaders in panels 1 to 4. Owing to the harsh environment in the panels, electronic barriers have been difficult to maintain. Mud slides and LHDs often damage them. Steel gates are still used extensively to restrict access to the remote areas. The electronic barrier system has been simplified: motion detectors are installed in close proximity of the steel gates and offer two additional levels of
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Following some large wet muck spills that filled up loading pockets and causing damage to feeders, Freeport realised that operators working in close proximity of the feeders and crusher were also at risk. All feeders and the two crushers are now operated from a control room adjacent to the one for the remote LHD units and rock breakers. All equipment being stationary, the design for the remote system to control feeders and crushers was much simpler than for the remote LHDs. Hard-wired CCTV cameras and monitors with standard electric and hydraulic components were selected.
PRODUCTIVITY AND FUTURE Productivity with the remote LHD units (79 tonnes per hour) is less than with conventional loaders (138 tonnes per hour). The main reason is the difficulty to remotely control loaders in very difficult conditions: water and mud in the drawpoints and limited visibility due the dimensions of the loaders compared to the size of the panel drifts and draw points. Freeport is currently investigating new technologies such as LHD auto-loading and guidance systems. Tire wear and equipment abuse would presumably be minimised. Freeport also believes that hauling time from draw point to grizzly could be improved. The reduction of the number of operators is currently not a priority. The DOZ cave will eventually connect with the IOZ/GBT one. Wet muck is foreseen as a potential problem and the DOZ mine will benefit from the IOZ experience and the in-house expertise developed over the last four years.
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FIG 12 - IOZ remote control room.
CONCLUSION
ACKNOWLEDGEMENT
Freeport has invested considerable time and money in the development of remote technology to safely handle wet muck. The efforts have paid off. Operators were displaced from dangerous workplaces to control rooms and several potential accidents were prevented. Mining in the IOZ mine has continued in spite of early concerns on the longevity of the operation. Freeport is prepared to tackle any wet muck challenge that may develop with the DOZ mine once in full production. Although not fully covered in this paper, aggressive dewatering efforts continue to minimise the infiltration of ground water into the cave. Even if the dewatering program is successful, wet muck remains unavoidable and remote control technology must be available. New machines and systems may remove workers from certain classes of hazards, but they often introduce new ones. Good safety operating procedures are essential.
The authors thank the management of PT Freeport Indonesia for permission to publish this paper. Gratitude is also expressed to colleagues who contributed to the project and reviewed this paper.
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REFERENCES Dirdjosuwondo, Suyono. 2000. IOZ Block Cave Mine. PT Freeport Indonesia Internal Document, 50 pp. HCI Hydrologic Consultants, Inc, Call and Nicholas, Inc. 1998. IOZ Wet Muck Study. Rachmad, Lufi. 1999. Mud Flow/Mud Rush Potential, PT Freeport Indonesia Internal Document, 5 pp.
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In Situ Stress Measurements Using Oriented Core E Villaescusa1, M Seto2 and G Baird3 ABSTRACT
METHODOLOGY
A low-cost methodology that allows the determination of in situ stresses using oriented core specimens has been investigated. The technique can be used to determine the stresses during the early stages of a project, even in undeveloped areas of a mine. The research aim was to compare the experimental results estimated by the Acoustic Emission and Deformation Rate Analysis methods with those estimated by conventional Hollow Inclusion cell measurements. The studies have been focused on full stress tensor determination from a single oriented rock core. In all cases the rock core specimens recollected the in situ stresses reasonably well compared with the standard overcoring method.
The Acoustic Emission (AE) and Deformation Rate Analysis (DRA) methods proposed for estimating in situ stresses from cored rock samples are based on the principle of the Kaiser effect (Kaiser, 1953). This phenomenon suggests that a previously applied maximum stress might be detected by loading a rock specimen to a point where there is a substantial increase in AE activity. The Kaiser effect is a recollection of the immediate previous maximum stress to which a rockmass has been subjected by its in situ environment. This is achieved by loading small cylindrical samples of rock which are undercored from conventional drill cores recovered from the site for which stress data are sought. The specimens are instrumented with two axial strain gauges and/or a pair of acoustic emission transducers. The sample is then loaded uniaxially, and while the strain gauges provide a measure of strain, the AE transducers provide a record of the number of internal ‘events’ with increasing load (and hence stresses). The principle behind both techniques is that changes in the rate of AE and of axial strain with stress occur at the maximum stress level (along the axis of the sample) to which the sample had previously been subjected (Kurita and Fujii, 1979; Seto et al, 1999). Provided that the original core had been accurately oriented and that a sufficient number of specimens can be obtained from the core at different orientations, it is possible to develop a full stress tensor at that point in the rock mass.
INTRODUCTION A number of mines in Western Australia are experiencing relatively high horizontal in situ stress fields, even at relatively shallow depths. Early signs of high stresses have been observed in several of the newly developed underground mines and this has stimulated the general interest for stress measurements. Currently, the attention is focused into the characterisation of in situ stresses at depth, in remote regions that are difficult to access from current mine workings, but where extensive geotechnical data from drillcore is widely available. In particular, a low-cost methodology to carry out stress measurements at a number of locations over a short period of time is being investigated. The measurement technique being considered allows the determination of stresses using oriented cored rock samples that can be drilled within remote and undeveloped areas of a mine. The method has attributes that can make it extremely useful if proven reliable. That is, a large number of measurements could be made at a reasonable cost, while achieving a very large coverage across an orebody. In addition, a representative and detailed knowledge of the in situ stress field could be determined during the early stages of a project (such as mine feasibility studies), even in areas where current access is not yet available. Furthermore, in cases where access is available, the potential exists to measure stress concentration in isolated pillars, and within stope crowns. The information can be used for input and calibration of numerical models and can lead to an optimisation of scheduled extraction sequences. As a result, safer and more economical mine designs are likely to be achieved. Consequently, the objective of this paper is to describe and investigate the application of two relatively new stress measurement techniques which use cored rock (Yamamoto et al, 1990; Seto et al, 1999). The results have been compared with those obtained by conventional overcoring, in a series of in situ stress measurement trials at a number of mine sites. Data were collected from a number of sites with different geological environments and, in most cases, the core was obtained from the same hole where a conventional stress measurement had been carried out.
1.
Western Australian School of Mines, PMB 22, Kalgoorlie WA 6430. E-mail: [email protected]
2.
National Institute for Resources and Environment, Japan.
3.
Western Australian School of Mines, Australia.
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Sample preparation In this study, oriented drill cores having a variety of sizes were used to determine the stresses. The diameter ranged from 51 mm for exploration core to 141 mm for core recovered from HI cell stress measurements. In all cases, each rock core was undercored into several 20 mm diameter samples. The sample lengths were trimmed to about 50 mm and ground at each end. Parallelism between the top and bottom faces of each sample, to within 1/50 mm, was accomplished in all cases. Usually, at least three specimens were prepared for each selected orientation of testing.
Sample loading In the present study, the rock core specimens were repeatedly loaded five times up to a certain stress level under a constant loading rate (7.5 MPa/min) by means of a servo-controlled testing machine. The specimens were usually loaded up to a maximum stress level that was decided taking into account the original depth of the core and the uniaxial compressive strength of the rock. During the repeated loading, axial strains were measured by two strain gauges (gauge length: 20 mm) attached on the surface of the specimen. During the cyclic loading, the Acoustic Emissions were measured by two AE transducers attached on the sides of the specimens. The AE monitoring system used in the test consisted of six NF-7661 AE modules capable of recording the full range of AE parameters as well as performing two-dimensional source location. The AE transducers used in this study were of differential type (5 mm diameter, NF-AE-904DM model) and the resonance frequency was 500 kHz. They had high gains between 200 and 550 kHz. The response frequency band of this system was between 50 kHz and 1 MHz.
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The signals from the AE transducers were amplified 40 dB by the pre-amplifiers, then sent to the AE monitoring system and were amplified further by 40 dB. The threshold level for AE counting was in the range of 150 to 200 mV that was slightly higher than the environmental noise. After the test, the AE characteristic parameters were analysed using the recorded AE event data, and the relations between each characteristic parameter and stress were investigated. In addition, a DRA study was carried out using the results of the strains measured during the test.
AE technique Acoustic emission is a burst of high frequency elastic waves caused by a localized failure such as microcracking within a material. The AE technique is generally based on the Kaiser effect that is an AE phenomenon briefly defined as the absence of detectable AE until the previously applied stress level is exceeded. Consider, for example, a simplified cyclic loading in which a rock specimen is subjected to two cycles of loading. In the first cycle of loading, stress is applied to the specimen at a constant rate up to a stress level (Sp) and then returned to zero. In the second cycle, stress is increased in a similar fashion, however, the previous maximum stress (Sp) is exceeded. During each cycle, AE activity is monitored and recorded as a function of applied stress. In the first cycle, AE activity is generated at all stress levels. However, during the second cycle of loading no AE activity is generated until the stress level (Sp) attained in the first loading cycle is exceeded. The effect was first observed by Kaiser (1953) in the experimental study of metal materials that had the capability to memorise the previous stress level. The Kaiser effect was also discovered in rock materials (Kurita and Fujii, 1979) and it has been found that the Kaiser effect in rock materials is closely related to the extension of microcracks that have been formed in the previous stress state (Seto et al, 1995). The extension of the microcracks induces the active AE and inelastic strain behavior after the previous stress level is exceeded. In addition, the AE is generated inside a rock core specimen by the irreversible movement of a discontinuity or a crack such as shearing and closure as well as microcracking. Since a rock specimen inevitably includes microcracks, the first loading cycle often produces noise associated with crack closure or compaction that can sometimes obscure the Kaiser effect. This noise in the first cycle of loading, however, can be suppressed by subsequent unloading-reloading cycles at stress levels below the Kaiser effect, thereby making the take-off in AE associated with the Kaiser effect more pronounced. Most noise is reduced by the second cycle and the previous stress can be estimated by the clear AE take-off in the second cycle of loading (Seto et al, 1999). Thus, in the present study the AE activity in the second loading is usually used to determine the in situ stress (see Figure 1).
Yamamoto et al (1990) performed cyclic uniaxial loading tests and measured the strain difference values (∆εi,j) during loading between two cycles as a function of the applied stress: ∆εi,j = εj(σ) - εi(σ); j > ι
(1)
where εk is the strain in the k-th loading and σ the applied stress. This function, called the strain difference function, represents mainly the difference of inelastic strain between two cycles. The mechanical behavior of pre-existing cracks in a rock specimen causes non-linear strains with respect to the applied axial stress. For instance, frictional sliding is expected to occur on a pre-existing shear crack when the shear stress exceeds a critical value. An isolated tensile crack may open or close elastically with the change in the axial stress. This type of non-linear behavior in strain is considered to be mostly reversible during many loadings, as long as the pre-existing cracks do not change their size (Holcomb and Stevens, 1980). The reversible components of strain are cancelled by the operation in Equation 1. The axial stress applied to a rock specimen may enlarge some of the pre-existing cracks and create new cracks. Considering what the Kaiser effect implies, this should occur especially when the applied stress exceeds the peak value of previous maximum stress. The strain resulting from this is irreversible for two successive cycles and not cancelled in the strain difference function defined by Equation 1. From the previous consideration it is clear that the use of the strain difference function has the advantage of emphasising the irreversible component of the measured non-linear strain by eliminating the reversible component. Using the strain difference function, therefore, we can detect a bending point of stress-strain curve to estimate the value of a normal component of in situ stress along the loading axis (see Figure 2).
CASE STUDY CANNINGTON MINE
Deformation Rate Analysis The Deformation Rate Analysis (DRA) method uses the behavior of inelastic strain of the specimen under uniaxial compression to determine the previously applied stress of a specimen. This method is similar to AE method in that both of them utilise the inelastic properties of rocks under compression. Yamamoto et al (1990) experimentally demonstrated that the previous stress can be obtained from a change in the gradient of stress-strain relation under cyclic uniaxial compression tests for a specimen, and then they named the procedure the deformation rate analysis. The gradient changes were not commonly determined in the stress-strain relations obtained by conventional techniques in the case of small previous stresses, because the change might be buried in the larger non-linearity of the stress-strain relation resulting from other sources (eg crack or pore closure).
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FIG 1 - A typical AE count rate for a cored sample, second loading cycle.
Core obtained from an Hollow Inclusion (HI) cell measurement hole was used to establish whether the complete stress tensor could be determined from AE and DRA test data. The conventional HI cell technique had been used at the Cannington Mine to determine the in situ stresses, although this information was not used in any way during the experiment. Nevertheless, the rockmass at Cannington is geologically very complex and different stress magnitudes and orientations were estimated by the HI cell within adjacent locations along a hole axis. Two test sites were analysed for stress measurements using cored rock. The 520 L site was drilled at approximately 521 m below surface with the hole oriented roughly horizontally North-South on the local mine coordinate system. The 605 L site was also drilled horizontally, but bearing East-West and slightly deeper at 596 m below the surface.
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IN SITU STRESS MEASUREMENTS USING ORIENTED CORE
TABLE 2 Summary of estimated stress from AE and DRA methods, 520 L. Drilling direction
Sample name
Stress AE (MPa)
Stress Ave DRA (MPa) Core (MPa)
1
5201C
25
29
1
5201D
33
29
1
5201E
23
28
1
5201S
27
2
5202A
21
22
2
5202B
19
19
2
5202E
22
25
3
5203A
21
25
3
5203B
21
-
3
5203D
21
26
4
5204A
21
24
4
5204B
31
-
4
5204E
26.5
23
5
5205B
27
27
28
21
FIG 2 - A typical strain difference plot for a cored sample.
The AE/DRA measurements were carried out to establish independent stress values to those determined by the HI cell. At least six independent orientations are required in order to determine the six independent components of the stress tensor. Table 1 shows the undercoring orientations chosen for the 520 L measurement site. TABLE 1 Summary of undercoring orientations – 520 L. Sampling direction
Plunge (°)
Bearing (°)
Depth into hole (m)
1
-05
002
10.55
2
85
002
9.61
3
00
272
9.61
4
45
272
9.61
5
37
002
9.61 - 10.55
6
00
320
10.55
The stress estimations from AE and DRA for the 520 L site are shown in Table 2. In most cases, up to three stress measurements were conducted for each method and individual orientation. The input to the computer program used to calculate the full stress tensor was the average of the individual AE and DRA values for each orientation. The stress tensor is six-dimensional (three normal stresses and three shear stresses) and therefore six independent normal stress measurements suffice to determine the full stress tensor with respect to a mine north, south, vertical co-ordinate system. The determination of the stress tensor uses equation (2.14) of Brady and Brown (1985). Table 3 shows the full stress tensor for the 520 L measurement. The principal stresses obtained by an eigenvalue analysis are shown in Table 4. Although the results are round to the nearest integer, it should be noted that all calculations and results were computed using four decimal places. At the 520 Level the orientations of the estimated principal stresses appear to match the localised orebody orientation very well. The main principal stress was found to be parallel to the orebody strike. The intermediate principal stress was parallel to the dip, while the minor principal stress was normal to the orebody. At the 605 Level, the orientation of the major principal stress was normal to the orebody, while the intermediate principal stress was also parallel to the dip. The minor principal stress was found to be parallel to the strike of the orebody. The estimated stress orientations using cored rock for both sites are shown in Figure 3.
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5
5205C
26
27
5
5205D
25
29
6
5206A
20
23
6
5206D
20
24
6
5206E
18
26
23
25
27
22
TABLE 3 Full stress tensor estimation – 520 L. Full stress tensor matrix AE results 27 5 -1 5 21 5 -1 5 20 Full stress tensor matrix DRA results 29 3 -1 3 26 0 -1 0 22 Full stress tensor matrix Combined AE and DRA results 28 4 -2 4 23 3 -2 3 21
The magnitudes and orientations for the principal stress directions obtained from the cored samples were then compared with those obtained using conventional HI cells installed within the same hole. The results for the 520 L measurement site are shown in Figure 4.
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E VILLAESCUSA, M SETO AND G BAIRD
σ
σ σ σ
σ σ σ
σ
σ σ σ
520 Level
605 Level
σ
FIG 3 - Principal stress magnitudes and directions - 520 and 605 Levels.
σ σ σ
σ σ σ σ σ σ σ σ σ FIG 4 - A comparison of principal stress magnitudes and directions - 520 L.
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IN SITU STRESS MEASUREMENTS USING ORIENTED CORE
TABLE 4 Principal stress estimation from cored rock, Cannington Mine 520 L. Stress
Test
Value (MPa)
Bearing (°)
Plunge (°)
σ1
AE
30
216
11
σ2
24
320
48
σ3
14
117
39
31
032
8
σ2
24
300
15
σ3
21
149
73
30
211
02
24
303
46
17
119
44
σ1
σ1 σ2 σ3
DRA
AE and DRA
CONCLUSIONS In situ stress estimations using Acoustic Emission (AE) techniques and Deformation Rate Analysis (DRA) were carried out using rock cores collected from oriented boreholes at the Cannington Mine in Australia. The time lag to testing did not deter the evaluation of the critical in situ stress condition. The rock core specimens recollected the in situ rock stresses reasonably well compared with the standard overcoring methods.
ACKNOWLEDGEMENTS The authors wish to thank the management of the Cannington Mine for their permission to publish the paper and the National Institute for Resources and the Environment (NIRE), Japan for the use of its facilities to carry out the stress measurements.
REFERENCES Brady, B H G and Brown, E T, 1985. Rock Mechanics for Underground Mining, 571 p (Allen and Unwin: London). Holcomb, D J and Stevens J L, 1980. The reversible Griffith crack: A variable model for dilatancy, J Geophysics Res, 85:7101-7107.
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Kaiser, J, 1953. Erkenntnisse unde Folgerungen aus der Messung von Gerauschen bei Zungbeanspruchung von metallischen Werkstoffen, Archiv Fur das Eisenhuttenwasen, pp 43-45. Kurita, K and Fujii, N, 1979. Stress memory of crystalline rocks in acoustic emission, Geophys Res Lett, 6(1):9-12. Seto, M and Villaescusa, E, 1998. In-situ stress determination by Acoustic Emission technique in McArthur River mine cores, in Proceedings 8th Australia New Zealand Conference on Geomechanics, Hobart, 2:929-934. Seto, M, Utagawa, M and Katsuyama, K, 1995. The relation between the variation of AE hypocenters and the Kaiser effect of Shirahama sandstone, in Proceedings 8th International Congress on Rock Mechanics, Vol 1, pp 201-205, Tokyo, Japan. Seto, M, Nag, D K and Vutukuri V S, 1999. In-situ rock stress measurement from rock cores using the acoustic emission and deformation rate analysis, Geotechnical and Geological Engineering, 17(3-4):1-26. Yamamoto, K, Kuwahara, Y, Kato, N and Hirasawa, T, 1990. Deformation rate analysis: a new method for in situ stress examination from inelastic deformation of rock samples under uni-axial compressions, Tohoku Geophys J, 33:127-147.
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Terratec Universal Raiseborer — Raisebore Machine Technology C Paterson1 INTRODUCTION Drilling underground can be a hazardous undertaking in many circumstances. Raiseboring and boxhole boring underground could form an integral part of the block caving production methodology and getting the right mix of technical capability, ergonomics and economical viability has been the goal of many machine manufacturers. With the introduction of the Terratec Universal Boring Machine concept, flexibility of raiseboring, downreaming and boxholing has been enhanced by the introduction of a single machine that will perform all three processes. In addition to its technical flexibility, additional safety features have been introduced that will eliminate many sources of injury by mechanising most of the operational procedures. Together with the one-man remote operation, this machine design addresses and solves many of the safety issues. This paper outlines the machine concept, details its safety features and shows the techno-economic advantages for use in the block caving environment. Block caving is a distinct caving method applied mostly to large massive orebodies because of the inherent low costs and high production capabilities. Areas of sufficient size are mined by undercutting so that the mass above will cave naturally. Extraction of the caved ore at the bottom of the ore column causes the caving action to continue upward until all of the ore above the undercut level is broken into sizes suitable for handling (Figure 1). Raise boring presents a significant level of specialist work in a block cave mine. Finger raises, blind slot raises, transfer raises, back-over branches, draw raises, ore passes, ventilation raises all provide ample application for a flexible raisebore machine. As the raises are at regular intervals throughout the levels at different phases of the mine development (Figure 2), it would appear a prudent investment decision to purchase a flexible, cost-effective and safe raisebore machine to perform these tasks.
RAISEBORE MACHINE DEVELOPMENT Raisebore machine technology today is a tried and proven method and very little has changed in the operation of the machinery. In the last few years high powered, compact and semi automatic machinery has been developed but Terratec felt there was room for even more. What was needed was a compact, multi-purpose machine that could cut rock effectively, inexpensively and most importantly safely. This prompted Terratec engineers to come up with the Universal Borer machine, which combines up reaming, down reaming and boxholing operations within the one machine (Figure 3). The operations can be defined as follows.
Up–reaming (traditional raiseboring) Top and bottom access is required for this process. A pilot hole is drilled to the lower access level and the bit is removed. A larger reamer head is subsequently attached and the head is pulled upwards rotating and cutting the rock to the final diameter as it progresses. Cuttings fall downwards (Figure 3).
1.
General Manager - Sales and Marketing, Terratec Asia Pacific Pty Ltd, PO Box 83, Blackmans Bay Tas 7052. E-mail: [email protected]
MassMin 2000
FIG 1 - Block cave isometric layout.
Down reaming As mentioned in the section above however, the drill string is retrieved once the pilot hole is finished and the head is attached at the top of the raise and pushed downwards by the machine to produce the final diameter hole. The cuttings fall downwards through the pilot hole (Figure 3).
Boxholing With the two options above, top and bottom access is required. In this process however, the pilot hole is drilled from the bottom to the top, the drill string is tripped back out and the head is attached and pushed upwards to drill the final hole. The boxholing procedure can also be done blind. It should be noted here that blind boring can cause steering and alignment problems as well as increased centre cutter wear. These problems don’t arise when boxholing with a pilot hole (Figures 6 and 7). The above procedures represent the non-entry method of vertical development along with long hole winzing and inverse raising. Entry methods include Alimak raise climbers, cage raising and ladder raising.
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FIG 2 - Block cave layout-Andina, Codelco, Chile.
method is the structural integrity of the completed raise and the improved ventilation flow or rock flow where used as a pass. Safety was a major criteria during the designing phase of this machine as the in-direct cost-savings from elimination of injuries, as well as more importantly saving fatal accidents from occurring, are substantial over the long-term.
TERRATEC UB1000 FEATURES AND BENEFITS As mentioned previously the machine is capable of three operations and can be changed from standard up reaming/down reaming to boxholing mode quickly. The main features are as follows.
Quick gearbox interchange Fast on-site change of machine application. The gearbox is mounted on bolts that can be dislodged, the gearbox turned 180 degrees and then rebolted to provide the set-up for the boxholing operation from the down reaming operation for example. The procedure should only take approximately half a shift with two experienced fitters.
Integrated torque/thrust cylinder These cylinders are designed with balanced thrust capabilities and the same pressure generates equal thrust force for both up and down operation. The four- thrust column configuration also results in clear space all round to provide the operator with unrestricted views at all times during the operation (Figure 4).
All purpose rod handler FIG 3 - Down-up-reaming/boxholing procedures.
It goes without saying that the non-entry methods of vertical development present the lowest levels of hazards to personnel and the surrounding mine infrastructure and offer the most readily controlled spectrum of risks. Raise drilling is well established as the regular method and a further benefit of this
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The rod handler is configured to operate in all applications. The machine is also capable of being operated by one person only therefore the traditional ways of handling rods via monorail and chain block had to be eliminated. A great deal of thought was applied to the problems of the different applications and the need for the bottom hole assembly (BHA) to be lowered into the worktable. To overcome these problems, a high power sophisticated rod handler was developed which produced a system with a larger handling capacity, relative to that of the largest raiseborers (Figure 5).
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TERRATEC UNIVERSAL RAISEBORER - RAISEBORE MACHINE TECHNOLOGY
manual work is required above head height. It also provides an extra safety block against falling debris (Figure 6).
FIG 4 - UB1000 schematic.
FIG 6 - Powered worktable.
Powered wrench tool This again is another safety feature of the machine while in the boxholing mode. The operator can make-up/break-out without putting himself in direct danger. It can all be done from the remote control panel and is hydraulically actuated. Manual operated wrenching is provided for the other operations.
Adjustable muck collector The muck collector is designed in two halves and can be opened by hydraulically actuated cylinders to allow the passage of the reamer and stabilisers. Water sprays are incorporated in the top end to act as dust suppression. The collector also incorporates a muck chute which deflects the debris away from the machine at any location left and right of the machine. This muck collector can be removed as one unit by removing four expansion pins (Figure 7).
Elastomer seal around pipe The seal prevents sludge falling into the gearbox in the boxhole mode, thus preventing possible premature failure of the gearbox.
Hollow shaft motor Allows direct flushing of the cuttings from the pilot hole and allows unrestricted flow of the bailing media. FIG 5- All purpose rod handler.
Terratec hydraulic digital variable speed drive/power pack Powered worktable doors Hydraulically powered worktable doors are located at the bottom and top of the machine to facilitate the support of the drill string and reamer weight. When used in the boxholing method, no
MassMin 2000
This system is robust, simple, quiet, reliable, cost-effective and helps to protect the machine and the drill string. The power pack design is tried and tested and is a feature of all Terratec raise drills. It is a low maintenance unit which is also fully PLC
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FIG 8 - Hydraulic power pack.
• worktable as added protection against falling debris; • hydraulically powered wrench; • integrated thrust/torque cylinders providing all round vision for the operator;
• one man operation throughout cycle; • fully hydraulic drive; • automatic shutdown in blocky conditions or system failure;
FIG 7 - UB1000.
(Programmable Logic Controller) controlled. This is also necessary for a one-man operation. It is possible to pre-program the machine to operate the tedious task of monitoring and controlling the reaming cycle. This feature is further employed at the end of the shift to allow the machine to continue the reaming process unattended, to the completion of the individual drill rod currently in the machine. PLC control, in conjunction with the adoption of the hydrostatic transmission, leads to a reliable method of controlling inertial forces imparted to the drill string thus reducing the phenomenon of drill string wind up during inattentive operation or blocky ground conditions. This feature allows relatively inexperienced operators to rapidly gain confidence in adverse ground conditions. One client claims this to be the ‘best hydrostatic transmission in the world’ (Figure 8).
Specifications The UB 1000 was based on general industry raise diameter requirements in the slot raises, fingers, draw raises etc and reflect a preliminary design. The concept can be modified slightly depending on the given circumstances. The UB 1000 can pull a standard raise of 1500 mm, can downream and boxhole through the upper and lower worktables at 1060 mm. The following specifications show the standard UB 1000 make up (Table 1). The plan shows the UB 1000 in the various modes (Figures 9 and 10).
COST-EFFECTIVENESS OF THE UB1000
Compact/modular design The compact design gives ease of transport and a low profile in relation to the drill string length. It also facilitates uncomplicated assembly and dismantling as well as ease of maintenance and repair.
Safety features As mentioned previously the machine incorporates a number of safety features which will make an active contribution to a reduction in many of the general injuries associated with this type of work and has the potential to eliminate fatalities. The main features are as follows:
• mechanised pipe-loading of all components; • no man access required above two metres height; • all functions above two metres height are hydraulically powered;
• muck collector as safety helmet against falling debris; • elastomer sealed drill string; • hydraulically actuated worktable doors;
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and
• counterbalance cylinders maintain position when flow is cut.
The UB 1000 is a piece of capital equipment and as with any piece of equipment it must prove to be of cost benefit to the mine operator. Most mines have tended to use manual means of excavation which are both time consuming and dangerous. It should follow that the quicker the raises can be finished and the blasting of the block cave starts, the faster the ore can be mined and the quicker the revenue can flow from the mine.
DIRECT MECHANICAL SAVINGS These are the predicted direct cost savings accrued by using this type of equipment and in particular the Terratec designed raiseborer series. The UB, although a new design model incorporates many of the Terratec tried and tested features which have improved reliability and performance and reduced maintenance and spare parts costs considerably. Similar Terratec machines working to date have effectively used less than $A10 000.00/year in spares on average. This coupled with increased performance and machine utilisation contributes significantly to the lowering of the cost per metre of drilling.
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TERRATEC UNIVERSAL RAISEBORER - RAISEBORE MACHINE TECHNOLOGY
FIG 9 - UB1000 down/up reaming site layout.
TABLE 1 UB 1000 specifications. Machine Specifications of the UB1000 Nominal Hole Diameter
Box Hole
Down Ream Back Ream
3.5 ft
1.06 m
3.5 ft
1.06 m
5 ft
1.5 m
8650 ft lb
11 700 Nm
Reaming
45 000 ft lb
61 000 Nm
Make-up
49 500 ft lb
67 000 Nm
Break Out
54 000 ft lb
73 000 Nm
Torque
Thrust RPM
Pilot Drilling
Down
250 000 lb
1100 kN
Up
250 000 lb
1100 kN
Pilot Drilling
0 - 80 Stepped
Reaming
0 - 16 (Stepped)
Power
162 HP o
Derrick Dip Angle
90 - 60
Maximum Drill Rod Length (S/S)
121 kW
o
48”
1219 mm
Nominal Down Drilled/Back Reamed Hole Depth
650 ft
200 m
Nominal Box Hole Length
280 ft
85 m
10”
254 mm
Standard Drill String Electric Voltage
440 V, 50 Hz, 3 Ph 120 V AC and 24 V DC Control
Electrical Power
115 KVa
Derrick Extended Height at 90o Dip
(Conventional Configuration)
114 - 1/2”
3670 mm
Derrick Extended Height at 90o Dip
(Box Hole Configuration)
173”
4392 mm
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FIG 10 - UB1000 boxholing site layout.
DIRECT OPERATIONAL SAVINGS These predicted savings include the increase in utilisation and performance as well as the decrease in downtime due to the PLC controlled system allowing faster uninterrupted changeover time. This not withstanding, the one man operation also reduces labour expenditure. Better utilisation of experienced personnel is the result. Direct operational savings are also attributed to the methodology of raise boring versus drill and blast. The extra ventilation, the reduced working time, the increase in size of shift teams, the extra ground support required and the speed of production all contribute significantly in the overall cost per metre drilled as well as the cost per tonne ore mined.
INDIRECT SAVINGS These are attributed to the inherent safety features of the methodology as well as the new safety features of the machine. It is predicted that the machine safety features will eliminate approximately 90 per cent of all the niggling injuries that are associated with this type of work. Owing to the nature of the working environment these incidents need to be reported and in most cases treated. This results in downtime, insurance claims, sick leave, etc. With the potential to be eliminated to a large extent, the cost benefits show themselves in machine utilisation, progress, reduced downtime, fewer claims and possibly a reduction in insurance premiums. This could be a real spin-off in that the insurance companies could reward the mining companies
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for utilising safe working practices and state of the art machinery incorporating the latest in safety features. We anticipate the regulatory authorities will soon prohibit manual excavation of ladder rises in Australia.
CONCLUSION Block caving mining presents a great opportunity to utilise the UB 1000 technology and reap the benefits mentioned above. Although the cost-savings are difficult for Terratec to quantify due to our position as an equipment supplier and not operator; they can be quantified by the people doing this every day. We believe the cost benefits from an operational sense as well as the safety issues deserve real consideration by the mining community out there, planning capital machine purchases. The UB 1000 provides a flexible cost-effective investment and is also a machine which sends the right message to the miners at the face; mine management cares.
REFERENCES Peach, T, Mulder, G and Adams, A, 1998. Introduction of Down Reaming in Australia, in Proceeding 7th Underground Operators Conference, pp273-276 (The Australasian Institute of Mining and Metallurgy: Melbourne). Hartman, H L, 1999. Mining Engineering Handbook. Torlack, J, 1990. Department of Minerals & Energy-Western Australia, Safety Bulletin, No 3. Torlack, J, 1998. Department of Minerals & Energy-Western Australia, Safety Bulletin, No 39.
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The Potential Use of Foam Technology in Underground Backfilling and Surface Tailings Disposal A J S Spearing1, D Millette2 and F Gay2 ABSTRACT Surface tailings are finely ground waste materials produced in all mines after the mineral or metal has been removed from the metallurgical process. Current methods for disposal of tailings involve constructing dams, dumping into existing waterways, and depositing as backfill underground. Each method has its drawbacks in terms of cost, level of complexity, safety and environmental concerns. Backfilling is the process whereby waste material (generally metallurgical tailings, prepared aggregates or a combination) is placed from the surface back underground into open mine workings via conveyor belts, hydraulic or pneumatic pipes or gravity. In the backfill process common difficulties are expensive placement costs, equipment wear and the ability to meet early and long-term strength requirements. Traditional tailings transportation in pipes involves the use of water as the transport medium. Foam technology involves the introduction of micro-air bubbles that replace the water as the main transport medium. This air can be removed after placement if necessary, by the addition of a defoaming agent. This could produce an under-saturated stable cake. Foam technology has the potential to revolutionise surface tailings disposal and backfilling by providing a less expensive and environmentally safe alternative to those just described.
SURFACE TAILINGS DISPOSAL Introduction to tailings disposal All mines (except aggregate or sand quarries), whether surface or underground, produce a finely ground waste material called tailings, after the mineral or metal has been removed during the metallurgical process. These waste materials can be disposed of in the following ways:
• into the sea or a lake (this method is now not generally favoured due to environmental considerations);
• into a river behind a suitable dam wall (this can be a costly method and can considerations);
also
has
adverse
environmental
• on to a specifically constructed tailings dam on surface; and • placing the waste back underground as a backfill material (as mentioned later this method cannot be used to dispose of all the tailings produced, and some surface disposal is still required).
Planning and engineering considerations for tailings dam design There are many considerations to be taken when planning a tailings basin especially with a new mining operation or in the case of major expansion. Some of the considerations for planning a tailings dam and basin are:
• the life and size of the facility, • water re-use needs,
1.
2.
Technical Director – UGC Americas – MB Inc, 23700 Chagrin Boulevard, Cleveland Ohio 44122, USA. E-mail: [email protected] Senior Scientist – UGC Americas – MB Inc, 23700 Chagrin Boulevard, Cleveland Ohio 44122, USA.
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• • • •
environmental issues, land availability and topography, cost of construction and operation, and reclamation.
Some of the considerations for dam design once the location is selected are:
• • • • • • •
spillway location, type, and sizes, water reclamation system, availability of materials for dam construction, controlling seepage (dam failure and pollution), foundations, abutment, and operations.
Methods for dam construction There are several methods of tailings dam construction that are used. The selected method can be a function of:
• • • •
terrain, operations, availability of materials, and cost. The common methods include the following.
Conventional method In this method a conventional earth dam is constructed in a suitable valley or ravine.
Downstream method As the volume of tailings stored increases with time, the retaining wall height is increased and the centre of gravity of the wall moves away from the centre of the dam (ie down dip).
Upstream method This is the opposite of the downstream method as the centreline of the wall moves into the dam as the wall height increases.
Tailings dam operation The tailings dam operators often control the tailings dam design, with respect to the method of dyke construction and availability of materials. Some of the considerations for tailings dam operations are:
• • • •
water pond location and elevation (within the dam), spillway, and tailings deposition. Reclamation
The planning for the design and operation of a tailings basin should consider the reclamation to the deposition of tailings, location of spillways, and final drainage patterns.
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In recent years, the costs associated with dam rehabilitation have increased dramatically due mainly to the environmental lobby in first world countries. A single large tailings dam at the now defunct Elliot Lake Mine cost over US$100 M to rehabilitate at the end of the life of the mine.
Aspects that need research, before foaming technology can be considered include:
• a new conceptual transportation and discharge design with foam;
• the rheology of the material during piped transportation to the deposition site; and
The problems with current tailings dams The problems with conventional tailings dams include the following:
• the inherent safety risk of liquefaction such as occurred in Spain and South Africa, causing severe property damage (and loss of life in the South African accident at Merriespruit Mine);
• a long-term rain water disposal system, although the system used by Kidd Creek may also be acceptable if foam is used. Figure 1 compares the conventional surface tailings method with the high density method (that foam could be a key element in).
UNDERGROUND BACKFILLING
• serious water losses on the dams in regions where water is a scarce commodity (such as in Australia and South Africa); and
• the large area needed for the dam and the cost of eventual reclamation. There is a move within the mining industry (Robinsky, 1999) to change from conventional tailings dam disposal to dry stacking (or thickened tailings disposal as it is also called). This dry stacking method involves the densification of the tailings usually using some form of high rate thickener, prior to disposal. This produces a near homogeneous material with some self-supporting capability. The potential advantages of such a method are:
• The placed tailings consolidate much faster and the chance of liquefaction is significantly reduced. The need for a settling pond is thus eliminated.
• More process reagents can be recovered at the metallurgical plant.
• The reduced water in the tailings reduces the seepage and hence also the pollution potential.
• Water recovery is much higher and this could become critical in certain dryer areas (Spearing, 1999).
• Rehabilitation can be concurrent with the stack operation. This method is not new and has been successfully used by the Kidd Creek Division of Falconbridge Ltd in Canada since 1974, when the first tailings thickener was installed (Sudbury, 1999). Other mines have (or are) implementing the method including:
• • • •
Ekati Diamond Mine in Canada, Bulyanhulu Gold Mine in Tanzania,
• • • • •
reduced tailings on surface, reduced surface subsidence, increased underground extraction, reduced rockburst (seismic) damage, improved support for mining excavations and/or working surfaces, and
• increased worker safety. Backfill transportation There are four main transportation methods (or combinations) used to move the fill from the point of preparation (generally surface) to the underground workings:
• Hydraulic via pipelines, with the fill as a slurry (low density) or paste (high density). This is the most common method.
• Mechanical via conveyor belts or trucks. • Pneumatic via pipelines as a dry (or damped) material. This method is seldom used today because of dust concerns, high energy requirements and high wear rates.
• Gravity via raises and rock passes. Criteria and requirements
Kubaka Gold Mine in Russia, and Hindustan Copper Ltd in India.
Dry stacking has the potential to revolutionise surface tailing disposal by removing the need for the traditional drainage pond system for water recovery. The concept does however require a totally different construction design. Foam technology can benefit this disposal method by:
• reducing the cost of transporting this higher density tailings product (and this is a significant proportion of the costs); and
• producing almost any desired final beach angle on the dry stack after deposition. This would be achieved in practice by altering the dosage of defoamer at the discharge nozzle. Many mines are already considering high density (paste) disposal for similar considerations, but if the foam concept works, it will be more effective. The international mining industry is so concerning with surface tailings disposal that two conferences have already been held in Canada and Australia and another is planned in March/April 2001 in South Africa (see the references at the end of the paper).
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Introduction to backfilling Backfilling is the process whereby waste material (generally metallurgical tailings, prepared aggregate or a combination) is placed back underground into open mine workings. This has many potential advantages for typical underground mining applications, such as:
The criteria and requirements for backfilling vary, depending on the site’s specific requirements. Following are some of the more common and important criteria and requirements:
• • • • • • •
the fill should be placed at the lowest possible cost; the risk of fill failure (eg liquefaction) must be minimised; early strength development needs to be adequate; long-term strength should be sustainable; delivery volumes must be reliable and adequate; after placement, dimensional stability must be achieved; and segregation should be minimised.
The backfill challenge and opportunity Many mines, especially in North America and Australia are moving towards high density (paste) fill. This material is typically over 78 per cent solids by weight as opposed to hydraulic fills that are usually a maximum of 70 per cent solids by weight. The main reasons for this are:
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COMPARISON BETWEEN CONVENTIONAL AND THICKENED “PASTE” SURFACE TAILINGS DAM CONSTRUCTION
NEW TECHNOLOGY THICKENED
CONVENTIONAL Tailings discharged from perimeter dam using spigots
Tailings discharge to be commenced from tower at north end, then proceed on ramp south
Reclamation can begin sooner from the north side
Natural slope
Water diversion ditch
Settling pond
Ditch
Pond
Low embankment
Recycle pumphouse
Perimeter dam to be built progressively using cycloned coarse tailings fraction
Recycle pumphouse
(After E. Robinsky, 1999)
FIG 1.
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• Total tailings can be used, minimising the problems associated with the surface disposal of the remaining material. All tailings cannot be backfilled due to the different density between intact rock and broken rock. At best only about 60 per cent can be returned underground, leaving 40 per cent (at least) for surface disposal.
• After placement, the high density tailings produces little or no bleed water resulting in reduced post filling shrinkage, reduced pumping (as the exuded water is minimised) from underground and more efficient cement hydration.
• Faster mining cycle because stopes adjacent to filling areas can be mined sooner after the fill placement.
• Reduced pipe wear if the backfill is transported underground in pipelines. The paste system is not without problems, however, such as:
• more difficult rheology as the paste has higher pressure losses per unit length of pipe as compared to hydraulic (low density) backfills;
• a higher beach angle when placed in a stope (ie the fill is not self levelling); and
Canadian Gold mine This is a medium sized gold mine operating at depths of approximately 1350 m. The mine has been using hydraulic tailings backfill but are consolidating their milling operations with another mine owned by the same company. With the closure of their milling operations, their tailings backfill source is also disappearing. The decision has therefore been made to investigate the use of alluvial sand as a substitute for the tailings. In order to keep the hydraulic sand fill from settling during transport, the fill was designed at 68 per cent solids. To enable attaining the required backfill strength of 1.35 MPa in the stopes, it will be necessary to add up to six per cent cement as a good portion off the cement will leave the stope in the decant water. By using a pre-generated foam (using a specific foam generator), MBT was able to attain better flow properties with the same sand at 90 per cent solids, and take a 50 per cent reduction in cement to attain the target strength. The mine is planning to conduct a full-scale field trial this year.
US Soda Ash mine
• relatively high operational and capital cost (on most mines the backfill operational cost of between US$3.00 to US$6.00 typically is the single highest mining input cost). MBT is already involved in selling admixtures for use in backfilling, and although cost-effective, a far better solution seems to exist for some applications. This solution is the use of foam technology as it can produce a backfill with the combined benefits of both a paste and a hydraulic fill because:
• for transportation the foam produces a ‘pseudo’ low density slurry making pipeline pressures low and reducing the pipe wear rates; and
• after placement the foam is destroyed producing a paste like material with desirable physical properties
RESEARCH FINDINGS Testing at MBT (Spearing, 2000) over the last year has enabled the company to establish certain advantages of foamed fill as follows:
• foam enables/improves the transport properties of relatively dry, coarse materials;
• foam reduces and in most cases eliminates segregation of the
The tailings of this mine are very fine and form a paste with a 125 mm slump at 37 per cent solids. During pumping trials at the mine site, the addition of a powdered foaming agent at the pump reduced pump pressures to 75 per cent of the non-foamed material. Had it been possible to add the foaming agent at the mixer, the pressures may have been even further reduced. These trials were to investigate the future potential surface deposition of the tailings material.
UK Potash mine A potash operation in Northern England was using subaqueous deposition for tailings disposal but due to increasingly more stringent environmental regulations are disposing of a large portion in the mined out underground areas. They are not using any cement in this backfill as it is not intended for support but as a means of disposal. The current transport of the tailings underground requires a maximum pulp density of 62 per cent solids. This leaves a great deal of troublesome brine to be returned to surface. During laboratory trials, the solids content of the backfill was increased to 75 per cent and using a pre-generated foam that was injected into the mixer, and rheological properties were obtained that surpassed those of the untreated 62 per cent solids material.
solids/water in high density backfills during transport;
• foam lubricates plug flow by forming an air annulus between the plug and the pipe wall;
• foamed material will compress under pressure but expand upon release of pressure but remains unproven in a full-scale pipeline of significant length;
• there is a significant decrease in compressive strength of backfill material that has not been subjected to de-foaming;
• the amount of foam generated during mixing increases with the degree of gap grading or void ratio of the material being foamed; and
• the addition of foam to material that is being pumped, regardless of particle size gradation, significantly decreases the pumping pressure.
CASE STUDIES
ON-GOING RESEARCH Research plans have been drawn up to investigate transport and placement of foamed tailings materials both for backfill and surface deposition. For surface deposition, roller compaction of the foamed material is thought to be a method of de-foaming the material. Also, tailings that have been treated with a foaming agent often tend to consolidate to a higher final density. One of the major foamed fill research endeavours now taking place at the MBT Research Centre in the US is the investigation of using foamed tailings backfill in a Canadian underhand mining operation. The mine’s current tailings backfill has a high bulk density. It is thought that the bulk density of the fill can be reduced to between 30 and 40 per cent of the current fill using a pre-generated foam. As the fill is only used to prevent sloughing of the stope walls above the mining, a lighter density material would be easier to support.
The case studies using foam technology are limited at present but tend to illustrate the potential for this technology.
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CONCLUSIONS
REFERENCES
From the limited experience to-date it appears that foam technology holds potential for the high density backfilling and surface tailings disposal on mines. Future research will establish the final viability and application scope for the technology. Costs will be an overriding factor.
Robinsky, E I, 1999. Thickened Tailings Disposal in the Mining Industry. E I Robinsky Assoc Ltd (ISBN 0-9686113-0-3). Spearing, A J S, 1999. Tailings disposal practice in South Africa. Paste Technology for Thickened Tailings – Mine Tailings Learning Seminar, University of Alberta, Edmonton, Canada. Spearing, A J S, 2000. The potential for foam technology in mining. Paste Technology 2000 – Australian Centre for Geomechanics – Perth, Australia. Sudbury, M, 1999. Kidd Thickened Tailing Operation. Paste Technology for Thickened Tailings – Mine Tailings Learning Seminar. University of Alberta, Edmonton, Canada.
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The Design, Testing and Application of Ground Support Membranes for Use in Underground Mines A J S Spearing1 and J Champa2 ABSTRACT The mining industry internationally seems on the verge of a major underground support revolution as suppliers, mines, consultants and researchers develop, test and use thin support (structural) membranes (or superskins as they are commonly called) mainly as a replacement for screen/wire mesh. Thick membrane support in the form of shotcrete (sprayed concrete) has been used since the 1950s in underground mines. Shotcrete, in mines, is generally applied at a thickness between 25 and 100 mm depending on the need. Thin membranes are however applied at a typical thickness between 2 and 6 mm. Certain thin membranes have been available for several years (Tannant et al, 1999), but the intense interest in such supports has only grown in the last few years. The interest lies mainly in the potential to safely reduce the mining cycle time and hence improve productivity. There is however considerable confusion concerning such supports, and this paper attempts to clarify certain aspects and dispel certain misconceptions and myths about thin support membranes.
SUPPORT MEMBRANE NEEDS The performance of any type of membrane support, including shotcrete, is governed by the following main physical parameters:
• bond strength (adhesion) to the various substrates (rock types),
• • • • •
tensile strength, elongation, tear strength, durability, and the effects of humidity and temperatures in mines (generally site specific).
The absence of compressive strength from the list is intentional. It is a property that is more important to thick and rigid membranes (ie sprayed concrete/ shotcrete). It can be easily tested and is used as a quality control test in shotcrete applications. It cannot be readily tested or yield meaningful results for thin support membranes (with relatively high elongation). The rate of gain and not only the ultimate value of the tensile strength, adhesion and elongation is important to the performance of thin structural membranes. A rapid gain of properties means that a more active type of support is applied to the rock earlier. The other critical parameters are the health and safety considerations of the material, before, during and after application:
• flammability, • toxicity, • chemical resistance,
Ease of application, including the ability to obtain the desired application thickness is also an important operational parameter. The cleaning of the equipment after use is also a consideration, unless water only can be used, because solvents tend to create additional problems during handling and usage. The physical properties of membranes are constrained by the chemistry of the raw materials available. As an example, high tensile strength and high elongation are both desirable, but generally this becomes a trade-off, as higher elongation generally results in lower tensile strengths and vice versa.
SUPPORT MEMBRANE APPLICATIONS The main advantages of support membranes, over other similar types of support such as meshing and lacing or straps are that:
• • • • • • •
reactive loads are generated with far less rock movement; high application productivity is achieved; remote application is possible (therefore safer); they are not prone to corrosion; they are easier to keep concurrent with the face advance; they are less prone to blasting damage; and weathering of the rock is virtually eliminated.
In addition, most thin membranes have the following advantages:
• less costly application equipment; • low volumes to transport and apply (ie easy logistics); and • some can be shotcreted over at a later and more convenient time, if required. requirement.
This
is
frequently
an
important
The above application advantages influence the thin membrane uses underground. Typical uses underground include:
• face support during tunnel development (both blasted or excavated);
• • • •
stope gully support; pillar reinforcement; sealing rock prone to weathering; as a replacement for meshing geotechnically approved; and
and
lacing,
where
• as a support component to help reduce rockburst damage.
1.
Head International Mining, Underground Construction Group – MBT International, 23700 Chagrin Boulevard, Cleveland Ohio 44122, USA. E-mail: [email protected]
2.
Staff Scientist, Underground Construction Group – MB Inc, 23700 Chagrin Boulevard, Cleveland Ohio 44122, USA.
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• handling and application requirements, • safe disposal of unused product, and • possible adverse effects in the metallurgical process.
Thin membranes are not generally a replacement for quality shotcrete where thickness’ of 25 mm or more are required. The combination however offers good potential in severe rock stability areas, as the membrane can give rapid support, and the shotcrete (generally fibre reinforced) provides the high compressive strength and durability. For this to be practical, the shotcrete needs to have good adhesion to the specific membrane, and few membranes are suitable for this.
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on the composite material. These tests can show the interaction between the various interrelated membrane properties and the rockbolts.
There are two main types of thin support membranes:
• one component membranes typically applied using a single piston pump; and
• two component membranes typically applied using a two piston pump.
One component systems One component systems tend to be only slowly reactive and therefore relatively slow in the development of the desired properties. The strength development generally requires moisture loss after the membrane placement. To help accelerate the cure, a cementitious addition is frequently made which helps effectively remove the water. Generally the one component systems consist of a liquid and a solid constituent that are mixed together before being pumped. The sprayed material begins to gain significant strength after typically about four hours. They are typically made from polybutadiene styrene dispersions (the liquid) with some form of cementitious addition (the powder). The powder addition tends to accelerate the strength gain and provide a means to maintain the sprayed thickness on the substrate, prior to the material setting and gaining strength.
Two component systems Two component systems tend to be highly reactive and therefore gain strength rapidly. They are typically made from polyurethanes, polyureas or acrylates. Some form of in-line mixing system is required at the spray nozzle to mix the two components. The sprayed material begins to gain significant strength in less than an hour. Being chemically reactive systems, appropriate safety and health systems must be in place, taking particular care if isocyanates are used (generally present in polyurethane and ployurea systems).
Typical performance Table 1 below gives the typical performance that can be expected from the one or two component systems:
TABLE 1 Typical properties of one and two component pump membrane systems. One component
Property
Two component
1.0 at 4 hours
Tensile strength (MPa)
1.0 at 30 minutes
20 at 4 hours
Elongation (%)
+ 50 at 30 minutes
0.5 at 4 hours
Adhesion (MPa)
0.5 at 30 minutes
TESTING STRUCTURAL MEMBRANES Any testing program must be simple, repeatable, cost-effective and at least provide qualitative comparisons with other similar products. The program should at least include testing the main parameters of tensile strength, elongation, adhesion and tear strength. Small-scale tests are usually performed on the support membrane itself, whereas large-scale tests are usually carried out
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Small-scale laboratory testing Standard tests that could be relevant, for laboratory scale results include:
• • • • • • • • •
ASTM D624-98 for tear strength; ASTM D412 for tensile strength and elongation; ASTM D4541 can be used for adhesion; ASTM E1619-95 for toxicity; ASTM E84-99 and IEC 707 for flammability; ASTM 162 for flame spread; NES 713 for smoke toxicity; ASTM C827 for water absorption; and ASTM D3045 for accelerated ageing.
Larger scale physical property testing Other non-standard tests that are being considered in various parts of the world tend to test the combination and interaction of various physical properties. The object of all such tests is to try to quantify the expected in situ performance by testing the interaction of the adhesion, tensile strength and elongation.
The MBT Membrane Displacement Test (Attiogbe and Ohler, 1999) This test is designed to provide load and displacement data on the performance of a support membrane, but is still small enough so that it can be undertaken easily in a laboratory with no major lifting equipment needed. The test is capable of both short and longer term (ie creep) performance evaluation. For repeatability considerations, pre-cast concrete patio slabs are used. These slabs are commercially available and are quite dense (relative to normal cast-in-place concrete) with a slightly textured finish on one surface.
Test apparatus The main components of the test frame (shown in Figure 1) are a concrete slab with dimensions of 610 x 610 x 51 mm. It is supported on four concrete cylinders of 102 mm diameter by 204 mm height, a loading plunger, and a sleeve to hold the plunger in place during frame set-up.
Membrane application The membrane is applied to the surface of the concrete patio slab either by hand (where a specific thickness is required) or by spraying. The membrane is applied over the entire surface of the slab except for the corner surfaces where the support cylinders are placed. Care should be taken to apply the material to a uniform thickness. Plastic strips of the desired membrane test thickness are placed at the edges of the application area to serve as guides. The hand applied or sprayed material is levelled to the thickness of the plastic. The thickness of the membrane is measured at several locations and the average thickness is recorded. The membrane is left to cure for the desired period of time, prior to testing.
Results The load and displacement data are recorded throughout the test. Typical results are shown in Figure 2. The first load peak is mainly a function of the adhesion of the membrane to the
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FIG 1 - The MBT test arrangement for thin membranes.
FIG 2 - Masterseal 840 R01I at a 3 mm thickness.
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FIG 3 - Block paving layout for the test.
1.2m
FIG 4 - The pull test frame.
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FIG 5 - Load displacement curve for Masterseal 840 R01 after one hour.
substrate. Subsequent to this, the membrane starts to peel and elongate, until it ultimately fails. Failure is a function of the interaction between the tensile strength, elongation and adhesion. This test method has the advantage that a certain thickness of membrane can be applied for a test (by pouring the mixed material onto the slab and finishing it with a scraper against a lip of the desirable height).
Membrane application
The INCO/GRC Membrane Test Method (Tannant, 1997)
Results
This test method was modified from an earlier method used by Tannant (1995) to characterise the load versus deformation behaviour of mesh/screen used underground.
The membrane is sprayed onto the concrete slabs from above and left to cure for the predetermined period. An hour is a good practical time after which to test reactive membranes, and between four and eight hours seems suitable for most unreactive membranes.
The load displacement data is collected during the test and then reproduced on a graph as shown in Figure 5. The results show the total load carrying capacity of the membrane.
The CANMET Membrane Test Test apparatus A 300 mm square steel plate is placed in the centre of an interlocking series of hexagonal concrete paving blocks (50 mm thick), as shown in Figure 3. A test frame, as shown in Figure 4 is lifted onto the paving blocks that have been sprayed with the support membrane, and a lifting load is applied to the base plate until the membrane has failed. The method appears repeatable and gives results that can be compared to the mesh test results.
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Test apparatus In this test, a 1.1 m square box, 0.3 m high is filled to the top, in three equal stages with specified aggregate and tamped after each stage. The top layer can be larger flat stone or paving blocks, and the gaps can be up to 25 mm between these blocks.
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Membrane application
Test apparatus
The membrane is applied over the compacted aggregate, but this process is difficult due to the varying gaps between the aggregate pieces. This results in a non-uniform membrane application thickness that makes repeatability difficult After a predetermined curing time, the box is inverted and placed under a compression machine and loaded to destruction. The load is applied to the centre of the box by a 0.56 by 0.30 m platen. The advantage of this test is that it is very useful as a tool to convince mining personnel that thin membranes are not only safe but also effective.
The layout is similar to the INCO/GRC test because the membrane is again applied onto several discrete concrete blocks. The loading mechanism is however different in concept (refer to Figures 7 and 8), and is as follows:
• Quasi-static loading is applied over the entire test area by means of an inflatable reinforced bag. The load area is typically taken as one metre square, which would represent the spacing of rockbolts underground.
• Dynamic loading is also transferred across the entire test section, and the load is applied by a fixed weight of 150 kg falling a predetermined distance onto the test panel through an impact plate and load distribution blocks.
Results The results are plotted as load versus displacement as shown on Figure 6.
The Dynamic Membrane Test (SRK) This test can provide dynamic and quasi-static results for a support membrane.
Results The loads carried in this test will probably be higher than in the other test methods outlined, because tear strength is a major consideration with this test, and the deformations are less severe in this case, because the area being loaded is larger (ie the 1 m²).
FIG 6 - Applied load versus axial deformation on an experimental 7 mm thick unreactive membrane after 24 hours.
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FIG 7 - The dynamic test arrangement.
Four 100 x 70mm steel channels 20mm thick plywood backing board
Substrate layer 250 x 250 x 100mm concrete blocks
20mm diameter tie-rods
Sprayed layer
125 x 125mm rockbolt washer
Inflatable Vetter bag (Inflated at 1 litre/min)
Spherical seat
Optically referenced displacement scale
150 x 80mm steel channel base frame
(after SRK)
FIG 8 - The quasi-static test arrangement.
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Comments on thin support membrane testing When conducting any support membrane test, it is recommended that the following are also always carefully recorded:
• The temperature during the spraying and regularly before and after the actual testing. The performance of all membranes is influenced by temperature, particularly the elongation.
• The relative humidity. • The thickness of the membrane at the first point of failure, in addition to the average thickness around the failure perimeter. With all the tests however, it must be remembered that the in situ performance is the final judge, and that membranes generally tend to outperform the expectations created from laboratory tests.
PRECAUTIONS AND CONSIDERATIONS Before deciding on using a particular thin membrane, the following need to be considered:
• No membrane can be applied over live (running) water. • A decision must be made as to whether support is needed rapidly (in which case a two component reactive system is needed), or can take sometime to develop (in which case the generally cheaper one component system can be used).
• The useful life of the thin membrane must be established for the specific application. Certain thin membrane systems are prone to significant embrittlement with time (particularly if they contain a cementitious addition), and this needs to be determined.
• As with any membrane application, the substrate must be free from dust, oil and loose materials. This is best achieved by the use of a pressurised water spray. It should be noted that certain membranes do not adhere well in the presence of even a damp substrate.
• If humidity is an issue, the long-term water stability of the product needs to be checked. Most membranes are water-sensitive and their properties are adversely affected in the presence of moisture.
• Appropriate protective clothing and equipment are required and can vary from:
-
goggles, gloves and a dust mask, to goggles, gloves and a charcoal full face filter, to full body suits, respirators and water curtains if isocyanates are used.
MASTERSEAL 840R01 Masterseal 840R01 is a highly reactive polymer membrane system based on methacrylates. It is the only reactive system presently on the market that does not contain isocyanates (ie is not polyurethane or polyurea based). Commercialisation of methacrylates began in the 1930s, and the first application was as a glass replacement in windows. Modern applications now include:
• • • • • •
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automotive coatings, dental resins, contact lenses (for eyes), adhesives, binders, and sealers and floor polishes.
In general methacrylates are considered non-toxic and several long-term exposure studies have shown the material to be non-carcinogenic. The membrane is unique in that it can oversprayed with shotcrete should that be necessary. Typical test results for a 3 mm thickness in a low humidity environment at 20°C are shown in Table 2: TABLE 2 Typical properties for Masterseal 840R01. Age
Tensile strength (MPa)
Elongation (%)
10 min
2.5
119
30 min
3.5
126
1 hour
3.1
124
1 day
3.2
124
90 days
3.1
119
The reaction is exothermic and therefore the temperature of the membrane is higher than the ambient temperature when samples are tested at 30 minutes or less. The typical application thickness is between 2 and 5 mm, with an optimum of 3 mm for most applications. The adhesion of the Masterseal 840R01 onto the substrate is dependent on the nature of the substrate, but typically exceeds that obtained when spraying shotcrete. The two components, Part A and Part B, are both supplied as liquids and are mixed at the spray nozzle at a 2.8 to 1 ratio by weight. The hardened polymer membrane meets the Class II rating for flame spread and smoke development of the Uniform Building Code as published by the International Conference of Building Officials, when tested to ASTM E84-98. It has a flame spread index of 21.4 when tested to ASTM 162.
A COMPARISON BETWEEN WELD MESH AND A SUPPORT MEMBRANE It is difficult to compare two totally different systems. The main difference being the fact that the rock must fail and dilate significantly before the mesh becomes reactive, whereas a membrane becomes reactive with only a small deformation. An additional benefit of a thin support membrane is that it eliminates weathering of the rock, caused by air and moisture over time. Comparing the load versus deformation performance of mesh and thin membranes is difficult as the test methods have frequently been different. Figure 9 compares the behaviour of (6 gauge) 4.88 mm diameter weld mesh (after Espley, 1999) against Masterseal 840R01. It should be noted that the membrane is very rapid to apply and achieves its properties in typically less than 20 minutes. It can be seen that the membrane offers superior support for up to 80 mm of rock deformation, according to laboratory test results. In practice underground however, the mesh is only in rock contact at the rockbolt positions (holding it in place), and is some distance away from the rock everywhere else. This means that the mesh is less effective in situ and that additional deformation must occur before the rock makes any initial contact with the mesh. Assume that as an average this additional deformation needed is an average of 50 mm, the comparative behaviour of the mesh and support membrane is given on Figure 10. The support membrane would therefore be clearly superior to the mesh under most underground conditions.
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THE DESIGN, TESTING AND APPLICATION OF GROUND SUPPORT MEMBRANES
FIG 9 - Idealised comparisons between weld mesh and a support membrane.
FIG 10 - Practical (in situ) comparison between mesh and a support membrane
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A J S SPEARING and J CHAMPA
THE FUTURE
REFERENCES
There can be little doubt that thin support membranes hold great potential as a support element as part of a support system, mainly to replace screen/wire mesh. The mining industry as a whole, in collaboration with suppliers and researchers need to establish objective testing criteria and performance requirements for thin support membranes. The performance requirements will take sometime to develop as in situ experience is gained in different application conditions. The lead must be taken by the industry itself, and not from parties with vested interests, although all parties involved should actively contribute.
Attiogbe, E and Ohler, J, 1999. The MBT membrane displacement testing method. In press. Brower, L, 2000. Personal communication. Espley, S J, 1999. Thin spray-on liner support and implementation in the hard-rock mining industry, MSc Thesis – Laurentian University. SRK Consulting - 2000. Comparative test procedure for sprayed membranes. Promotional Brochure – Johannesburg, South Africa. Tannant, D D, 1995. Load capacity and stiffness of welded-wire mesh. 48th Canadian Geotechnical Conference – Vancouver, Canada. Tannant, D D, 1997. Large scale pull tests on Mineguard. Report to INCO Ltd by the Geomechanics Research Centre. Tannant, D D, Barclay, R and Espley, S J, 1999. Two fields tests on MineGuard, CIM – AGM, Calgary, Canada.
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Remote Monitoring of Rock Mass Deformation During Mining M T Gladwin1, R L Gwyther1 and M Mee1 ABSTRACT
Timeframe
Long Term Pit Slope Stability
3 year
Pillar Stability
100 day
Borehole Tensor Strainmeters Remote monitoring
Remote
Accurate measurement of rock mass response during the mining of massive underground orebodies is essential, as mines become larger and deeper. These measurements enable: provision of quantitative data for optimal pit slope design and engineering; monitoring of the stability of deep pit slopes, the long-term integrity of shafts and large-scale underground infrastructure, and long-term measurements of deformation or rock mass creep following the mining process for environmental management. Strainmeters traditionally used for deformation measurement have neither the long-term stability nor the high sensitivity to perform measurements in these applications. This paper describes a new precision borehole strain monitoring system (GTS) currently deployed by CSIRO Exploration and Mining, which has the potential to solve these new issues in rock mass deformation monitoring. This technology was originally developed for use in hard rock mines (Gladwin, 1977) and refined considerably for earthquake research (Gladwin et al, 1994; Gwyther et al, 1992). A case study of the use of the GTS in monitoring longwall coal mining over the period 1993 - 1998 is presented.
Extensometers
10 day
Foil Gauges
1.0 %
0.01%
1.0 µε
10 nε
Strain Sensitivity
INTRODUCTION Modern mining activities require optimisation of the mine plan for efficient production within well understood stability and safety criteria. Performance of these structural designs needs to be monitored so that the safety margins of the overall mine plan can be assured while the reserve losses implicit in pillars are minimised. This is particularly critical when mining is closely associated with other cultural developments or population centres, where the long-term performance of the total underground structure needs to be continuously and non-invasively monitored for elastic failure, creep and plastic failure events. Commercially available strain gauges are routinely used to measure deformation associated with mining activities, for example in monitoring pillar stability or roof caving. However there is an increasing need to monitor rock mass deformations over a long-term (for example associated with pit slope stability), or remote from mining operations (for example, surface infrastructure or nearby buildings, roads, etc). Previously this has not been possible because the sensitivity and long-term stability of the commercial strain gauges is not adequate. A new precision borehole tensor strain monitoring system (GTS) currently deployed by CSIRO Exploration and Mining has the potential to solve these new issues in rock mass deformation monitoring. The range of sensitivity and stability of strain measuring technology is described in Figure 1. The GTS technology was originally developed for use in hard rock mines (Gladwin, 1977) and refined considerably for earthquake research (Gladwin et al, 1994; Gwyther et al, 1992). The technology has now been refined to a cost level appropriate to large-scale mining operations, and is currently deployed in monitoring of massive sandstone strata deformation above a long-wall coal mining operation in the Bulli basin. In this paper we detail the operational basis of the borehole tensor strainmeter, describe possible applications of the technology in hard rock and open pit mining, and detail a case study of deformation monitoring over a five-year period in the Bulli basin. 1.
CSIRO Exploration and Mining, Pinjarra Hills Qld 4069.
MassMin 2000
FIG 1 - Domains of application in monitoring deformation associated with mining - commercial strain gauges (eg Foil gauges or extensometers) and the borehole tensor strainmeter (BTSM).
THE BOREHOLE TENSOR STRAINMETER (GTS) GTS instruments provide nanostrain sensitivity by measurement of all three-strain components in a plane perpendicular to the borehole axis, induced by tectonic or engineering processes. Earth tilt measurement at nanoradian sensitivity was also available in the instrument modules. Data is available at the measurement site in real time, and can also be used to monitor elastic failure or rock creep processes over extended periods of time. Instruments are commonly installed from the surface in vertical boreholes, and there is no interference in the mining operation. Sites must be carefully chosen to provide representative strains in the monitoring environment, and the target depths are core drilled to ensure adequate coupling conditions for the transducer elements, with the core samples used to determine the elastic moduli of the rocks at the measurement sites. Operational integrity of the instrument is verified in real time by direct continuous measurement of the solid earth tides. These have a peak to peak strain amplitude of approximately 0.05 to 0.1 microstrain (see Figure 2) and are also used to provide in situ calibration for the measurement system Instrumental deformation, u, is monitored in three directions perpendicular to the axis of the borehole. These deformations are related to surrounding rock deformation via shear and areal strain hole coupling factors c and d, and quantities tij which describe the influence of topography and geology in mapping a regional strain field to strains at the instrument site. These parameters are all determined after instrument installation. An example of a shear strain field applied to the instrument is illustrated in Figure 3. The original circular cross-section is deformed into an elliptical shape, and measurements are taken of change in diameter R1, R2 and R3 as indicated in the figure. Raw transducer readings are processed to provide estimates of the measured engineering areal strain in the plane perpendicular to the borehole, εa = εxx + εyy. The engineering shear strains, γ1 and γ2 are calculated in an axis system with the x-axis to the east,
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where U is the deformation vector consisting of the three gauge deformations, G is a direction matrix calculated from the gauge azimuths θi , H is the coupling matrix, T is a matrix describing local topographic and geological influences, and s is the regional strain state at the measurement site. The coupling matrix can be determined by calibration using harmonic analysis of earth tidal signals present in the data. This equation enables determination of the regional strain state in terms of the instrument deformations, and the experimentally determined matrices described here. These measures are then used to calculate any other horizontal strain parameters, for example in mining applications the maximum and minimum compression axes for the strain ellipse can be measured at a site for each point in time. The principal axes of strain in the horizontal plane are defined by directions ϕ0 such that shear strain across planes in these directions is zero, and so:
Solid Earth Tide observed at site 503 60 40
Nanostrain
20 0 -20 -40 -60 20
21
22 23 Days, June 1993
24
tan 2ϕ 0 =
25
FIG 2 - Solid Earth tide observed on strainmeter.
γ2 γ1
For these directions, the extension is: e(ϕ 0 ) = 12 ( ε a ± γ 21 + γ 22 ) It is useful to define two further strain parameters, tensor areal or planar hydrostatic strain:
Shear stress field
P = 12 ( ε xx + ε yy ) = 12 ε a and maximum tensor shear strain (>0 by convention): S=
R1
1 2
γ 21 + γ 22
The principal strains are thus: R2
R3
Reduction in R1 increase in R2 and R3 FIG 3 - Operation of the instrument to measure a shear strain field. Initial state is the circle indicated. A shear field across planes at 45° results in the elliptical shape as indicated. Measurement of the three diameters R1, R2 and R3 enables determination of the magnitude and direction of the strain field.
and the y-axis to the north. The γ1 is standard engineering shear, (εxx - εyy), and γ2 is 2exy. γ1 produces maximum extensional strain in the N-S and E-W directions, whereas γ2 produces maximum extensional strain in NE-SW and NW-SE directions. Strains are analysed from an arbitrary zero after installation has equilibrated. The deformations are related to these quantities by the matrix equation: u1 u = 2 u3
1 2 1 2 1 2
cos 2θ 1 cos 2θ 2 cos 2θ 3
1 2 1 2 1 2
sin 2θ 1 c 0 0 t11 t12 t13 ε a sin 2θ 2 0 d 0 t21 t22 t23 γ 1 sin 2θ 3 0 0 d t31 t32 t33 γ 2
or ~ U = G.H.T. ~s
210
(axis of maximum extension)
ε1 = P − S
(axis of maximum compression)
APPLICATIONS IN LARGE-SCALE MINING
Shear stress across 45° planes results in:
1 2 1 2 1 2
ε1 = P + S
The high resolution of the GTS instrument (1 in 109), and long-term stability (better than 1 in 107 over one year) enables a range of monitoring applications in deep hard rock or open pit mining, for example:
• long-term monitoring of ground deformation in the vicinity of surface mine infrastructure, buildings, pumping stations, etc which are at a distance of 100’s of metres to some kilometres from the mining operation;
• long-term
monitoring of viability of underground infrastructure, eg hoisting and ventilation shafts, declines, large voids such as workshops, crushing excavations, etc;
• monitoring of deformation of the rock mass separating multiple coal seams during coal extraction, providing engineering data for planning extraction of overlying or underlying seams;
• measurement of slope stability in deep pit mines, for example in circumstances where the pit wall surface is decoupled, and monitoring of changes in deformation at distance from the operation; and
• optimisation of pit slope by measurement of actual performance allowing for slope modification. The circles numbered 1 to 4 illustrate some of these uses in Figure 4. Stability of deep pit slopes is a continuing critical concern, with currently deployed technologies based on laser EDM distance measurement, and surface strain and tiltmeters. High precision GPS is also under investigation (Ding et al, 1998). All of these technologies enable monitoring of slope changes of mm
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REMOTE MONITORING OF ROCK MASS DEFORMATION DURING MINING
FIG 5 - Strain field resulting from a crack of length 100 m, and depth of 100 m, opening to a maximum of 10 mm at the surface. Modelled contours of maximum shear strain in a horizontal plane at a depth of 100 m are indicated over an area of 1 km2 around the crack. The instrument resolution over a period of tens of days is better than ten nanostrain.
4500
CBT
FIG 4 - Applications of borehole strainmeter in monitoring a large-scale open pit mine: for example; 1) Deep pit slope stability during mining, and optimal pit slope design; 2) Long-term shaft integrity; 3) Stability.
CASE STUDY: SOUTH BULLI LONGWALL MINING GTS instruments were deployed in 1993 at Cataract Reservoir, south of Sydney, to provide long-term monitoring of longwall mining induced strain variations beneath the reservoir (Gladwin et al, 1998). A mine plan of the extraction sequence monitored is shown in Figure 6. The coal seam is at 450 m depth, and instrument installation was at approximately 100 m depth in order to evaluate strains at the critical depth of the reservoir floor. Sites were chosen to provide an early estimate of maximum mining induced strains (CCT), a forward analogue of the final approach to the dam wall (CAT), and direct measurement of any effects related to the dam abutment itself. Target depths were core drilled to ensure adequate coupling conditions for the transducer elements.
MassMin 2000
Dam Wall
3500
3000 517
516 515 514 513 512 511
2500 Meters
size. The inherently higher resolution of the GTS technology enables it to be installed at distances of 100’s of metres from a pit slope without loss of monitoring capability. Slope failure is often preceded by surface cracking. To illustrate the rock deformation resulting from a typical crack, Figure 5 shows modelled contours of maximum horizontal shear strain in the surrounding rock mass resulting from a 100 m long crack with a depth of 100 m, opening by 10 mm at the surface. A simple assumption of elastic deformation in a half space was used for the modelling. The values shown are for a depth of 100 m, a typical installation depth for the GST instrument. The GTS resolution and stability (better than ten nanostrain over periods of weeks) is well within the expected strain magnitudes even at distances of 500 m.
4000
2000
N
1500
509
CAT
508
507
1000
506
505 504 503 502 501
500
0 -1500
-1000
-500
0 Meters
CCT 500
1000
FIG 6 - Mine extraction plan at South Bulli, with instrument sites shown as CAT (above panel 508) and CCT (above panel 503). The dam wall site is 3 km to the north.
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Results The measurement program began in 1993 immediately prior to the beginning of the extraction sequence for the first panel in the series (501). Measurements at all three sites were taken at 30 minute intervals, recorded locally at the site, and every three hours transmitted via a VHF link to a central recording facility at the operations centre at the Cataract dam. This site provided a redundant recording system and offsite direct access to the data via a dial in modem interface. The control software at the site included remote intervention for maintenance of the data, for
storage and general processing of the data. At regular intervals the data was extracted from the unmanned base station, network performance and timing was verified, and the data archived and processed at CSIRO. Figure 7 shows a time sequence of principal axis strain data measured at site CCT (above panel 503), with the upper plot (a) showing the maximum shear (S) and hydrostatic strain (P), and plot (b) the maximum and minimum compression axes of the strain ellipse relative to the undisturbed state. Initially the longwall was about 1200 m to the south of the instrument. The
FIG 7 - a) Maximum shear and hydrostatic strains observed at site CCT (panel 503). b) Maximum and minimum compressional strains. c) Principal axes of observed strain. d) Progression of mining.
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REMOTE MONITORING OF ROCK MASS DEFORMATION DURING MINING
lower plot panel indicates the principal axes interpretation of the strain sequences measured at CCT, at each point in time marked with a principal axis plot. Arrowheads on the principal strain axes give the direction of strain, and the azimuth of the axis plot gives its normal geographic azimuthal orientation (east to the right, north up). The bottom plot shows the position of the advancing longwalls measured from the western end of each panel. The longwall passed under the CCT site about 30 July 1994 (1994.6), at which time the site went into approximate cross panel uniaxial compression. Mining operations in panel
509 have an influence at the CCT site (1400 m distant) at a level of less than 10 microstrain. The maximum horizontal compressive strain reached approximately 80 microstrain as the 504 panel was mined, shear strains reached 50 microstrain, and the mean hydrostatic strain was approximately 20 microstrain. Strain changes were evident at the CCT site in the 503 panel immediately the extraction of 501 panel was begun. Strains over the same interval measured at the site CAT (above panel 508) are shown in Figure 8. Borehole recovery effects (associated with the local strain relief surrounding the borehole
FIG 8 - a) Maximum shear and hydrostatic strains observed at site CAT (panel 508). b) Maximum and minimum compressional strains. c) Principal axes of observed strain. d) Progression of mining.
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M T GLADWIN, R L GWYTHER and M MEE
214
1800 511 1600
509
1200
508 CAT
100 Microstrain
1400
507
1000
Meters
during drilling) are superimposed on observed strains in borehole instrumentation. These effects have been removed to facilitate analysis of the mining induced strains (this procedure was not possible for CCT because mining of 501 began immediately after installation). Longwall operations passed beneath site CAT in November 1996, and the magnitudes of strain offsets observed were similar to those observed at CCT site (panel 503) when operations were in close proximity to the CCT site. The detail of strain response at this site was markedly different to that observed at CCT. In particular the relaxation of north/south strain (εyy) observed at CCT during mining in the preceding panel was not observed at CAT site during mining of the preceding panel. The strains at CAT resulting from the mining operations are almost fully shear strain. The changes in azimuth of principal axes indicate that the maximum shear strain (S) is orientated at angles of approximately 10° and 100° East of North. Note that the hydrostatic strain term (P) indicates net dilation (ie positive hydrostatic strain) at CAT site after mining of panel 508 was complete, with a current overall increase in dilation of 10 to 20 microstrain since commencement of mining panel 506. This result contrasts with an overall net compression of approximately 10 microstrain as was observed at CCT. There is also clear evidence of a rate increase in dilation or tensional strain (ie decrease of compressional strain) beginning approximately at the commencement of mining panel 506. This rate of decrease of compressional strain continued until at least 1998. A map view of the principal axis strains measured at site CCT due to the progression of the longwall is shown in Figure 9. The figure shows the position of longwall panels 501 to 511 in map coordinates. The mining proceeded from east to west in panels 501 to 508, followed by panels 511 and 509. A series of principal axes plots have been superimposed to indicate the state of strain (relative to the undisturbed state) observed at the CCT measurement site when the long wall sequence had reached the particular position. The stars indicate the sites CCT and CAT. The maximum strains above the panel (at the instrument depth) can be estimated directly from the figure. Azimuth is relative to the north point shown. The result clearly indicates a significant rotation of the principal strains with progression of the longwall. Detail of strain meter response to longwall operations over a 12-day period when the face operations were directly below the instrument (a depth of approximately 400 m) are shown in Figure 10. The periods of mining activity are shaded. Note that they reflect the fact that face cutting occurs in two shifts separated by a maintenance shift interval, which can clearly be identified. The major strain changes induced by the mining are restricted essentially to the periods of mining activity. The figure indicates the direct causal relationship of the observed loads to mining operations. Detailed examination indicates that the immediate elastic response followed by a short-term plastic response is consistent with expectations from measurements of surface subsidence data. Tiltmeters installed in the instrument provide a measure of earth tilt with a resolution of better than one nanoradian. Figure 11 shows North and East tilt records from two observation sites, with the convention is that a +ve tilt in a direction is a tilt of the upward vertical towards that direction. The data indicate that tilts of order 200 microradians occurred at site CCT during the closest approach of the longwall mining to the instrument site, with a total tilt amplitude change of 400 microradians due to the mining activity, predominantly downwards to the east (along panel). The tilt records are well correlated with the mining process. Tilt changes at site CAT during close longwall operations show remarkably similar behaviour to those observed at site CCT. Following completion of panel 508, mining of the relatively distant panel 511 was carried out prior to mining of the
506 800 505 600
N
504 400
503
200
501
0 -1000
CCT
502
-500
0
500
1000
Meters
FIG 9 - Map view of principal axes of strain observed at site CCT during longwall operations. Each set of strain axes shown at a position in a panel represents the strains observed at CCT when the longwall face was being mined at that position.
FIG 10 - Strainmeter response to longwall mining directly beneath the instrument. During this period mining was carried out for two shifts, with a maintenance shift each day (with no mining activity). The regular goaf fall, with consequent deformation, is evident during the two working shifts each day.
adjacent panel 509, and the overall tilt amplitude observed at CAT is higher (of order 600 microradians), again predominantly downwards in an easterly direction. The dam wall site tilt records are constant at these scales. The complete tilt data set from panel 503 indicates that significant subsidence induced tilts occurred by longwall activity at least two panels from the measurement site. Large tilts (>400 microradian) occurred during the mining, but returned approximately to initial conditions after two further panels were mined. These plots show that tilt is predominantly influenced by longwall mining of five adjacent panels. Detailed investigation of tilt changes as each panel is mined indicates that significant tilt changes occurred 15 days before panel 502 face passed site CCT and 20 days before panel 503 face passed by CCT. These times
Brisbane, Qld, 29 October - 2 November 2000
MassMin 2000
REMOTE MONITORING OF ROCK MASS DEFORMATION DURING MINING
indicate that significant tilt changes were initiated when the face position was approximately 180 m distant from the observation site. Reversal in the azimuth of the maximum tilt direction, with a temporary tilt northwards, are evident each time a panel face passes the observation site, with the exception of panel 502 face
The CAT site, mined in November 1996, indicates strains of similar amplitude but different character to those observed at site CCT (mined in mid-1993). Changes in the strain components (εxx, εyy and εxy) as mining is carried out in the preceding panel have differed significantly between the two sites.
Discussion
• At site CCT (panel 503) there was a significant relaxation of
Data from both CCT and CAT sites indicate that, at the measurement depths, horizontal strains of order 100 microstrain are locally associated with the progression of the longwall. These strain estimates are the change of strain state relative to the initial (undisturbed) environment at the beginning of the measurements.
cross-panel strain (εyy) as the face in panel 502 passed by CCT, and the strain due to the mining of panels 502/503 was essentially cross-panel uniaxial strain (see Figure 4) as expected.
FIG 11 - East and North tilt records at observation sites CAT and CCT.
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M T GLADWIN, R L GWYTHER and M MEE
• At site CAT (panel 508), the cross-panel strain (εyy) decreased (ie added tensional strain) as the face in panel 507 passed by CAT site and along-panel strain (εxx) increased (extensional strain). These changes indicate that the strain field associated with the mining of panels 507/508 was essentially shear strain at the observation depth of 90 m (see Figure 9), with directions of maximum shear of approximately 10° and 100° east of North. Site CAT (508) showed an unexpected decrease in compressional strain as adjacent panels were mined, resulting in an overall decrease of 10 to 20 microstrain, compared to an overall increase of 10 microstrain in compressional strain at CCT (503). It should be noted that these strain changes occur in the presence of a significant lithostatic stress load. At 100 m, this produces compressive strains of order 100 microstrain in the horizontal plane, which increase linearly with depth. Strains from these gravitationally induced horizontal compressive stresses are added to the strain changes observed to determine the total strain field. If the current rate of decrease of compressional strain continues, the site would go into net tension in ten years. Tilt signals at the two sites showed similar behaviour at each observation site during the progression of longwalls. Behaviour of changes in the maximum tilt direction varied significantly at the two sites, however both indicate approximately east-west tilting, consistent with the geometry of the panel layout. The tilt changes recorded at site CCT (503) indicate that subsidence close to the surface is predominantly influenced by longwall activity in five adjacent panels, with significant tilt occurring at a site directly above a panel when mining operations are approximately 180 m from the site. Strains observed at site CAT during mining of CCT were less than ten microstrain, and strains observed at CCT during mining of CAT were similarly less than ten microstrain and indicate the expected strain at a distance of 1 km from the mining operations. Strains observed at the Dam Wall site show that mining induced strain effects are of magnitude less than one microstrain at this time.
CONCLUSIONS Observations of the type detailed in this case study allow planning engineers to modify initial mine plan towards higher recovery ratios with the benefit of data on the actual performance of the in situ rock. In the case reported here, panel widths were increased by five per cent for panels following 507, after the
216
observations from mining of panels 503 - 506 indicated that strain amplitudes were well below rock failure stress conditions at the observation depth. Applications requiring higher sensitivity are unlikely in mining. Creep and plastic deformation of rock very close to mine workings imposes loads on intact rock remote from these workings. These loads result in elastic deformation of this relatively remote rock mass, and the observations of a GTS instrument installed in this rock mass can provide information in real time on potential failure of the rock mass structure to perform as expected. In summary, the extremely high sensitivity of the GTS instrument enables remote monitoring of deformation, with no disturbance to mining operations. The high stability of the GTS measurements provides reliable data on the bulk response of the engineering structures over a long timescale, in cases where mining operations extend over a number of years.
ACKNOWLEDGEMENTS The authors acknowledge funding support for this research from Bellambi Collieries Pty Ltd. We thank Terry Cheeseman of Sydney Water for his ready assistance in field maintenance, and Mark Kochanek and Zak Jecny of CSIRO Exploration and Mining for assistance during installation and field repairs.
REFERENCES Ding, X, Montgomerey, S B, Tsakiri, M, Swindells, C F and Jewell, R J, 1998. Integrated Monitoring Systems for Open Pit Wall Deformation, ACG Meriwa Report No 186. Gladwin, M T, Gwyther, R L and Mee, M, 1998. Precise Deformation Monitoring for Stability of Mine Structures, in Proceedings International Conference Geomech and Ground Control in Mining and Underground Construction. Gladwin, M T, Breckenridge, K S, Gwyther, R L and Hart, R, 1994. Measurements of the Strain Field Associated with Episodic creep events at San Juan Bautista, California, Journal of Geophysical Research, 99(B3):4559-4565. Gladwin, M T, 1984. High Precision multi component borehole deformation monitoring, Reviews of Scientific Instruments, 55:2011-2016. Gladwin, M T, 1977. Simultaneous Monitoring of Stress and Strain in Massive Rock, Pure and Applied Geophysics, 115:267-274. Gwyther, R L, Gladwin, M T and Hart, R H G, 1992. A Shear Strain Anomaly Following the Loma Prieta Earthquake, Nature, 356(6365):142-144.
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MassMin 2000 Block and Panel Caving The Palabora Underground Mine Project
K Calder, P Townsend and F Russell
219
Cave Mining at Premier Diamond Mine
P J Bartlett and A Croll
227
Panel Caving Experiences and Macrotrench (Macrozanja) — An Alternative Exploitation Method at the El Teniente Mine, Codelco, Chile
G Diaz and P Tobar
235
Evolution in Panel Caving Undercutting and Drawbell Excavation, El Teniente Mine
J Jofre, P Yáñez and G Ferguson
249
The Pre-Undercut Caving Method at the El Teniente Mine, Codelco Chile
E Rojas, R Molina, A Bonani and H Constanzo
261
Esmeralda Mine Exploitation Project
M Barraza and P Crorkan
267
‘An Underground Air Blast’ — Codelco Chile Division Salvador
R de Nicola Escobar and M Fishwick Tapia
279
Freeport Indonesia’s Deep Ore Zone Mine
J Barber, L Thomas and T Casten
289
Excavation Design and Ground Support of the Gyratory Crusher Installation at the DOZ Mine, PT Freeport Indonesia
T Casten, R Golden, A Mulyadi and J Barber
295
Commissioning of Two 750 kW Centrifugal Fans at PT Freeport Indonesia’s Deep Ore Zone Mine
F Calizaya, T Casten and K Karmawan
301
A Case History of the Crusher Level Development at Henderson
M F Callahan, K W Keskimaki and W D Rech
307
The New Henderson Mine Truck Haulage System — The Last Step to a Totally Trackless Mine
W D Tyler, K W Keskimaki and D S Stewart
317
Application of Block Caving System in the Tongkuangyu Copper Mine
Zhou Aimin and Song Yongxue
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The Palabora Underground Mine Project K Calder1, P Townsend2 and F Russell3 ABSTRACT
GEOLOGY, HYDROLOGY AND ORE RESERVES
Palabora Mining Company operates an integrated open pit mine, concentrator, smelter and refinery complex located in the Northern Province of South Africa. The Palabora open pit went into production in 1966, mining 30 000 t/d to produce some 62 000 tpa of copper. With continuous modification and improvement, a mining rate of 82 000 t/d, producing up to 135 000 tpa of refined copper, has been achieved. While these improvements have also enabled the mine life to be nearly doubled, the open pit will reach the end of its economic life by 2003. The Palabora orebody is vertical and the reserve extends to a depth of 1800 m over an area of 700 m by 200 m. Studies to assess the economics of mining the orebody by underground means began in the mid-1980s. These studies indicated that block caving is the lowest cost underground mining method that would generate a suitable return. In March 1996, a decision was made to develop a 30 000 t/d block cave mine that will extend the life of the Palabora Mining Company by a further 20 years. In terms of the relatively high rock strength and the resultant coarse fragmentation that is expected at the drawpoint, the Palabora cave can be considered rather unique when compared to other block cave operations. This paper describes the project in broad terms and details the mining method, particularly with respect to the cavability of the orebody and the design features to deal with the coarse fragmentation. The project commenced in July 1996 and final development of the undercut and drawbells is scheduled for completion in November 2003. This paper describes the status of the project and planning as of April 2000.
The Palabora copper orebody is an elliptically shaped, vertically dipping volcanic pipe (Palabora, 1976). The pipe measures 1400 m and 800 m along the long and short axes, respectively. The orebody is open at depth with reserves proven to 1800 m below surface. Copper grades of approximately one per cent are found in the central core of the orebody and decrease gradually towards the peripheries with no sharp ore/waste contact. Analyses during the feasibility study determined that the optimum grade boundary is 0.8 per cent Cu (Kear, 2000). This cut-off resulted in a mineable reserve of 245 Mt at 0.68 per cent Cu. An additional resource of 467 Mt grading at 0.57 per cent Cu, is situated adjacent to and below the 30 000 t/d reserve base. This ore is available for mining and constitutes a resource base for a life of mine extension or an expansion in production rate. The cut-off grade for the mine has been calculated at 0.31 per cent Cu. Mineralisation is hosted by three main rock types. Transgressive and banded carbonatites form the central core of the orebody and are made up of magnetite rich sövite with minor amounts of apatite, dolomite, chondrodite, olivine and phlogopite. Barren dolerite dykes, with a steeply dipping north-east trend are present and account for approximately eight per cent of the 245 Mt reserve. Two sources of water inflow have been identified, groundwater and rain falling within the catchment area of the pit and flowing through the caved rock. The total amount of water is likely to increase over the period to full production and the establishment of a cave through to surface, in approximately four to five years. The flow will increase in stages:
INTRODUCTION Palabora Mine is located in the Northern Province of South Africa about 560 km north east of Johannesburg. Elevation of the pit rim is about 400 m above sea level. The climate is subtropical with an average annual precipitation of 480 mm. Palabora Mining Company was originally established as a joint venture between Rio Tinto (formerly RTZ) and Newmont Mining Corporation in 1956 to exploit the copper resource identified around a hill known as Loolekop, where there was archaeological smelting evidence from the 8th century. Construction of an open pit mine started in 1963 and processing of ore began in 1966. Subsequent expansions increased milling capacity to 82 000 t/d ore and 135 000 tpa of cathode copper. Capacity increases in the metallurgical plant allowed for economies of scale to be realised in the mining operation and the cut-off grade has been lowered from the initial 0.30 per cent Cu to the present 0.10 per cent Cu. Open pit life was similarly expanded, ultimately to nearly double the originally planned 20 years. The present mining plan calls for open pit operations to cease in the year 2003. Although the orebody continues below the bottom of the final pit shell, the stripping ratio precludes further open pit mining. Various studies into the viability of underground mining were carried out from the mid-1980s culminating in a final feasibility study in 1996 with the project commencing in the same year.
1.
Project Director, Palabora Mining Company, South Africa.
2.
Project Manager Technical, Palabora Mining Company, South Africa.
3.
Principal Consultant, Rio Tinto Technical Services, Castlemead, Lower Castle Street, Bristol BS99 7YR, England. E-mail: [email protected]
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• initial flows from groundwater only - estimated to be about 16 L/s;
• groundwater and possibly catchment water flowing through fissures connecting the pit to underground; and
• groundwater and rainwater flowing through the caved rock estimated to average 66 L/s. Installed pumping capacity is 300 L/s and the pump station is protected by watertight doors that are closed in the event of inflow exceeding pumping capacity, currently predicted to occur when rainfall is in excess of the 1:100 year event.
GEOTECHNICAL INFORMATION The average uniaxial strength of the carbonatites is about 120 MPa, with a variation in values depending on mineralogy between 90 MPa and 160 MPa. Dolerite is a strong brittle rock with a uniaxial strength of 320 MPa. Adjacent to the major faults, dolerite is locally weathered with a marked reduction in strength to around 80 MPa. The in situ state of stress is assumed to be hydrostatic and approximately equal to the overburden load of 38 MPa. As part of the numerical and fragmentation analyses, parametric studies were carried out between the limiting values of horizontal stress ratio of 0.75 and 1.5 to determine whether the solutions were sensitive to the inherent assumptions about state of stress. The structure of the carbonatites is dominantly subvertical jointing. These joints are open or infilled with weak material, planar and through-going. There are three steeply dipping sets,
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striking approximately 010 (dip direction 290), 310 (dip direction 040) and 050 (dip direction 140). The flat lying joints have a different morphology from the vertical sets: wavy, rough and of limited continuity. There are two sets oriented approximately 20/160 and 45/350. For the purposes of the fragmentation analysis, the rock fabric was idealised to three major sets. The fracturing of dolerite is closely spaced and blocky. Within the cave area, the fracturing in dolerite was established to be zoned by proximity to the major faults. For the purpose of mine design the orebody has been divided into Less and Well jointed zones based on the formation of primary fragmentation at a size cut-off of 2 m3. RMR (Laubscher) values were estimated from core samples (Table 1).
TABLE 1 Rock quality. Zone
Average RMR
Less Jointed
70
Well Jointed + Dolerite
57
Orebody Average
61
MINING Summary The underground mine will exploit ore below the final open pit shell using mechanised block caving. The undercut level, currently being mined at an elevation of 1200 m below surface and approximately 460 m below the ultimate pit bottom, is the uppermost level of the mine. The production level, containing the drawpoints and other infrastructure, is being developed 18 m below the undercut. During the early stages of the cave propagation, it is anticipated that a significant percentage of caved material will be too large to be handled by the ore extraction equipment and specialised rigs will be used to drill and break the oversize, using both emulsions and non-explosive techniques. A fleet of 11 diesel-powered LHDs with a 14 t payload will muck from the drawpoints directly to four crushers located along the northern periphery of the cave. The average one-way length of haul is 175 m. The crushers will discharge onto sacrificial conveyors which in turn will feed onto a horizontal section of a single 2000 tph inclined conveyor, delivering the ore to two 5000 t capacity production shaft silos. The ore will be hoisted out the mine using four 32 t payload skips. All major underground fixed equipment and the fully automated service and rock hoists will be monitored and controlled from a control room situated on surface using tele-remote systems where applicable. The underground project has four shaft systems including the ventilation shaft. The first shaft to be sunk was the exploration shaft, which was sunk initially from bench 30 in the open pit to 889 m below surface elevation to develop an exploration drive to facilitate exploration of the deeper lying ore. After project approval, the shaft was deepened to the production level to allow an accelerated development program to be put in place concurrent with the sinking of the two main shafts. The exploration shaft is a 4.8 m diameter shaft equipped with two 6 t skips and a 27 man service cage. The shaft has a 4000 t/d hoisting capacity and will remain in production until the end of 2000, after which it will be maintained as an emergency exit.
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Sinking of the new concrete lined service and production shafts to the final depth of 1280 m below surface was completed in the third quarter of 1999. The shafts are situated 72 m apart. The 10 m diameter service shaft has an 86 m high concrete headframe and is equipped with a large single deck cage running on fixed guides and a 20 man capacity auxiliary cage running on rope guides. The main cage has a 35 t pay-load and can accommodate 155 people. It is licensed to operate at 12 m/s and the auxiliary hoist operates at 8 m/s. The production shaft has a diameter of 7.4 m and a 106 m high concrete headframe (Taljaard, 2000). The shaft is being equipped with four 32 t pay-load skips running on rope guides. Maximum hoisting capacity will be in the order of 42 000 t/d. The production shaft, first crusher and conveyor system will be commissioned early-2001. All the main hoisting systems utilise tower mounted friction winders with integrated motors. The two rock hoist winders are each powered by 5500 kW motors. Both shafts are sited outside the open pit shell beyond the influence of the block cave. All the hoists are fully automated and will be operated from the surface control room. Intake air is downcast through the production and service shafts and is exhausted through a 5.76 m diameter, 924 m deep raise bored ventilation shaft. The two main 1250 kW upcast ventilation fans are installed on bench 28 of the open pit. At present, only one fan is being utilised and is exhausting 340 m3/s. Both fans will exhaust a total of 500 m3/s via the ventilation shaft. Two 850 kW booster fans have been installed underground which will enable 600 m3/s to be downcast through the production and service shafts and will put the cave area under positive ventilation pressure, minimising the ingress of dust and heat into the workings through the cave. An 18 MW refrigeration plant supplies chilled water to a bulk air cooler adjacent to the main shafts. The intake air to the underground workings is cooled to offset the high underground ambient rock temperatures of 50°C and heat generated by the diesel equipment. Major underground infrastructure, including workshops and offices, is being constructed on the production level, close to the mine workings but outside the zone of abutment stresses. Current production from the open pit is processed at the concentrator, consisting of crushing, grinding and flotation circuits. The concentrate is then smelted in a conventional furnace and converter with a capacity of 135 000 tpa of fine copper. The final process is the electrical refinement of the anode copper to produce cathode copper. When the underground mine is in production surplus process plant capacity will be utilised to increase overall recovery through changes in grind, increased retention time in the flotation circuit and reducing concentrate grade. An Environmental Impact Assessment (EIA) was completed as part of the underground project feasibility study. The conclusion was: ‘No impacts were identified on the mine property, resulting from the proposed development of the (underground) mine that cannot be mitigated adequately’. The key issues highlighted in the study were: reduced water consumption, reduced waste disposal and reduced emissions.
Undercut design The design and operation of the undercut is the key mechanism for initiating the cave. The features of the undercut design at Palabora are:
• An advanced undercut whereby the undercut is developed ahead of opening the drawbells to provide a ‘stress shadow’ to protect the production level.
• A narrow undercut 4 m high. This is considered sufficient to initiate cave while minimising the amount of swell material that has to be removed during excavation using an advanced undercut.
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THE PALABORA UNDERGROUND MINE PROJECT
• An inclined face is made over the major apex creating a chevron shape to the undercut, to facilitate undercut initiation following the construction of the drawbell by promoting the flow of blasted undercut material into the bell. The resultant design is shown in Figure 1. The undercut development is supported with a combination of resin grouted bolts, mesh and shotcrete, taking into account the large stress change that will occur as the undercut area is enlarged.
15m
12m
55
O
U NDE R CU T L E V E L
18m
P R ODU CT I ON L E V E L 13m
FIG 1 - Undercut design.
Production level The optimum drawpoint spacing is a balance between obtaining good draw and recovery characteristics while maintaining the strength of the drawpoint structure with a layout appropriate to equipment size and type. The offset herringbone layout was selected principally as a function of LHD manoeuvrability and the potential to use electric LHDs with trailing cables - with the crushers located on one side of the production area, all the drawpoints point towards the incoming LHD. In addition, with the return ventilation on the opposite side to the crushers, any dust from the loaded LHD bucket is coursed away from the LHD and the driver. The drawpoint spacing was designed at 17 m centres as a function of the isolated draw zone characteristics of the coarse ore. The drawbells are rectangular with inclined walls and offset between production drives spaced at 34 m. There is scope for minor modification of the drawbell shape to improve draw characteristics, if necessary. The bell layout leaves substantial pillars for support and protection of the drawpoint structures. Numerical modelling has been carried out to check the integrity and the reinforcement for the drawpoints. Production tunnels will be 4.5 m x 4.2 m with drawpoint cross-cuts sufficiently long to allow the 6.5 m³ LHDs to load without articulating. The design is shown in Figure 2. The production level development is supported with a combination of resin grouted roof bolts, cable bolts and fibrecrete. Steel sets may be placed at the bow of drawpoints in weaker ground.
PRODUCTION LEVEL N
NORTH MAIN ACCESS
CRUSHER No4
CRUSHER No2 CRUSHER No1
CRUSHER No3
INTAKE FANS
NORTH RIM DR IVE
MAIN WORKSHOP
PRODUCTION SOUTH INNER SERVICE DRIVE 0
50
100
150
200
250
PRODUCTION SOUTH OUTER SERVICE DRIVE
MAIN RETURN AIRWAY CONNECTION
Scale metres
FIG 2 - Plan of production level.
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Loading LHDs will load from drawpoints to crusher tips accessed from the main production drive located on the northern periphery of the cave. Initially, LHDs will be operated manually but development into using semi-autonomous LHDs, operated from the control room on surface, is well advanced. Secondary breaking equipment and services will access the production area via two service drives along the southern rim. The activities of the underground fleet will be monitored and directed from the central control room on surface via an auto dispatch system. LHDs will be dispatched to drawpoints and crusher tips based on a production schedule that will be generated using information stored in a cave management database (Penswick and Dasys, 1997).
Crushers and conveying level Because of the coarse fragmentation that is expected, the decision was taken to deliver the ore direct to a crusher (as at Northparkes mine) rather than risk potential hang-ups in ore passes and problems with loading systems. To optimise LHD travel distance and provide flexibility if a crusher is not available, four crushers will be installed along the northern side of the production area. Each of the four crusher stations will have two tipping bays to enable two LHDs to tip simultaneously. Bays will be equipped with a scalping grizzly with a maximum aperture of 1.4 m. Oversize will be broken by a hydraulic rock breaker operated tele-remotely from the control room on surface. Ore will be crushed to minus 200 mm by Krupp 1700 mm x 2300 mm jaw crushers. A total of 750 t of live storage capacity is provided below each crusher. Ore is then delivered via sacrificial conveyors onto the main conveyor inclined at 9° (reducing the production shaft depth of wind by 117 m) and discharging into two storage bins at the production shaft. The incline conveyor haulage will not be used as an intake airway to the mine workings due to the risk of dust pick-up and toxic fumes in the event of a fire. The ventilating air will discharge directly into the return airway system.
Dilution control The estimate of mineable reserve includes all material in an 85° draw angle from the cave perimeter and is low-grade ore rather than waste. To minimise any further dilution it is planned to draw down all drawpoints evenly at a rate not to exceed 200 mm/d. A cave management system that includes a dispatch system linked to LHD tagging will be used to ensure that the cave is drawn at the required rate. Estimates of production tonnes and grades by period were made using an in-house program and later verified using Gemcom’s PC-BC program (Diering, 2000).
Cave height
Cavability With the rock quality being higher than any other block cave, extensive studies were carried out to gain confidence that the undercut area would cave. At the time of feasibility study the model with the greatest industry acceptance to predict cavability was the Laubscher Stability Diagram. This diagram (Figure 3) consists of a graph that plots rock strength, as measured by Mining Rock Mass Rating (MRMR) against hydraulic radius, which is the size of area available for undercutting. The Stability Diagram is an empirical predictor which experience at other mines has proven to be conservative but reliable. Figure 3 shows the MRMR for the rock characteristics listed in Table 1. Assuming global average rock strength the model indicates that caving will commence when a hydraulic radius of 35 m, equating to an area of two hectares, has been undercut. The total Palabora footprint comprises an area of 12.6 hectares. Figure 4 shows the plan of the undercut level and the sequence of opening up the undercut.
Fragmentation With the rock quality being higher than any other block cave, studies were made to determine the size range of the rocks reporting to the drawpoint. A program (Block Cave Fragmentation - BCF) was specifically developed by D H Laubscher and G L Esterhuizen to generate rock size distribution using, as input data, the cave face orientation (dip and dip direction), stress data, MRMR, joint set orientations and joint set spacing. The program also caters for secondary fragmentation, that is the breakage of primary blocks into smaller fragments, through repeated splitting, corner rounding and splitting when temporary arches fail. Another program was developed in-house by R M Kear to estimate the hang-up potential for the rocks generated by the BCF program. Fragmentation problems will be most severe when the cave is initiated. Estimates using the BCF program indicate that in the first year of production over 70 per cent of the rock will be greater than 2 m³ and will therefore require secondary breaking before being loaded. As the cave height is increased, the action of broken rock moving toward the drawbells will further reduce rock size, hence, the number of hang-ups that could occur will be reduced. As part of the feasibility study, fragmentation data, hang-up prediction and clearance times were input into a dynamic simulation of the entire production process. Results of the simulation, using Systems Modelling ARENA program, indicated that, in the year of worst fragmentation, the target production rate of 30 000 t/d could be achieved.
Manning levels
The design column height for Palabora is approximately 460 m, varying according to the intersection with the pit walls. This is high in comparison to most caving operations, however, at these operations the limitation of cave height is sometimes a function of orebody configuration and drawpoint life. In turn, drawpoint life can be seen to be more related to damage created by abutment stresses and subsequent damage by bad secondary blasting practice rather than wear during operation. The high rock quality at Palabora, a disadvantage in terms of fragmentation, becomes an advantage when related to potential drawpoint wear. By using an advanced undercut and suitable secondary breaking equipment to minimise damage it is not expected that
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the relatively high cave height will present a problem to the operation.
Manning levels for the underground mine are as shown in Table 2.
Capital cost The projected capital cost to construct and commission the underground mine is $US410.2 million in escalated terms (Table 3).
SAFETY AND HEALTH RELATED ISSUES Palabora Mining Company places great importance on the safety aspects of all operations. In this respect block caving is seen as a
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THE PALABORA UNDERGROUND MINE PROJECT
FIG 3 - Laubscher Stability Diagram.
TABLE 2 Underground mine manning levels.
TABLE 3 Estimated capital expenditure ($US ‘000).
Area
Number
Mining Operations
210
Exploration shaft
2.8
Mining Technical
45
Mine surface facilities
15.5
Engineering Services
94
Refrigeration and ventilation
13.6
TOTAL
349
Production shaft
48.6
safe mining method. The only activities carried out during operations will be secondary breaking, load-haul-dump followed by crushing, conveying and hoisting. The following indicates the measures that have been incorporated into the design to create a safe working environment for these activities.
Estimate Area
$US M
Service shaft
36.9
Mining
142.9
Materials handling
20.4
Services
25.8
Control and instrumentation
14.4
Indirect costs
73.8
Subtotal
394.7
Secondary breaking
Exchange rate allowance
15.5
Because there will be occurrences of drawpoint oversize and blockages beyond the reach limitations of conventional equipment, a specialised high reach drill rig has been developed to drill and charge these high hang ups remotely without personnel entering the drawpoint (Penswick, 1997) (Figure 5). The high reach rig has the capability of accessing blockages 21 m above footwall elevation. It is equipped with a three-dimensional video system and is operated remotely from a mobile control module that detaches from the main unit. The rig
Total
410.2
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is equipped with an emulsion charging system to load the drilled holes. Non-explosive breaking techniques will be used to break drawpoint oversize. This technique may be extended to the drawbell hang-ups.
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N 9
8 7
6 5 4 3
2
3
4
5
6
1
FIG 4 - Plan of undercut level and undercut sequence.
FIG 5 - High reach drill rig.
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Loading
ACKNOWLEDGEMENTS
LHD operations will be segregated from all other operations. The LHDs will operate between the drawpoints and the crushers located in the north main production drive. Personnel will not be allowed access to areas where LHDs are operating. Secondary breaking and service units will enter the production cross-cuts from the South access drift but only when the LHD has completed production activities. A personnel detection warning system is already employed on the development fleet which warns the LHD operator that personnel are in the proximity of the LHD. This will also be installed on the production fleet of LHDs. The use of a semi-autonomous loading fleet is currently under investigation (Penswick and Kirk, 1976). Tele-remote loading will be used in conjunction with an advanced guidance system that will automatically take the LHD from the drawpoint to the tip and return. This greatly reduces the number of operators required and removes them from the underground environment.
Conveying and hoisting All crushing, conveying and hoisting operations will be controlled from a central control room located on surface.
FUTURE DEVELOPMENTS The mine design is based on existing technology, but the overriding philosophy is to use proven developments in technology and automation to the benefit of the operation in terms of safety and productivity. On-going research and development in automation of LHD operations could result in this technology being implemented within the next 18 months. Automation of many other routine tasks is being actively investigated on an on-going basis. It is believed that the mine being developed at Palabora will determine the viability of block caving in rock previously considered too difficult to cave. This has the potential to open up many other deposits around the world using this low-cost, but safe, mining method.
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The paper presented here is the work of all those who have participated in the Palabora Underground Mine Project and the role of the authors has simply been to present this work to a wider audience. The permission of Palabora Mining Company and Rio Tinto Technical Services to present this paper is gratefully acknowledged.
REFERENCES Diering, T, 2000. PC-BC A Block Cave Design and Draw Control System, in Proceedings MassMin 2000, pp 469-484 (The Australasian Institute of Mining and Metallurgy: Melbourne). Kear, R M, 2000. The Use of Evaluation Surfaces to assist in the Determination of Mine Design Criteria, in Proceedings MassMin 2000, pp 57-62 (The Australasian Institute of Mining and Metallurgy: Melbourne). Palabora Mine Geological and Mineralogical Staff, 1976. The Geology and the Economic Deposits of Copper, Iron and Vermiculite in the Palabora Igneous Complex: A Brief Review, Economic Geology, 71(1):Jan/Feb. Penswick, D P, 1997. The Palabora High Reach Rig, SAIMM Journal, 97(2):March/April. Penswick, D P and Dasys, A, 1997. Developing Dispatch for Underground; Palabora Looks to the Future in CIM’97, 99th Annual General Meeting, Vancouver, Canada, 27 April -1 May. Penswick, D P and Kirk, R L, 1997. The Potential for Automation at Palabora in 4th International Symposium on Mine Mechanisation and Automation, Brisbane. Taljaard, J J, 2000. The State of the Art Shaft System as Applied at Palabora Underground Mine. Paper to be presented a Mine Hoisting 2000, Johannesburg, 2000.
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Cave Mining at Premier Diamond Mine P J Bartlett1 and A Croll2 ABSTRACT
INTRODUCTION
A panel cave using LHDs for ore extraction came into production at Premier in 1990 at a depth of 630 metres below surface on the western side of the mine in the BA5 mining block. A post-undercut mining sequence was used with all development on both the undercut and production levels completed prior to the undercut being advanced over the pre-developed drawbells. Support and rock mass damage on the production level was extensive as the undercut moved overhead. Undercutting and coarse fragmentation created problems. The mining sequence was changed to an advance undercut with the minimum of development and support undertaken on the production level prior to the undercut moving overhead. Only once the abutment stresses had stabilised were drawbells developed and final support installed. The height of the undercut drill design was decreased from 18 to 13 metres. The changes resulted in markedly less rock mass and support damage on the production level. Lessons learned in the BA5 mining block on the 630 level were incorporated into planning and scheduling the BB1E mining block which came into production on the 732 metre level on the eastern side of the mine in 1996. An advance undercut mining sequence was planned, undercut height was minimised and undercutting parameters were optimised in terms of the BA5 experience. Numerous problems still occurred. As the undercut area increased in size severe spalling from the cave back resulted in vertical propagation of the cave to a height of 30 metres instead of the planned 13 metres. The resultant coarse fragmentation created secondary blasting problems and an excess of undercut ore that had to be loaded and trammed, slowed the rate of undercut advance. The planned mining sequence was compromised and considerable support and rock mass damage resulted on the production level below. Once these problems had been overcome the undercut advanced at the planned rate and continuous caving initiated at the predicted hydraulic radius. At a later stage the unexpected ingress of water hampered undercut drilling and slowed the rate of undercut advance in a section of the cave. Leads and lags between adjacent tunnels and the overall face shape could not be developed as planned. The rate of natural horizontal cave propagation exceeded the rate of undercut advance and three undercut drilling tunnels were extensively damaged. More tunnel development on the production level prior to undercutting than was planned resulted. The slow advance of the undercut and high abutment stresses caused extensive damage to installed support and the rock mass on the production level, 15 metres below. Experience in the BB1E mining block on the 732 metre level has taught that an advance undercut mining sequence reduces damage on the production level markedly. Extremely tight control on the rate of undercut advance, leads and lags between adjacent tunnels and overall undercut face shape as well as extraction ratio, tunnel development and support sequence on the production level below is required. Premier Mine plans to implement a block cave at a depth in excess of 1000 metres within the next decade. All the experience gained on current mining levels together with data collection and numerical modeling is being used to plan the proposed mining operation. The geotechnical challenge of mining at this depth in weak rock has been identified as the greatest risk to the project. Challenges centre around adopting the correct mining sequence, limiting the extent of rock mass damage associated with undercutting and effective support. The hydraulic radius needed to initiate caving, fragmentation, drawpoint spacing, secondary blasting and the possibility of seismicity are geotechnical concerns that are being addressed. Results to-date suggest that a pre-undercut mining sequence with all development and support on the production level carried out only after the undercut abutment stresses have passed overhead will have to be implemented.
Premier Diamond Mine, located 45 kilometres to the east-north-east of Pretoria in South Africa, exploits the Premier diamond pipe, the largest kimberlite occurrence in South Africa, with a surface area of 32 hectares. The pipe is geologically unique in that it is cut at a depth of 400 metres below surface by a 75 metre thick, dipping, gabbro sill. Figure 1 shows the location of the mine. Mining operations started in 1902, first by open pitting to a depth of 189 metres, followed by underground mining which started in 1948. Block cave mining was first introduced in the early-1970s with four caves operated above the sill producing 85 million tons of ore using scrapers to move the ore from finger raises sited below drawpoints to ore passes. The four caves were sited at depths of between 390 and 420 metres below surface. Mining infrastructure was sited in competent kimberlite and gabbro with rock mass ratings ranging from 55 to 75. Mining operations first started below the sill in 1979 using an open stope mining method. Problems with the method forced a change to block caving to exploit the reserves that existed below the sill. Cave mining using LHDs to transport the ore from drawpoints to passes started in 1990 on the 630 metre level on the western side of the mine in the BA5 mining block. A second cave was established in 1996 at a depth of 732 metres below surface in less competent ore on the eastern side of the pipe in the BB1E mining block. Since mining operations started the Premier Mine has produced 322 million tons of ore and 124 million carats of diamonds. An inferred resource exists below current mining levels and a recently completed evaluation program and feasibility study have been undertaken to better define and investigate mining of the resource. Mining of the C-cut would involve the exploitation of 170 million tons of ore using cave mining methods with a production level 1000 metres or more below surface. Mining at Premier would be extended by at least 17 years and could produce an additional 85 million carats of diamonds. Figure 2 diagrammatically illustrates the geology of the pipe and the location of the actual and proposed mining blocks.
1.
Chief Geologist Premier Diamond Mine, Private Bag X1015, Cullinan 1000, South Africa.
2.
C-Cut Project Manager – Premier Diamond Mine, Private Bag X1015, Cullinan 1000, South Africa.
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BA5 MINING BLOCK The BA5 mining block was planned as a panel retreat cave with a post-undercut mining sequence. An area large enough to ensure caving was prepared. Tunnels on both the undercut and extraction levels were developed in a north-south direction across the pipe. Undercut tunnels were developed directly above production tunnels and spaced at 30 metres. All development and support on the extraction level, including drawbells, was completed prior to undercutting. The block was then undercut by drilling and blasting a slot at right angles to the undercut and extraction levels 120 metres long and 30 metres wide. The initial undercut height was 20 metres. As undercut rings were blasted broken ore dropped directly into the pre-developed drawbells. As the undercut area approached the hydraulic caving radius stress levels on both the undercut and extraction levels increased. It became increasingly difficult to drill, charge and blast the long undercut drill holes and the rate of undercutting slowed. On the extraction level rigid shotcrete linings were extensively damaged by the abutment stresses associated with the now slow moving undercut face above. Footwall heave was widespread, damaging
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FIG 1 - Regional map showing Premier Diamond Mine.
concreted roadways and disrupting production. When continuous caving initiated stress levels stabilised, but an extensive programme of roadway and tunnel support rehabilitation was needed to ensure the safety of men and equipment and guarantee uninterrupted production. The rate of caving of the kimberlite was much as expected increasing to a final rate of two metres per day. When the cave back reached the competent sill caving ceased until a hydraulic radius of 47, measured at the base of the sill, was exposed. Caving of the competent sill was intermittent and slower than anticipated resulting in a large and potentially damaging air-gap. The rock above and surrounding the air-gap was monitored on a daily basis using drill holes that had been drilled into the air-gap. Access into the level where the collapse of the air-gap had the potential to vent was by permission of the mining manager only. Tunnels into the actual air-gap were sealed off with re-enforced concrete walls. Draw control and monitoring ensured that a column of ore at least 70 metres thick existed between the production level and the air-gap to ensure that production personnel would not feel the effects of an air-blast. As the panel retreat cave extended eastwards the high abutment stresses associated with the advancing cave face continued to damage roadways and support on both the undercut and extraction levels, necessitating costly and time consuming repairs. Blasted ore from the high undercut flooded undercut tunnels. Subsequent loading of ore to allow access to drill, charge and blast the undercut face slowed the rate of undercut advance. A decision was taken to adopt an advance undercut mining sequence with only the production tunnel and drawpoint break-aways developed and partly supported prior to the undercut moving overhead. Once the area had been de-stressed, development on the extraction level, including trough opening, was completed. Final support included concreting of roadways and application of a shotcrete lining. The height of the undercut drill design was reduced from 20 to 12 metres.
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The measures resulted in reduced support requirements and a marked reduction in rehabilitation of tunnel support and roadways. Production targets were met and the undercut face could be advanced at the planned rate. The overall face shape and leads and lags between adjacent tunnels were maintained in terms of the mining plan. The main lessons learned in the BA5 were that:
• High and damaging abutment stresses are associated with the undercut advance.
• Where the total stress approaches 20 per cent of the rock
•
mass strength extensive support damage can be expected. Rigid support in the form of concreted roadways and shotcreted or concreted tunnel linings will be damaged as the undercut moves overhead. Where the total stress exceeds 50 per cent of the rock mass strength tunnels are often crushed.
• The damage induced by high stresses can be alleviated by changing the mining sequence.
• An advance undercut reduces the extent of rock mass and support damage.
• Stress levels can be minimised by advancing the undercut according to the mining plan.
• The rate of undercut advance must be at the planned rate. • Every effort must be made to ensure that undercut blasting achieves complete undercutting with no remnant pillars (stubs) left behind. • Leads and lags between adjacent tunnels must be eight metres or less. • The overall face shape must be maintained. Prior to mining the BA5, numerical stress modeling of the proposed mining sequence had been undertaken. Only with hindsight was correct interpretation of some of the numerical stress model results possible.
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FIG 2 - Diagramattic section and plan showing geology and mining block.
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TABLE 1 Rock mass parameters. Rock Type
Tensile Strength (MPa)
Uniaxial Compressive Strength (MPa)
Young’s Modulus(Gpa)
Poissons’s Ratio
Specific Gravity
Rock Mass Rating
Brown Kimberlite
7.9
50 - 80
16.0
0.27
2.67
45 - 55
Grey Kimberlite
6.7
80 -130
29
0.17
2.67
50 - 60
Black Kimberlite
13.7
73 - 193
34 - 88
0.2 -0.4
2.8
55 - 72
Gabbro sill
24.0
180 - 400
119
0.33
2.81
65
Felsite
16.8
240 - 300
62
0.29
2.6
72
Norite
16.8
140 - 220
74
0.25
2.8
45 - 60
5
60 - 240
30 - 167
0.15 - 0.33
2.5 - 2.8
40 - 65
Meta-sediments
TABLE 2 Block caving parameters. Parameter
BA5 Mining Block
BBIE Mining Block
80 - 140 metres
148 - 164 metres
Rock mass rating
45 - 65
45 - 55
Hydraulic radius
30
25
Mining sequence
Post and advance undercut
Advance undercut
Rate of undercutting
900 square metres per month
1100 square metres per month
Tons in Mining block
42 million tons
23 million tons
Tons per drawpoint
50 000 to 120 000 tons
100 000 to 200 000 tons
Drawpoint spacing
15 × 15 metres
15 × 18 metres
22.6 metres
23 metres
Average rate of draw
180 mm per day (109 tons)
165 mm per day (120 tons)
Initial fragmentation
30 % > 2 cubic metres
30 % >2 cubic metres
Fragmentation after 20 % drawn
12 % >2 cubic metres
7 % >2 cubic metres
Cable anchors, rockbolts, mesh tendon straps and shotcrete
Cable anchors, rockbolts, mesh tendon straps and shotcrete
Brow wear
0 to 2 metres wear after 50 000 tons drawn
1 to 3 metres wear after 50 000 tons drawn
Tunnel size
4 × 4.2 metres
4 × 4.2 metres
Diesel and electric 5 and 7 yard Toro's
Diesel and electric 5 and 7 yard Torro's
Column height
Distance across major apex
Drawpoint support
Lhd Type Tons per LHD per hour
118 tons
131 tons
154 metres
134 metres
Hangup frequency
30 % of drawpoints per shift
25 % of drawpoints per shift
Fragmentation Initial 20 drawn Secondary blasting
30 % > 2 cubic metres 12 % > 2 cubic metres 40 grams per ton
30 % > 2 cubic metres 10 % > 2 cubic mtres 30 grams per ton
LHD average tramming distance
Calibration of the model did, however, show that some of the empirical lessons learned in the BA5 cave could have can be pre-determined for a planned cave by careful core logging, laboratory testing and, as underground exposures become available, cell mapping to determine accurate parameters for numerical stress modeling. Stress and displacement monitoring as the cave moves towards the hydraulic caving radius must be undertaken to calibrate the stress model, which can then become an accurate predictive tool. Some of the lessons learned, such as optimum rate of undercut advance and leads and lags between adjacent tunnels, are peculiar to an orebody and can only be learnt by monitoring, experience and intelligent observation. Interaction with other cave mines can accelerate progress along the learning curve considerably.
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The layouts of the BA5 undercut and extraction level are illustrated in Figures 3 and 4 respectively.
BB1E PRODUCTION BLOCK Lessons learned in the BA5 were applied in the BB1E mining block. The block was planned as a panel retreat cave with an advance undercut mining sequence. Only the production tunnel and drawbell breakaways were developed on the extraction level prior to the undercut passing overhead. Support was installed in two phases. In the first phase all required rockbolts were installed and tunnels supported with mesh and tendon straps. Shotcrete was applied only where this was required to prevent damage to the installed steelwork by LHD loading and secondary blasting operations. Roadways were compacted rather than concreted.
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FIG 3 - Layout of BA5 undercut level.
FIG 4 - Layout of BA5 production level.
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The overall shape of the undercut face, leads and lags between adjacent tunnels, the rate of undercut advance, undercut drill ring design and the infrastructure and equipment to move blasted ore was planned and scheduled in detail. As the undercut initiated vertical propagation of the cave was faster than anticipated. Drill ring design blasted out the undercut to a height of 13 metres, but spalling from the cave back extended this to a height of between 18 and 30 metres as soon as undercutting started. Far more ore than anticipated, much of it in the form of coarse fragments that had to be drilled and blasted, had to be moved on the undercut level. The rate of undercut advance slowed dramatically and resulted in tunnel convergence, footwall heave and support damage on the partly developed extraction level 15 metres below. Time consuming and expensive rehabilitation of production tunnels had to be done before production could start. Completion of extraction level development and drawbell opening was carried out in adverse mining conditions. Only when the cave reached its hydraulic radius of 25 and stress levels reduced could the mining plan in terms of production tons, scheduled development and ring drilling be achieved. The second phase of support which included rehabilitation of any damaged support, shotcrete lining of the entire tunnel and concreting of roadways was completed only after the undercut had moved overhead and the area had been destressed The mining block had initially been dry, but as mining operations progressed stress changes allowed existing fractures and joints to open and water moved into the kimberlite down the pipe contact and through the surrounding wall rock on the northern margins of the cave. Decomposition of the clay-rich kimberlite followed, hindering undercut drilling and blasting operations. The rate of undercut advance slowed, leads and lags between adjacent tunnels became unacceptably large and the overall shape of the undercut was affected. Undercut advance was no longer in phase with development on the extraction level below and part of a production tunnel on the extraction level was crushed as the undercut moved slowly over the now fully developed, but partly supported, tunnel. The rate of natural horizontal propagation of the cave exceeded the rate of undercut advance. Large induced fractures developed parallel to the undercut face across several undercut tunnels destroying support and making the tunnels unsafe for drilling and blasting operations. Contingency plans had to be developed and implemented. An attempt to control the inflow of water into the cave was made by developing tunnels up to the pipe contact at 30 metre intervals on the undercut level. Tunnels in the kimberlite were graded to allow water to flow towards the pipe contact and holes were drilled to allow water to flow under gravity through return air and waterways to the water sumps. Efforts were only partly successful. The undercut is now advancing at an increased rate and lead and lags and overall face shape are again according to plan. There is, however, considerable damage on the extraction level below which has affected production in the area. This has had to be rehabilitated. Where the advance undercut has been implemented as planned an effective cave mining operation has been established. Experience has, however, taught that the planned mining sequence must be achieved. This demands accurate planning , availability of resources and careful control and implementation of the entire process. The sequence shown in Figure 6 must be completed within six months after the undercut has passed over a drawbell or production tunnel to avoid compaction of ore. Maintaining such a schedule can be extremely difficult. The risk of not being able to achieve the planned schedule is high and the economic consequences of failure equally high. Effective undercutting to ensure that no remnant pillars (stubs) are left unblasted and maintaining the planned rate of undercut blasting is essential. The undercut ring design as shown in Figure 5 has
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proved effective for undercut blasting at Premier Mine. The geometric relationship of the undercut to the extraction level is illustrated in the same section. The undercut ring design is based on experience gained on mine in undercutting over a number of years, discussions with Dr D H Laubscher and new insights brought to undercutting by a Russian mining engineer. The expeditious opening of drawbells once the undercut has moved overhead is important to avoid compaction of ore. Drawbells are opened by ring drilling from the drawbell crosscut on the extraction level to the undercut level above. A blind-hole bore drilled from the extraction level into the broken ore on the undercut level serves as a free breaking face. The time taken to open a drawbell has been greatly reduced by pre-drilling the rings, increasing the diameter of the blind hole bore from 600 to 1000 millimetres and using electronic delay detonators in blasting. Once ring drilling and blind-hole boring have been completed the drawbell can be opened to its final dimensions of 13 metres by 15 metres in three blasts. The drawbell geometry is illustrated in Figure 7.
Phase 3
RAISEBORE HOLE
FIG 7 - Isometric view of drawbell.
77m2
Phase 2 7m
Phase 1
12m
5m
9m
7m
9m
4m 7m
15m
1.5m
4 m 4m
1.8m 30m
FIG 5 - Section showing undercut ring design and geometry of undercut and extraction level.
Partly develop Extraction level Limited Production to avoid compaction
Partly support Extraction level Complete final support
Advance undercut overhead
Develop drawbells
Concrete roadways
FIG 6 - Advance undercut development and support sequence.
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Steel tendon straps and chainlink wiremesh covered with 100 millimetres of shotcrete has proved an effective support system, but is time consuming and labour intensive to install. The mine is looking at using fibre re-enforced wet mix shotcrete as an alternative. A decision to implement a pre-undercut has been taken in principle in the BB1E as the geotechnical benefits are overriding. Experience and numerical simulation of support and development at a depth in excess of 1000 metres in weak kimberlite has shown that a pre-undercut mining sequence will have to be implemented at this depth. Experience gained in pre-undercutting will be invaluable for the C-cut.
C-CUT MINING BLOCK The C-cut mining block will be implemented at a depth in excess of 1000 metres below surface. A new mine will be re-engineered with two new shafts being developed from surface. The ore hoist will house four 30 ton skips and the men and materials shaft will be designed to move 50 ton loads. A new ore processing plant using modern technology will be built. The mine will be designed to extract and process nine million tons of ore a year with a labour force of some 700 employees. Shaft sinking is planned to commence in 2001 and ore extraction in 2004 building up to nine million tons by 2008.
The column height will be between 350 and 450 metres. Ore in the block will total some 170 million tons. Each drawpoint will produce between 200 000 and 275 000 tons of ore. All the cave mining experience gained to-date in the BA5 and BB1E blocks is being used to design the C-cut. A detailed risk assessment has shown that the greatest risk to the project is the geotechnical risk associated with the development of the cave mining infrastructure in the kimberlite. Both experience and numerical simulation have shown that a pre-undercut mining sequence will have to be implemented. Several changes are being implemented in current mining blocks to test their effectiveness with a view to implementing these ideas in the C-cut. Figure 8 shows the layout of the C-cut extraction level. All geotechnical aspects are being examined to understand and minimise risks. Numerous geotechnical parameters are being considered to see whether any advantage can be gained by optimising these relative to current practice on the mine. Specific areas that are being addressed include: a.
Tunnel size Tunnel size will be 4.2 × 4.2 metres on the production and the undercut levels. The overriding concern is that efficient equipment that will allow the mine to achieve planned rates of development, drilling, support and production can be
FIG 8 - Layout of C-cut production level.
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used in the tunnels. Any increase in tunnel size is associated with a commensurate increase in the depth of the zone of fracture propagation around the tunnel. b.
Layouts The mine currently uses an offset herringbone layout -originally copied from the Henderson Mine - on the extraction level, motivated primarily by concerns around ventilation and electric LHD loading constraints. Numerous benefits, especially during development, are associated with the El Teniente layout and the geotechnical benefits of the two layouts were evaluated. Geotechnically there appears to be little long-term benefit to be derived by moving to the El Teniente layout.
c.
Mining sequence Both experience and the numerical simulation of stress changes have clearly indicated that a pre-undercut mining sequence will have to be implemented in the C-cut to avoid extensive support damage and even crushing of production tunnels. The logistical implications consequent on this decision are being evaluated and the mining sequence in the existing BB1E mining block is being changed to a pre-undercut mining sequence to allow personnel to gain experience in implementing the mining sequence.
d.
• Experience has shown that modern cave mining methods can be used to extract kimberlite ore cost effectively from the diatreme zone of a kimberlite pipe.
• Increased stress levels that accompany progressive deepening of the mining operations from the 390 metres to in excess of 1000 metres in a geologically consistent orebody have forced a change in mining sequence and support methods in cave mining.
• Modern technology in the form of efficient LHDs, long reach secondary blasting drill rigs, shotcrete, electronic draw control and numerical modeling have allowed increasingly productive cave mining operations.
ACKNOWLEDGEMENTS The author would like to thank the management of Premier Mine and De Beers Consolidated Diamond Mines (Pty) Ltd for permission to publish this paper. The extensive discussions and contributions made by personnel involved in cave mining at Premier Mine and other cave mining operations to this paper is gratefully acknowledged. The use of diagrams prepared by colleagues at Premier Mine personnel is acknowledged.
Drawpoint spacing
REFERENCES
Numerical simulation shows that there is little stress-related geotechnical advantage to increasing drawpoint spacing from 15 metres by 15 metres to 15 metres by 18 metres in terms of a decrease in tunnel and pillar damage. Current planning is that a 15 metre by 15 metre drawpoint spacing will be implemented in the C-cut. Drawpoint spacing at Premier is currently 18 metres by 15 metres resulting in a maximum drawpoint spacing of 23 metres diagonally across the major apex. The BB1E mining block is investigating increasing drawpoint spacing to 18 metres by 18 metres. Should this spacing prove successful the advantage of implementing this increased spacing will be re-evaluated for the C-cut. e.
CONCLUSIONS
Bartlett, P J, 1992. The Design and Operation of a Mechanised Cave at Premier Diamond Mine, in Proceedings Massmin 92, Publication Symposium Series S12 (South African Institute of Mining and Metallurgy: Johannesburg). Bartlett, P J, 1994. Geology of the Premier diamond Pipe, in Proceedings XVth CMMI Congress (Ed: H W Glen) Vol 3, pp 201-213 (South African Institute of Mining and Metallurgy: Johannesburg). Leach, A R, Naidoo, K and Bartlett, P J, 2000. Consideration for design of production level drawpoint layouts for a deep block cave, in Proceedings Massmin 2000, pp 357-366 (The Australasian Institute of Mining and Metallurgy: Melbourne.
Effect of increasing vertical distance between undercut and extraction level Numerical simulation shows little benefit in increasing the spacing beyond 18 metres as, with increasing depth, the extraction level moves progressively out of the stress shadow of the pre-undercut. Practically an increased distance between the undercut and extraction level would make some hang-ups less accessible. The results of numerical modeling have been summarised and presented in a paper by Itasca at this conference.
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Panel Caving Experiences and Macrotrench (Macrozanja) — An Alternative Exploitation Method at the El Teniente Mine, Codelco-Chile G Diaz1 and P Tobar2 ABSTRACT The development plan (long-term plan) of El Teniente Mine, Codelco-Chile, includes 100 per cent primary ore, located at increasing depths, which may involve several difficulties related to the application of the current panel caving method, such as a high seismicity level, a coarse fragmentation and ore pass damage. Therefore, it is a strategic priority for the El Teniente mine, to design and to test alternative mining methods, at similar or lower costs to ensure an efficient and safe ore extraction. In this context, the Macrotrench method is developed, as a competitive alternative to the current method. The document describes the panel caving method, the experience achieved in the primary ore exploitation from the geomechanical point of view. It also describes the ‘macrotrench’ method as an alternative to the current exploitation method. Finally, the comparative advantages of macrotrench in relation to panel caving are presented. The results obtained during the engineering stage include the following aspects: • the mining design, • the material handling, • a geomechanical evaluation of the method, and • a comparative analysis.
INTRODUCTION El Teniente Mine is located southeast of Santiago and 60 km northeast of Rancagua City in Central Chile. Figure 1 shows the location of all Codelco Chile Divisions and El Teniente Mine. The mine is owned by Codelco-Chile and 1.100 million tonnes of ore have been mined out during almost 100 years of exploitation. Its production has been mainly obtained by a block-caving method with grizzlymen (secondary ore exploitation). In 1982, a mechanised panel caving method was introduced, with LHD equipment, in order to extract the primary ore. A currently production of 100.000 tonnes/day is achieved, with a projected production of 126.000 tonnes/day by the year 2004, which would mean 450 Kt of fine copper yearly.
THE GEOLOGY AND GEOTECHNICS El Tenient is one of the biggest deposits of porphyry copper in the world. The main rock types of the deposit are: Andesite (73 per cent), Diorite (12 per cent), Dacite (nine per cent) and Brecchia (six per cent). Different rockmass types are included within these lithologies and they are characterised by a different degree of fracturing and geomechanical properties. There is a great difference between secondary ore and primary ore. Schematically, El Teniente deposit is formed by a central brecchia body or pipe (1.0 to 1.2 km diameter), surrounded by a mineralised rockmass over a variable radial extension between 400 and 800 m (see Figure 2). 1.
Chief Mining Engineer, Codelco-Chile El Teniente Division, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
2.
Project Mining Engineer, Codelco-Chile El Teniente Division, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
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There are two systems to classify the rock mass and they are both used indistinctly: Rock Mass Rating (RMR by D Laubscher) and so called ‘Q Index’ from L Barton. Those indexes, are applied on a rock mass zonation defining the geotechnical units. The lithology, the mineralogical characteristics and the structural domain define these units. These geotechnical units define regions of homogeneous strength properties of the intact rock, fracture frequency and discontinuity shear strength. Table 1 shows the classification of the rock mass.
THE PANEL CAVING METHOD In general terms, the primary rock has been exploited applying a panel caving method, which consists of five levels: undercut, production, ventilation, reduction and train haulage. The following one is a brief description of the design:
• Undercutting level: 18 m above the production level, undercutting galleries are developed parallel to the haulage drifts, on 30 m spacing centre to centre. From this level, drilling and blasting are performed to produce the undercutting.
• Production level: The layout consists of parallel haulage drifts on 30 m centres, intersected by a set of 17 m spaced parallel drifts (trenches) at a 60º angle. Ore passes are located in the haulage drifts every 80 – 100 m.
• Pickhammer level and ore passes: The ore passes are 3 m diameter, from dumping points in two neighboring haulage drifts (calles) converging to a stationary pickhammer sublevel station 35 m below. Blocks dumped by the LHD will be reduced at this station.
• Train haulage level: It consists of looped cross-cuts at different distances (75 to 105 m) capable of fitting trains with 30 tonnes capacity ore cars.
• Handling of materials: In the production level 7 yd3 LHD dump ore down ore passes located each 100 m. The material arrives through the ore passes to reduction stations where it is reduced using breaker hammers. Finally, it arrives (via ore pass) to the transport level where it is carried by trains with 30 tonnes cars. Figure 3 shows a typical panel caving block.
EXPERIENCE IN THE EXPLOITATION OF PRIMARY ROCK IN EL TENIENTE MINE From the beginning of primary rock exploitation there have been geomechanical problems, related either to the exploitation method designs or to rockbursts that have occurred in that environment. El Teniente mine, started the exploitation of primary ore in 1982, applying a panel caving exploitation method, using LHD equipment. El Teniente mine is a pioneer in the exploitation of primary rock. It has gained a great amount of knowledge related to mining a primary rockmass, when compared with different behaviour of the secondary rockmass.
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FIG 1 - Location of Codelco Chile divisions and El Teniente Mine.
It has had to face some great challenges during its productive process, due to the rockmass physical-properties and the deepening of the production sectors. This environment includes higher stresses and a complex geology and geotechnics (lithology and structures). The rockmass have presented diverse instabilities, such as: slabbing and over-excavation of galleries, structural blocks formation, collapses, subsidence and the rockburst phenomenon. Simultaneously to the application of panel caving, a series of studies and tests have been conducted, mainly in the geotechnical-geomechanical field, in order to answer the questions that have arisen during the application of this method.
PRIMARY ROCK IMPACT ON THE EXPLOITATION METHOD Outstanding instabilities at El Teniente Mine The main instabilities that have been attributed to the exploitation method are the following:
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• • • •
slabbing and over-excavation, structural blocks formation, collapses, and rockbursts.
A brief description of these instabilities is presented and the applied measures to prevent its occurrence and to attenuate its consequences, mainly oriented to the production process and to protect the infrastructure and service facilities.
Slabbing and over-excavation of galleries This type of instability is the most common one and it affects horizontal and vertical galleries. Along horizontal drifts, it is associated with rock falls, mainly during the excavation stage. In order to minimise this problem, the designs consider some stress reduction measures in the surrounding rock, such as:
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FIG 2 - Isometric view, El Teniente Deposit.
TABLE 1 Classification of the rock mass. Units Secondary andesite
ff/m
3
RMR
10 - 20
30 - 35
Primary andesite
5-9
40 - 62
Primary diorite
4-9
55 - 65
Primary dacite
4-6
65 - 70
Brecchia Braden
1-4
65 - 75
• to place drifts appropriately in order to reduce the formation of small wedges due to minor structural features; and
• changes in the drift section geometry (shape and size). Corrective measures intended to maintain support for the drifts, range from a development fortification with grouted bolts and mesh, to a definitive support that includes grouted bolts, cables, mesh and shotcrete in the productive areas. For vertical galleries, such as ore passes, applied measures include: locating them in zones without structural relevant features, support involving cables, steel linings and a suitable mining design.
Structural blocks formation. It corresponds to major over-excavations and it is characterised
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by the formation of large wedges movements along relevant structural features and the limits of caved areas. They are also present in large-scale galleries or at their intersection areas. Applied measures aim to anticipate the structural block movements through changes of the geometry and orientation of the caving front, the drift location versus the structural block location and an appropriate support. This last measure sometimes becomes corrective, especially when the presence of a structural block is detected too late. In the case of ore pass systems, measures are related to the mining design, excavation method, ore passes location, support with cables, and the ore passes filling as a working favorable practice.
Collapses This corresponds to the gradual failure of a rockmass over a great area, usually on the production level, with observed damage at the crown pillars and production level, whose maximum expression is the total closing of the affected drifts. As a consequence, the production area is reduced. Several studies have been conducted to understand the causes of this instability phenomenon in order to apply the corresponding actions to reduce its occurrence and to mitigate the observed damages. According to subsequent observations, preventive actions are to keep a strict draw control. Some corrective actions are the recovery of affected drifts and to increase draw within the surrounding areas in order to avoid the formation of big structural wedges associated to the collapse.
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FIG 3 - Isometric view of panel caving with LHD.
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Reserves loss associated to collapses are recovered through the development of new production levels located under the affected production area.
Rockbursts Beyond any doubt, this stability phenomenon is the most complex one, mainly due to the rupture and the sudden violent rock displacement. These characteristics do not allow immediate actions as in other instabilities that evolve slowly making possible some actions in order to reduce or to mitigate its effects. The actions applied by El Teniente Mine have included strategical and tactical measures. The strategical measures include the application of geotechnical and geomechanical criteria in order to reduce the risk of the phenomenon occurrence, such as the definition of the initial caving zone and its shape, the exploitation sequence (undercutting-extraction), the mining design (geometry, size and drifts arrangements) and the definition of the undercutting and the extraction rates. The tactical measures are adopted after a rockburst, such as the drift support for the rockmass reinforcement and the retention of the projected material. Thus, depending on the severity of the damages, a bolt support is used (low severity of damages) or a combination of bolts, cables, mesh and shotcrete (high severity of damages). A Processes Control framework has been developed at El Teniente mine, searching for a relationship between the induced seismicity and the mining activity in order to understand the origin of the seismic induced activity in terms of the mining parameters. This approach allows the control of the seismic rockmass response through the modification of the mining parameters in order to reduce the associated seismic risk. Mine planning parameters that contribute to the generation of induced seismicity are the undercutting rate, the extraction rate and the production trench mining schedule. The basic assumption of this option is to assume that the mining activity produces a rockmass disturbance, including a rockmass breaking process. The seismic events are included in the rockmass response to the mining activity and arise from the radiated energy coming from the primary rock ruptures. Regarding the time evolution of this framework, two research approaches were developed: a static one and a dynamic one. The static approach consisted of identifying some outstanding factors that condition the seismic response of the rockmass to the mining. The main identified factors were the geology and the rockmass geometry as they are the main rockmass parameters involved in the rupture process. A seismic risk zoning of the mine was achieved, with three risk categories were defined: low, medium and high. This work was based on two main parameters: rockmass column height and volumetric fracture frequency. The dynamic approach introduces a process control concept that relates the seismic rockmass effect to the mining process. The impact of the whole mining process is considered over the rockmass under exploitation and the rockmass characteristics that contribute to the seismic rockmass response parameters. A Mining Activity Index has been developed as an estimation of the seismic rockmass response accordingly to the mining and the rockmass parameters. Simultaneously, an Alert Criteria is used, which indicates to the mine personnel that the seismic response is far from the expected one. The subsequent analysis of the seismic activity should determine the necessary mining activity modifications in order to restore a controlled seismic rockmass response. This application is based on the information provided by the seismic system installed in the mine.
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Drawpoint and ore pass damage Based on experience and on the observations made of the drawpoint and ore pass behaviour, the damage occurs before and during the passing of the undercut face. In the first stage, the damage during this period is the result of the effects of the abutment stress concentration generated by the undercut face line. The second stage corresponds to a transition period, in wich damage is caused by various phenomena such as relaxation generated by the undercut, undercut and trench blasting. The main damage occurs for an extraction ranging from 0 to 25 per cent of ore.
PANEL CAVING EVOLUTION RESULTS Mining of the primary ore exploitation has undergone serious difficulties during the last decade. The most serious situation took place at the Teniente Sub-6 area, located completely in primary rock. This area starts its production in 1989 and temporarily shut down its operations in 1992 due to some seismic episodes with associated rockbursts. The rockbursts involved casualties and relevant economic losses as the result of the production shut down and the considerable infrastructure damages. Later studies and developments allowed the reopening of Sub6 production in 1998. In the near future, the situation will become more difficult due to an increasingly complex geomechanical behaviour as mine exploitation gets deeper. In addition, a significant increase of the primary rock exploitation volumes is expected. Currently, 50 per cent of extracted ore corresponds to primary rock, a percentage that will increase to 100 per cent by the year 2006. Codelco-Chile, and specifically División El Teniente, has dedicated many years to the study and practice of primary rock exploitation, assigning considerable technical and economic resources. During this period, the division has gained valuable experience. Solutions with direct application are hard to find due to the particular features of our deposits and the Codelco exploitations on a large-scale, forcing an internal development of the solutions. The experience and the knowledge of Codelco regarding the primary ore exploitation are recognised by the world mining industry. The previously presented information proves the importance of enhancing deep mining methods in primary rock, in order to be able to create and to implement actions that allow us to change the mining standards of our underground exploitations. This action ranges from the consolidation of caving knowledge and the ore breaking process to the tests of alternative exploitation methods. El Teniente Mine has been able to validate its conceptual framework through the empirical evidence developed in an experimental productive area, obtaining important results for the mining activity. These include:
• Seismic activity is part of the response of a competent rockmass to the way mining is performed, that is, how ore extraction and undercutting are carried out.
• There should be a planning and operation personnel commitment to the results related to seismic control because of their influence on the handling of the mining variables which impact the induced seismic response.
• Based on the acquired knowledge in relation to the exploitation of the primary rock with a panel caving method, it is concluded that the abutment stress generated at the caving front is the main source of the damage to the drifts located at the lower levels.
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• It is important to keep a periodic control of the whole mining process in order to reduce the possibility of rockburst occurrence.
• The experience gained at the El Teniente Mine during the last few years has allowed an increase in the theoretical and practical knowledge about the way primary rock mining should be performed. The control of the negative effects of the seismic activity associated to a primary rockmass exploitation has been learned.
• After almost two decades of a learning process, the division can deal with the challenge of a larger-scale exploitation with a decreasing risk.
• The definition of Geotechnical-Geomechanical Criteria and Parameters for Mining planning based on the environment where the productive areas are located. They include:
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mining sequence (caving starting zone and area expansion); and - undercutting and draw rates. - Undercutting rates, should be defined per unit of the active front, allowing the new area incorporation rate as the replacement of the exhausted area. This parameter also is involved in the seismic response of the rock mass due to the mining activity. An average rate of 2000 3000 m2/month has been set. - Differentiated extraction rate up to 30 per cent of the solid primary rock column has been reached. The extraction rate must allow keeping the seismic activity in an admissible range up to the complete breaking of the solid primary rock column. Empirically, it has been estimated that the solid primary rock column is completely broken after an extraction height having reached 30 per cent of the solid column. The geomechanical restriction for extraction rates is an increasing function of the primary ore extraction percentage due to the fact that the induced seismicity moves upwards as the extraction increases. Rates of 0.25 tonnes/m2 to 0.6 tonnes/m2 have been set.
• Caving angles (overhang and breaking); • Safety stripe (undercutting limits regarding developments, construction and extraction); and
• Open area: The maximum open area size must be quantified considering that there is relation between this area and some geotechnical instabilities that have occurred in panel caving areas in a primary rockmass. Besides, a greater production area including a greater rockmass volume in a breaking process has a direct impact on the seismic response of the rockmass under the mining activity. There is a direct relation between the active area and the open area through the availability factors and the use of the production area. Therefore, improving (increasing) these factors will result in a reduced open area. Consequently, the corresponding active volume (volume of rockmass in a breaking process) can also be reduced with an improvement of the geotechnical-geomechanical condition of the area.
CAVING METHOD TENDENCIES Analysing the evolution of the caving exploitation system at El Teniente Mine and if a comparison with the current world tendency is set, we observe that there really have been innovations, both internally and externally. They lead us to establish that mining methods are changing in a constant search for an answer to challenges arising from the particular features of each ore deposit.
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For the El Teniente Mine, substantial change started in 1982, as the rockmass change from a less competent secondary rock to the hard primary rock. It was characterised by the introduction of greater mechanisation and new mining designs that allowed handling ore with a proper fragmentation and hardness. However, the greatest challenge is how to adapt this type of mining operation to the rockmass response. This response includes the induced seismic activity that has caused damages and losses in the normal productive process. The experience and the study of this phenomenon lead to a mining parameter control until recently unknown in the performance of underground mining, already mentioned and related to the mining design, the operational sequence and the production volumes. The El Teniente Mine has modified the conventional panel caving method by determining that the stresses induced by the undercutting produce damages, mainly to the infrastructure at the production level (draw points, dumping points, drift over excavation). A variant of the panel caving exploitation method was defined, called previous caving, which consists of sequencing the undercutting front ahead of the preparation of the production level (behind the front and under the undercut area). This variant of the panel caving method has been applied to a mine sector since 1996, with production starting in 1997. Today, almost 80.000 m2 have been undercut, achieving a production of 15.000 ton/day with 40.000 m2 of drawbells under production. A significant reduction of damage to the drifts located under the undecut level and rockburst occurrence has been observed associated with the passing of the undercutting front. Within this framework, El Teniente Mine has developed a new exploitation method for primary rockmass, called ‘macrotrench’, as an alternative method with a great potential for deep primary ore extraction.
MACROTRENCH METHOD General concepts The macrotrench exploitation method consists basically in starting the exploitation through a four level sublevel cave that begins from a central slot with retreat to both sides in order to leave a big trench around the initial slot. The retreat motion of the sublevel cave stops at a position that leaves the upper levels more advanced than the lower ones, defining a distribution of regular drawpoints from a plan view. The next stage begins the extraction in a regular way from these drawpoints in order to start the undercutting and the later caving at the block upper level. This method involves a major sublevel initial stage in order to initiate ‘caving’. The first extraction stage will include broken material from blasting and the second stage will correspond to the extraction of naturally fragmented material (see Figures 4, 5 and 6).
Design description The macrotrench exploitation method is a combination of the sublevel and inclined caving methods. It consists of a working scheme arranged into four production levels plus one haulage level using conveyor belts. Undercutting and cave growth are produced using sublevel caving method from the four production levels, which are separated by 15 m vertically. The area over the upper level pillar is undercut from a conventional drift, located 20 metres above this level. Once undercutting has been completed, the extraction of the rest of the column is performed through the panel caving method with a regular draw mesh.
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FIG 4 - Isometric view of macrotrench.
Ore is extracted from the drawpoints located at 1st, 2nd and 3rd levels (enumerated in ascending order), with 13 yd3 scoops and with 7 yd3 scoops at the fourth level. These scoops unload the material into low profile crushers located one per level. The crushed ore is transferred through vertical ore passes to a conveyor belt level, located 30 metres beneath level 1, from where it is carried by the main transport system to the surface. Among the main design characteristics of the macrotrench method, the following ones can be mentioned:
• Greater strength of draw points: It means a greater availability of the drawpoints and therefore a lower repair cost. Besides, greater draw regularity can be obtained with this greater availability of the drawpoints.
• Wider drawpoints: With wider drawpoints, the handling of greater blocks can be improved and the hang up frequency can be reduced
• Regular draw mesh: It allows for greater draw ore efficiency increasing the ore recovery.
• Larger equipment: The use of 13 yd3 scoops increases the operations efficiency at the production level due to easier handling of larger blocks. The same equipment allows increased extraction rates at the upper levels compare to the panel caving rates.
• Use of the low profile crushers: The low profile crushers allow the reduction of bigger blocks than conventional crushers, with a simplified handling without the need for further excavations. Additionally, the handling of coarse ore
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with blocks between 12″ and 18″, reduces the ore pass damages.
• Greater global stability: Geomechanical analyses reveal that an adequately oriented macrotrench design in relation to the stresses and the structures is more stable than a panel caving design. The analysis of potencial unstabilities was performed using wedge analysis and numerical modelling was used for stress analysis.
Aspects to be considered • Mining design: As mentioned before, the design has four production levels plus one undercutting level and one haulage level using conveyor belts, which could also be used as ventilation exhaust and drainage drift. For this case, a 140 m x 130 m modulus of the macrotrench design is defined. The area to be caved is then 18 200 square metres (see Figure 7).
• Accesses: Connections between the different levels will be made by ramps with a 12 per cent gradient.
• Levels - Level 4 For this production level a 25 x 20 m pillar and a ‘fishbone type’ layout, with drawpoints spaced every 20 metres has been defined, thus setting a 400 m2 mesh for the level.
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FIG 5 - Macrotrench – initial stage.
FIG 6 Macrotrench – final stage.
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FIG 7 - Macrotrench plan view.
The drift section is 4.0 x 4.0 m and 7 yd3 LHD equipment will be used. The 400 tph capacity horizontal impact crusher is located at one end of the level.
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Level 3 This level is located 15 metres beneath Level 4 and incorporates the typical scheme for the macrotrench, with two independent production galleries connected by two crosscuts, one for communication and the other one for the crusher installation, located at one end. The production gallery section is 5.0 x 4.0 m, while sublevel caving load stabs and drifts are 4.5 x 4.0 m. In this level and the lower levels 13 cubic yard LHD equipment will be in use, dumping into the 600 tph horizontal impact crushers.
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Level 2 and Level 1 Level 2 and Level 1 are similar to Level 3, with two
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independent galleries and their communication and crushing crosscuts. Level 1 is used as fresh air injection system for the upper levels.
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Conveyor Belt Level The ore pass system consists of a series of vertical or subvertical galleries that connect crushing stations of the 1st, 2nd, 3rd and 4th levels with the conveyor belt level. These ore pass have a diameter of 1.8 m and the lengths ranging from 30 to 70 m, depending on the level they serve. The conveyor belt level has a 5.0 x 4.0 m gallery section.
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Undercutting level The undercutting level has a 3.6 x 3.6 m section, located 20 m over the Level 4 production drift. The drilling and later blasting are performed from this drift to produce the undercutting. A typical macrotrench block is show in Figure 8.
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FIG 8 - Development of a typical macrotrench.
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• The proposed design has less degree of interaction potential
Caving forms There are two types of caving involved in the macrotrench method, caving by sublevels and conventional panel caving. 1.
Undercutting by sublevels This type of undercutting will be used for the greater portion of the test area, drilling from the sublevel cave drift from all production levels. Although it is true that this undercutting methodology is known as an exploitation method, in this case the purpose is to perform a major basal undercutting, thus assuring an adequate fragmentation of the first metres of the rockmass column.
2.
Conventional undercutting Conventional undercutting will be used only to cave the area immediately over level 4 pillar. This undercutting will be based on the drilling of 64 mm (2 ½″) radius boreholes from the single undercutting drift.
Slot construction No previously mined area with an upper undercutting level will exist. With primary ore column heights close to 300 m, it is convenient to help the initial caving by including a slot in one side of the area to be caved. Figures 9 and 10 show the progressive stages of sublevel caving and Figure 11 shows the production stage.
Draw The draw study has developed an independent analysis of the extraction mesh proposed for the macrotrench method and a comparison with the 15 m x 20 m mesh of the panel cave. The main comments regarding this subject are:
than the one used by El Teniente, being in the so-called ‘isolated-interactive’ draw area.
• Increasing the distance between drawpoints to 18 m, pushes the design close to the interaction area limits. If the fragmentation is smaller than expected, it could fall in the isolated draw area.
• The particular features of the proposed design are a fewer than those included in the 15 x 17 m panel caving design. In the panel caving method, they reach 56.7 per cent of the area and in the macrotrench method only 18.8 per cent. This is clearly favorable, since a more regular geometry allows for a more homogenous draw and therefore better extraction results.
• The effect produced by a geometry of inclined drawpoints is more favorable from a draw point of view in order to reduce hang ups as well as to increase the interaction potential. This is further favored when using equipment with a greater capacity (13 cubic yard LHD).
• Regarding the extraction rate to be used, the design itself does not impose any other restriction than the geomechanical ones accordingly to the environment where it is implemented. Once these restrictions are eliminated, the method must bear the maximum possible extraction rate from the operational viewpoint, taking into account only the recommendation of a uniform draw. The results of an operation simulation reveal that this rate is close to 1.2 tpd/m2 as an average. Rates close to 2.2 tpd/m2 are reached at lower levels. The extrapolation of restricted extraction rates from the panel caving to the macrotrench method is considered to be very conservative for this new geometry, in which the volume of broken rock between galleries and the undercutting is greater for the same extraction percentage.
FIG 9 - Sublevel caving stage.
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FIG 10
Sublevel caving progressive stage.
FIG 11 - Production stage.
• Regarding the waste-ore angle, it is recommended to work with a 50º value, which has to be considered for mining planning.
• One of the main advantages of the macrotrench design is the
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strength of the drawpoints. Owing to this, it is estimated that an availability of drawpoints greater than in the panel caving areas favors draw uniformity and therefore, a greater reserve recovery.
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• The stresses and the wedge blocks with a smaller size allow
Material handling This item focuses its analysis mainly on the extraction, reduction, ore passes system, and the intermediate haulage activities. The main purpose of the analysed options for the material handling is to reach high production rates. For this purpose and according to geomechanical studies that allow the development of drifts with larger sections, loading and transport equipment of a high productivity and a larger size have been considered.In order to meet these requirements, 13 yd3 LHD equipment have been defined at levels 1st, 2nd and 3rd, which characteristics make them able to reach high productivity levels and to handle large blocks. Owing to the design restrictions, the use of 7 yd3 LHD is planned at the upper level (Level 4). After a detail analysis, including preparation, operation and investments costs, it was decided to incorporate low profile crushers inside the macropillar. Main foreseen advantages are:
• a high productivity of LHD equipment; • a flexible capacity of the system, which allows a potential production increase through a higher number of scoops;
• an efficient and economic intermediate transport system, due to the use of conveyor belt;
• it allows for its complete automation; and • a comparatively less investment level.
for the drifts with a bigger section that the panel caving.
• The orientation of the macrotrench should not affect the size of the wedge blocks to be formed.
Global stability • The analysed structural pattern does not produce wedge blocks that affect the global stability of the macrotrench system.
• In this design, diamonds shaped zones are produced which protect the drawpoints. The analyses reveal that the diamonds can lose part of their geometry due to the formation of wedge blocks, but even then they protect the drawpoint adequately. This could affect the drawpoint design, allowing for a lighter design, without such a heavy reinforcement.
• There is a possibility of wedge blocks formation at the macrotrench ‘Apex’, which would substantially reduce the pillar over it. This situation, in conjunction to a bad orientation of the macrotrench can seriously compromise the stability of the upper part of the macrotrench. This confirms the relevance of the macrotrench orientation and it recommends a more rigorous reinforcement of the upper level 4.
COMPARATIVE ANALYSIS OF MACROTRENCH VS PANEL CAVING
Geomechanical evaluation The geomechanical evaluation was analysed considering three main aspects: the stresses, the drift stability and the global stability. The most relevant comments are as follows.
The comparison presented between both methods is based on the operation and preparation costs. The operational aspects are not considered because they have been mentioned previously.
Stresses
Costs
• The stresses affecting a macrotrench design are very sensitive
Table 2 shows cost values for mine preparation and operation. Development, support, construction, drilling and blasting costs are included as preparation cost. Extraction, ore transport, secondary blasting, crushing and maintenance costs are included as operational cost.
to macrotrench orientation. By orienting the main axis of the macrotrench perpendicularly to the principal stress, the stresses affecting the tunnels are substantially reduced.
• In a general way, according to its geometrical arrangement, the macrotrench method accumulates less energy than the panel caving method.
• Through numerical modeling, a design for the macrotrench was analysed, oriented with the main axis parallel to the principal stress, which turned out to be the worst orientation. Within this framework, the differences in the behaviour comparing to a panel caving were not relevant. The stresses at the top and at Level 3 were similar or higher to the panel caving stresses. At the Level 1 and Level 2, they were smaller or equal to the panel caving situation.
TABLE 2 Cost values for mine preparation and operation. Item
Macrotrench
Panel caving
Preparation ($US/m2)
760
830
Operation ($US/ton)
2.8
3.3
• The second analysis was conducted with the macrotrench oriented perpendicularly to the principal stress. Smaller stress concentrations than in panel caving were founded, showing an approximately reduction of 60 to 80 per cent.
Tunnel stability A structural system has been extrapolated for the macrotrench area based on the observations from adjacent sectors. This has an impact on tunnel stability. The most relevant comments of this analysis are:
• The macrotrench method has a better drift orientation than panel caving and smaller wedges are produced. This is due to two effects, a smaller number of directions and the perpendicular intersection between the drifts.
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According to this cost information, it can be noted that in both items, the macrotrench cost are lower, especially regarding the operation costs with a 30 per cent difference between both methods. This difference lies mainly in the lower cost of the ore transport using belts, which compensated the higher cost activity of a previous crushing. Regarding the preparation costs, the difference is only ten per cent and therefore, this item must be considered similar for both methods, given the accuracy of the study engineering level. Finally, as a conclusion regarding the involved costs, the macrotrench looks more favorable than the traditional panel caving design.
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• The results obtained in this paper confirm the potential of
CONCLUSIONS
this new method and the need to continue with the industrial piloting stages of the method at the mine.
Advantages • The ‘macrotrench’ method offers enough comparative advantages and technical arguments to be considered competitive regarding the extraction of primary ore at El Teniente Mine. At this engineering level, the method offers lower costs and better stability conditions.
• There are operational advantages in favor of the ‘macrotrench’, which are summarised as follows:
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Higher draw rates Better global and local stability Better ore handling when using bigger scoops (13 cubic yards) Reduction of the quantity and length of ore passes. Better availability of the draw points, which means higher draw uniformity, with the subsequent benefits, such as a less dilution and a greater reserve recovery. Besides, the better availability of the draw points the higher draw rate should mean a smaller open area is needed, and therefore, a greater operation concentration with the corresponding advantages. This situation creates a highly productive method when compared with the alternative method of panel caving. The existence of draw points at different levels allows for a better handling of hangings up. This method allows the use of the transport and the ore reduction equipment in the production level.
• After the analysis of the panel caving and the macrotrench extraction meshes, its advantages completely justify accepting the disadvantages. It is estimated that the fact of having more robust and resistant drawpoints mean better availability, which allows for a more uniform draw and a smaller open area. These two factors lead to a better reserve recovery and a better operation concentration, generating a better global operation.
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Disadvantages • It has not been proven yet. • A non flexible orientation regarding the stresses, the structures and the upper Apex stability.
• An increase of the equipment capacity in order to set similar productions.
• Developments and production at more than one level. • The use of a undercutting methodology through sublevel •
caving. A high amount of developments in order to allow for production continuity.
REFERENCES Cavieres, P, 1999. Geomechanical report for El Teniente development plan. El Teniente Division internal report. Díaz, G and Córdova, N, 1986. Block Caving at El Teniente Mine, Journal of Chilean Mining Engineer Institute. NCL Ingeniería y Construcción SA, 1997. Study and analysis of alternative methods for primary ore exploitation at El Teniente Mine. Report for El Teniente Division. NCL Ingeniería y Construcción SA, 1998. Conceptual developments for exploitation using macrotrench. Report for El Teniente Division. Rojas, E, Cuevas, J and Barrera, V, 1992. Analysis of the wear in drawpoint at El Teniente Mine, in Proceedings Massmin 92. Rojas, E and Díaz, G, 1998. Know-how at El Teniente Sub 6, El Teniente Division Internal Report Tobar, P, 1998. Macrotrench: an alternative method, Journal of Chilean Mining Engineer Institute. Yevenes I, 1985. Macrotrench, El Teniente División Internal Report. Varas, F, 2000. Undergroung mining methods. Codelco-Chile División El Teniente, internal report.
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Evolution in Panel Caving Undercutting and Drawbell Excavation, El Teniente Mine J Jofre1, P Yáñez2 and G Ferguson3 • expanding of the production on 100 per cent during the last
ABSTRACT The El Teniente mine produces more than 340 000 tonnes of fine copper per year from an industrial complex of underground mine, concentrator and smelter. Mass caving methods are employed to deliver approximately 95 000 tpd of ore to the mill, from a number of sectors underground, each sector being in effect a large-scale mine in its own right. This mining activity has faced challenges over the course of the last century such as: • the capacity of the underground mine has been expanded to maintain and increase output of fine copper; • the increasing size of the ore fragmentation forced to change the mining method to achieve the high productivity and profitability required in the business; and • to modify the mining methods to continue mining at depth with safety. Fundamental activities in these mining methods, for the successful initiation of caving and subsequent draw, are: • undercutting the base of the block or panel by blasting; and • excavating of drawbells. The undercutting design evolves mainly on the height of the blasting to manage the change on ore fragmentation. The methodology move away from incorporating massive areas, typical of the block caving method, to a continuos blasting, typical of the panel caving method. Drawbell design changes on geometry, dimensions and sequence of excavation to deal with a corse ore fragmentation and the resolution of mining difficulties in the form of major rock bursts and collapses of production facilities. This paper focuses upon the evolution of the undercutting and drawbell design used at El Teniente mine and techniques under consideration to deal with future caving operations to provide the bulk of the daily production for the foreseeable future.
INTRODUCTION The ending of exploitation within the weak and well fragmented ore zone of secondary mineralisation (secondary ore) and the start-up of extraction in the much harder and poorly fragmented ore zone of primary mineralisation (primary ore), has forced El Teniente management to mechanise mining operations. In this way, it has been possible to achieve the high productivities and production rates required to maintain profitability. Such circumstances gave rise to new mining geometries and initiated the changeover from block caving to panel caving. Fundamental activities in both these mining methods, for the successful initiation of caving and subsequent draw, are the undercutting the base of the block or panel by blasting, and the excavation of drawbells. The mining of primary ore by panel caving began in 1982, and since that date, El Teniente has studied and incorporated a number of different methods and sequence of undercutting and drawpoint excavation. The changes were made with the aims of:
20 years;
• improving the safety of personnel; • maintaining production despite facing the closure of a major sector for a number of years brought about rock bursting;
• reducing the cost of mining and improved productivity; • minimising the risks associated with the formation of pillars through imperfect blasting;
• • • •
enhancing the flow of ore; improving the management of fragmentation; securing better crown pillar stability; and reducing the time and cost of preparing sectors for caving.
As the mine grade has fallen over the course of the last century, the capacity of the underground mine – and associated treatment facilities – has been expanded to maintain and increase output of fine copper. Considerable investment is obviously required to achieve such a substantial contribution to the world supply of copper. The main technical risks involved in approving expansions invariably focus upon the mineral resources and the mining plan proposed to deliver annually 36 million tonnes of ore to the metallurgical plant.
GEOTECHNICAL CHARACTERISTICS – EL TENIENTE DEPOSIT The bulk of mineralisation within the El Teniente deposit is typical of massive, homogenous copper porphyries. However, at some time during its history, the deposit has been affected by supergene alteration through percolation of meteorological water close to surface, which gave rise to secondary mineralisation. This secondary ore can be characterised as high in copper grade, but low in strength that produces good fragmentation and caveability. In contrast, primary mineralisation is low in copper grade, and is a high strength material that has only moderate fragmentation and caveability. Such contrasts can be clearly observed in Andesite, the predominant rock type within the El Teniente deposit (see Table 1, after Quezada, 1998). Caving experience in those types of ore has shown that the height of an undercut excavation can facilitate the initiation of caving, but must be considered in relation to rock mass characteristics and time.
UNDERCUTTING AND OPENING DRAWBELLS IN SECONDARY ORE Exploitation of secondary ore from El Teniente has taken place since start-up of operations in the early part of the last century, until the present time. Currently, the last mining sector located in secondary ore, Quebrada Teniente, has a rate of production of 22 000 tpd and is planned to shutdown in 2003 (Tobar, 1997).
1.
Geomechanics Consultant, 9 Cox St, St Albans Vic 3021. Formerly Codelco-Chile, Division El Teniente and IMC Consultores Ltda, Chile. E-mail: [email protected]
2.
Project Chief, Codelco-Chile, Division El Teniente, Chile, Colon Alto, Millan 1020, Rancagua, Chile.
Block caving
3.
Mining and Geomechanics Consultant, Seltrust Associates Limited, United Kingdom.
The block caving method has been used for the mining of secondary ore, using mining areas of from 3600 m2 to 9000 m2 and block heights greater than 100 m. There are five main levels
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TABLE 1 Principal characteristics - primary and secondary andesite - El Teniente. Parameter Specific gravity (t/m3)
Secondary Rock
Primary Rock
2.2 - 2.7
2.7 - 2.9
10 - 30
5-9
Fracture frequency (ff/m3) Fragmentation (m)
0.05 - 0.20
1-5
< 40
50 - 60
RMR Laubscher Uniaxial compressivestrength (MPa)
25 - 50
90 - 150
Porosity (%)
2 - 10
0.09 - 0.88
RQD
< 25
> 80
Characteristics
• Weak material, highly altered and well fractured, located close to surface
• Hard and dense material, located at depth
• Occurrence of collapses
• Occurrence of collapses
• No occurrence of rock bursts
• Occurrence of rock bursts
FIG 1 - Traditional block caving - undercut design.
utilised in block caving related to the following activities: undercutting, production, ventilation, rock handling and haulage. Undercutting is carried out some 8 m above the production level from drives, typically 2.4 m wide by 2.1 m high in cross section. Undercut heights evolved from 5.5 m to 9.1 m, which are considered to be low, but which gave a good result from the viewpoint of initiating and propagating caving in secondary ore. Thus, the undercut height could vary from 2.5 to 4.5 times the height of the undercut drill drive, or from six per cent to nine pre cent of the height of the block. Ore funnels, developed from the production level, connect to the overlying undercut level for the extraction of ore. Excavated funnel dimensions are typically 8 m in length and, in cross-section, either a circle of 3.6 m diameter, or a rectangle of
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1.2 m by 1.7 m. Given that the rock fragment size is less than 325 mm, the rate of draw through funnels is high. Block caving was adopted in a number of sectors, including Mina El Teniente, Teniente 1 Mina Sur, Teniente 4 Mina Norte, Teniente 3 Isla and lastly, Teniente 6 Quebrada Teniente (Pasten et al, 1999b). The design and layout of Quebrada Teniente is typical of traditional block caving (see Figure 1).
UNDERCUTTING AND OPENING DRAWBELLS IN PRIMARY ORE A corse ore fragmentation forced the change from block caving to a LHD mechanised panel caving. The LHD is the only mechanised system handling the primary ore at El Teniente mine.
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EVOLUTION IN PANEL CAVING UNDERCUTTING AND DRAWBELL EXCAVATION, EL TENIENTE MINE
The aspects of the design and layout of panel caving that have changed most since the inception of the method in 1982, are:
• the height of undercut, and • the sequence of mining the undercut in relation to the development of the production level and excavation of the drawbells. These differences are outlined in the following sections.
Traditional panel caving The principal characteristic of panel caving is the continuous excavation of the undercut level and the opening of the underlying drawbells. Thus, panel caving may be said to be a dynamic method differing fundamentally from the concept of block caving which encompasses a static or single undercutting operation. Once the undercut has been blasted in a block caving sector, it no longer plays a part in the caving operation.
The undercut level in traditional panel caving is located 18 m above the production level. Undercut blast holes, approximately 50 mm in diameter, are drilled from drives 3.6 m wide by 3.6 m high, so forming a void into which panel caving may be induced, Figure 2. Undercutting in traditional panel caving requires the complete development and construction of the undercut and production levels. Drawbells are opened before the passage of undercut blasting. Unfortunately, this sequence exposes the production level to high levels of abutment stress from the advancing undercut face, provoking damage in the production level crown pillars (JKRMC, 1993). The main characteristic of this method has been the evolution of the undercut drilling layout, which defines the height of the undercut excavation. Thus, the height has varied significantly over the years (see Figure 2 and Table 2, after Pasten et al, 1999a).
FIG 2 - Evolution of undercut height – traditional panel caving.
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TABLE 2 Undercut height – traditional panel caving.
should be remembered that the method of traditional panel caving induces the damage which brings about the collapse of significant areas, so interrupting production and demanding additional expenditure to recover the ore.
Year
Height (m)
1987
16.6
Pre-undercutting panel caving
1988
13.6
1991
10.6
1994
8.6 and 3.6
1998
10.6
One of the main reasons for introducing this method arose from the necessity to prevent the effects of the high abutment stresses associated with the traditional panel caving method. The pre-undercut method mitigates this effect by generating a distress shadow below the undercut. It is confidently expected that this no doubt contribute to a much lower occurrence of pillar failure in the future on the production level, as compared to the traditional method (Karzulovic, 1998 a and b). From the construction point of view, the main difference between the pre-undercut and the traditional panel caving is the sequence of development, undercutting and the excavation of drawbells (Cavieres and Rojas, 1985). As the name suggests, this variation of panel caving excavates the undercut prior to the development and construction of the production level. At El Teniente mine, a flat undercut is mined with the same height (3.6 m) as the undercut drives. Swell material from each undercut blast is removed by LHD to provide a free face for the next blast. Thus, two mining fronts have to be managed in this method, one on the undercut level, and the other located on the production level, which are related to the advance of the excavation of the drawbells and construction of the drawpoints. The undercut comprises drives, 3.6 m wide by 3.6 m high, developed parallel to each other on 15 m centres. The excavation of the undercut is achieved by blasting three or four hole fans, drilled into the sidewall some 7 m to 10 m in length. The drill holes are fanned slightly to ensure an undercut height equal to the height of the drives. Various blasting layouts have been attempted in the Esmeralda sector, the first area to implement the pre-undercutting style of panel caving (see Figure 3). Three working zones are defined by this method:
The trend of the undercut height is 3.6 m, but 10.6 m is recommended in the last studies and applications for traditional panel caving. A reduction of the undercut height has a number of advantages:
• less drilling, lower explosives consumption; and • good visibility to view the results of blasting. On the downside, a low undercut height:
• is less favourable for the initiation of caving; • provides less volume into which the cave may expand; and • produces a coarser initial fragmentation and oversize blocks. Consequently more secondary blasting is required initially which may damage the drawpoint brow, support and floor, and so suffers frequent interruptions to production during the start-up of caving. In accordance with the experience of the largest panel caving sector, Teniente 4 Sur, the most successful undercut height, on average, has been 10.6 m above the undercut level floor. The main sectors which have adopted traditional panel caving are Teniente 4 Sur (B, C and D), and Teniente Sub 6, with to-date, undercut plan dimensions of approximately 347 000 m2 and 37 000 m2, respectively. Experience of the method has demonstrated operational flexibility with the ability to make changes to resolve difficult situations. For example, the collapse and loss of the production level over a considerable area was overcome by developing an additional level below the production level designated SNV Teniente 4 Sur Panel 1 and 2. However, it
• the undercut zone; • the preparation zone on the production level (maintained at a distance of 22.5 m from the undercut front); and
• the production zone located some 45 m to 60 m behind the undercut front.
FIG 3 - Undercut blasting configurations – Esmeralda sector.
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Owing to the destress shadow, excellent conditions prevail on the production level in the pre-undercut panel caving sectors. However, this method also has difficulties, such as:
• minor damage to access drives in the undercut level; • cut-offs in radial drill holes; • increased percentage of swell removal (30 per cent to 40 per cent of the volume blasted);
• problems with opening drawbells; and • significant interference with respect to the co-ordination of the three working zones. Pre-undercutting has been applied in El Teniente for some eight years in minor sectors on a trial mining basis and in the Esmeralda project since 1997, with successful results. The main sectors utilising this method are Teniente 3 Isla (see Figure 4) and Teniente Sub 5 Esmeralda (Yáñez et al, 1998), with approximately 15 000 m2 and 72 000 m2 undercut to-date, respectively.
Drawbell design in primary ore Changes to the sequence of initiating panel caving has led to changes in drawbell geometry and excavation. In traditional panel caving, there are two stages of drawbell construction: partial development from the production level to a height of some 12 m to 15 m vertically above, followed by final construction from the undercut level by means of downholes, as shown in Figure 5. In the case of pre-undercut panel caving, drawbell excavation must be performed solely from the production level (see Figure 5). It is thus necessary to create a free face or slot, drilling and blasting holes 14 m to 18 m in length and fully charged in order to break through to the closed undercut level above. Such confined blasting is inefficient and, in some cases, can damage the crown pillar and drawpoint. The design of drawbell has encompassed changes in height, length, width, geometry, spatial location and method of excavation, mainly to assist layout changes with respect to:
FIG 4 - Teniente 3 Isla design layout.
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FIG 5 - Evolution of drawbell design.
• • • • •
orientation of drives; old workings; access cross-cuts; drawpoint configuration; limits of mineralisation; and finally
for improvements on down hole blasting design and variations in the system of exploitation, which has maintained the difference of 18 m between the undercut and production levels (see Figure 2). It should be noted that between 1985 and 1994, numerous ‘special’ drawbell designs were implemented, mainly in Sector Teniente 4 Sur. Their main goals were to deal with many specific differences in the mining design, such as: access cross-cuts, change in direction of drives and drawpoint cross-cuts, old workings, and limits of mining areas (see Table 3).
OPTIONS FOR FUTURE PROJECTS IN PRIMARY ORE Engineering studies for the mining of future sectors are based upon the best experience found in both El Teniente and worldwide. With respect to undercutting, the most successful technique is that practised in the Esmeralda sector. The condition of the production level rock mass in the shadow of the pre-undercut level is greatly improved. However, engineering studies continue to search for ways to improve current methods, focussing upon undercutting and the excavation of drawbells. Technically viable options for implementation in future projects are outlined in this section.
Panel caving – advanced undercutting This variation of panel caving is applied in other mines, one of the best known is Northparkes mine (Dawson, 1995). It requires the development and partial construction of the production level prior to the excavation of the undercut (NCL, 1994). Once the production level is constructed, a narrow height undercut is mined, followed by the excavation of drawbells from the production level. Three or four blasting phases are needed to completely open the drawbells, in a similar fashion to the pre-undercut method of opening drawbells. The method derives its name from the fact that the undercut is mined some distance in advance of the drawbells, rather than in advance of the production level (see Figure 6).
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A zone of abutment stress is generated by the advanced undercut, but is predicted to have less effect upon the pre-developed production level cross-cuts below than in the case of traditional panel caving. However, the method will suffer from two disadvantages: the passage of the undercut front will subject the pre-developed production level to high stress levels, and the same drilling and blasting difficulties experienced in the pre-undercutting method of Panel caving will be encountered in opening drawbells from the production level. In contrast, the method should not experience the high levels of interference that exist in the pre-undercut method between the development and construction of the production level and the excavation of the pre-undercut.
Panel caving – SLC undercut and continuous trough drawbell Engineering work for future projects is considering the design of panel caving with a sublevel caving style of blasting for both undercutting and opening of drawbells in the form of a continuous trough (see Figure 7, after IMC, 1999). Continuous blasting is the main advantage, from an operational viewpoint, of the proposed sublevel caving excavation of the undercut and draw troughs, particularly with respect to the present pre-undercut method. A sublevel caving style of blasting is in strong contrast to the excavation of single drawbells. The latter, presents drilling difficulties and confined blasting conditions for the large volume of rock to be removed. Another advantage is that blasted material flows towards a central drive, as opposed to having to remove the swell material from below blasted material, as is the case with the narrow pre-undercut configuration of panel caving. The continuous drawbells scheme under investigation plans to maintain the sequence and distance implemented in the Esmeralda style of pre-undercut panel caving. Nevertheless, these factors will be revised in the next stage of engineering design with respect to the re-distribution of stresses and the co-ordination of undercutting, construction and production activities. The 30 m high sublevel caving drilling and blasting pattern would result in a significant increase in the volume of finely blasted material during the start-up of caving operations, compared to current undercutting methods. A positive impact is thus expected with regard to the initial fragmentation and progress of caving.
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TABLE 3 Development of drawpoint design – panel caving.
The geometry of the proposed design encompasses two continuous walls, whereas current practice is based upon four walls. Thus with the proposed design, early interaction is likely between drawpoints located in the same production drives, whilst the same level of interaction should be maintained between drawpoints in the opposite direction. Consequently, the proposed design should have a favourable impact on the progress of initial caving. Further, although a solid wall would not exist between adjacent drawpoints, it is surmised that fewer hang-ups would occur between neighbouring drawpoints. Provided that the design proves to be stable geomechanically, it has the potential to deliver high productivity using fully automated equipment. Recent engineering work has suggested that this SLC undercut style should take into account the characteristics of the flow of fragmented material, in order to define better the height of the
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undercut excavation and crown pillar. Current and past layouts have maintained the crown pillar height more or less constant. The work of Kvapil (1992) and Hustrulid (1999) requires a knowledge of the ellipsoid of movement of broken material, the angle of draw, width and height, which in turn are influenced by the configuration of the drawpoint horizon. A key objective of any design is to achieve the greatest interaction possible between draw ellipsoids located over the crown pillars, taking into account:
• the characteristics of the broken material; • the production level layout – spacing and angle of the production drives and drawpoint cross-cuts;
• the crown pillar height; and • the height of the undercut excavation.
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FIG 6 - Advanced undercut panel caving.
Considering this approach, it has been found that this SLC undercutting design, based upon the classic Teniente production level layout (ie drawpoint cross-cuts and production drives spaced on 20 m and 30 m centres respectively, angled at 60°), should be optimum with a crown pillar height of 18.0 m and an undercut height of 8.5 m. This estimation takes into account the fragmentation characteristics of primary ore and a wider draw point. Such a configuration encompasses a crown pillar height some two-thirds of the height of the extraction ellipsoid. If, however, the layout of the drawpoint cross-cuts and production drives were to be configured at right angles, (maintaining the same 20 m and 30 m spacing), a 30 m diameter ellipsoid of loosening is required to achieve full ellipsoid interaction. In this case, the height of the crown pillar could be set at 19 m and the height of the undercut at 10.5 m. Engineering work is at present
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investigating the many possible variations and configurations that arise from the application of Kvapil’s work on material flow and the practical experience of El Teniente mine (see Figure 8). Although the proposed SLC and continuous trough method appears advantageous from both technical and operational viewpoints, doubt exists with respect to the stability of the production level. In order to develop a continuous trough, it would be necessary to excavate a drive between the main production drives which would cut the pre-undercut pillar in half, so removing 30 per cent more solid rock, in comparison to present designs. This has to be compared with a lower load and better rock mass properties than the traditional panel caving pillar. Obviously, major engineering studies will be needed to evaluate the proposal, together with operational trials.
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FIG 7 - SLC undercut and continuous trough drawbell.
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FIG 8 - Estimation of the extraction ellipsoid and crown pillar height.
Front caving, a similar technique to the continuous trough method, has been used in El Teniente to recover pillars from a single level. Approximately 50 000 m2 have been mined successfully by front caving. Other examples worldwide, include the Radenthein mine in Austria (Weiss, 1981), Premier mine in South Africa (Owen, 1981; McMurray, 1979) and the San Manuel mine in the United States of America (Jackson, 1978).
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Single level SLC panel caving El Teniente mine designed this variation for Sector A1 at Teniente 4 level, to extract an ore column comprising 40 m to 60 m primary ore overlain by secondary ore. In this variation of panel caving, undercutting would be carried out from the production level, via a continuous trough, excavated by drilling fans of longholes, in a similar fashion to sublevel caving, as shown in Figure 9, after IMC, 1999.
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FIG 9 - Single level undercut and continuous trough drawbell.
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Following behind the undercut front, production drives would be developed in the pillar formed between the drives used for undercutting. Drawpoints would be excavated from the pillar drives, breaking against the blasted material of the undercut, or until the roof stability indicates that breakthrough has been achieved. The main disadvantage to this method is that undercutting cannot be visually confirmed and, at the same time, pillars are minimised over the future production drives, so reducing stability. However, the variation has the significant advantage of eliminating the undercut level, so reducing the initial cost of development and should speed up the start of production. Of course, any stability problems experienced would potentially negate these apparent advantages.
CONCLUSIONS The evolution of dynamic mass caving methods, over the 18 years period from 1982 has been described. Such developments have been possible, against the backdrop of the collapse of significant areas of the production level, severe rockbursting and the need to deliver tens of millions of tonnes of mineral to the plant every year. It is clear that El Teniente is determined to seek further improvements in panel caving in order to meet the twin challenges of declining copper grade and metal price and that they will in all probability succeed.
REFERENCES A Karzulovic and Asociados Limitada, 1998a. Altura de Socavación: Consideraciones Geotecnicas, in Internal report DT-CG-98-001 submitted to El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). A Karzulovic and Asociados Limitada, 1998b. Evaluación Geotécnica Métodos de Socavación Previa y Avanzada, Mina El Teniente, in Internal report DT-CG-98-003 submitted to El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). Cavieres, P and Rojas, E, 1985. Hundimiento Avanzado: Una Variante al Método de Explotación de Hundimiento por Paneles en Mina El Teniente, Internal report El Teniente mine, Codelco-Chile, División El Teniente, Estudios y Métodos Operacionales (in Spanish). Dawson, L, 1995. Developing Australia’s first Block caving operation at Northparkes mines – Endeavour 26 Deposit, in Proceedings Underground Operator’s Conference, pp 155-164 (The Australasian Institute of Mining and Metallurgy: Melbourne). Hustrulid, W, 1999. Contribution to the Pre-study ‘Drawpoint configuration for large LHD machines’ at El Teniente, in Internal The Itasca Consulting Group report dated November 10, 1999 submitted to El Teniente mine, Codelco-Chile, División El Teniente (in English).
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IMC Consultores Limitada, 1999. Proyectos de Explotación Pipa Norte y Diablo Regimiento – Ingeniería Conceptual, Vol I a V, in Internal report dated Diciembre 1999 submitted to El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). Jackson, D, 1978. Block caving keeps San Manuel competitive with neighbouring open-pit copper mines, Engng Min J, 179, (6):127-136. JKMRC, 1994. Project P93E. Research performed for El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). Kvapil, R, 1982. The Mechanics and Design of Sublevel Caving Systems, Underground Mining Methods Handbook, Chapter 2, AIME. Kvapil, R, 1992. Sublevel Caving, Mining Engineering Handbook, Chapter 20.2, AIME. McMurray, S, 1979. The design of a mass mining system for use below the grabbo sill at Premier mine, in Circ Assoc Mine Managers S Afr, no 4/79, 29 p. NCL Ingeniería y Construcción SA, 1994. Proyecto de Explotación Esmeralda – Ingeniería Básica, in Internal report dated July 1994, submitted to El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). Owen, K, 1981. Block caving at Premier Mine, in Design and Operation of Caving and Sublevel Stoping Mines, Chapter 15, AIME. Pastén, O and Cuevas, J, 1999. Altura de Socavación en Mina El Teniente, in Internal El Teniente mine report PL-I-023/99, Codelco-Chile, División El Teniente (in Spanish, May 1999). Pastén, V O, Pastén, M O, Medrano, S and Cuevas, J, 1999. Estudio Diseños de Zanjas y Socavación en Mina El Teniente, in Internal El Teniente mine report PL-I-069/99, Codelco-Chile, División El Teniente (in Spanish, November 1999). Quezada, O, 1998. Antecedentes de Geología y Geotecnia Proyecto Diablo Regimiento, in Internal El Teniente mine report GL-286/98, Codelco-Chile, División El Teniente (in Spanish, November 1998). Subterra Ingenieros Limitada, 1993. Análisis de Alternativas de Métodos de Explotación para Mineral Primario con Restricciones Geomecánicas, Volumen I y II, in Internal report dated Enero 1993, submitted to El Teniente mine, Codelco-Chile, División El Teniente (in Spanish). Tobar, P, 1997. Métodos de Explotación, Mina El Teniente, paper submitted by CODELCO-Chile, División El Teniente to the I Congress of Mining Engineering, University of Antofagasta, Octubre de 1997 (in Spanish). Weiss, P, 1981. Development System for Block Caving Under Severe Conditions, in Design and Operation of Caving and Sublevel Stoping Mines, Chapter 12, AIME. Yáñez, P, Pastén, O, Molina, R, Jofré, J, Cuevas, J and Rojas, E. Análisis de Inicio y Propagación del Caving, Mina Esmeralda, in Internal El Teniente mine report PL-I-066/98, Codelco-Chile, División El Teniente (in Spanish, October 1998).
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The Pre-Undercut Caving Method at the El Teniente Mine, Codelco Chile E Rojas1, R Molina1, A Bonani1 and H Constanzo1 CONVENTIONAL PANEL CAVING AND PRE-UNDERCUT CONCEPTS
ABSTRACT With conventional undercutting sequence the infrastructure on the production level often has a short serviceable life before major rehabilitation is necessary and is characterised by high maintenance costs. The abutment stress zone is the main cause of this situation. The application of the pre-undercut caving method variant can increase the availability and utilisation of the production level When a pre-undercut is used development and construction are carried out underneath a previously undercut area, far away from the abutment stress zone, under more favourable stress conditions. This paper describes the design principles and the relevant results from the application of the pre-undercut variant over three years to the Esmeralda sector of El Teniente mine. The aspects described are the undercut-extraction, the undercut-development and construction distances, the geometry and drawbells opening, cave extraction rates, and the rate and height of undercutting.
INTRODUCTION Since 1982, the El Teniente Division of Codelco-Chile has been mining in the primary ore on the lower levels. The mechanical properties of the primary ore greatly differ from the secondary ore, which was mined during the exploitation of the upper levels of the orebody. This mining has been carried out mainly using the panel caving method (Rojas and Cavieres, 1993), which was applied for the first time on the Teniente 4 South level. Extensive experience and knowledge has been gained from the mining of some 200 Mt in sectors with wide and dynamic caving fronts under high stress conditions (40 - 60 MPa). Among other relevant factors, the moving caving front has been identified as the main source of damage to the tunnels and infrastructure on the lower levels. Therefore, a series of studies and controlled tests has been carried out with the aim of improving significantly the working conditions on the production level and to increase productivity. The improvement has been achieved through the introduction of an innovation to the panel caving undercut sequence, called ‘pre-undercut’. In essence, this innovation uses an advanced undercutting front in relation to the development and preparation of the production level. Thus, the damage to the tunnels and infrastructure due to the abutment stress zone generated by the undercut front can be avoided. The pre-feasibility studies for this new variant were initiated in 1982. Different test scale applications were carried out in some mine sectors and under different conditions. In 1992, the conceptual engineering stage of the new Esmeralda sector was initiated. The planned rate of full production is 45 000 tpd. The sector started production during 1997 and at present, 80 000 m2 have been undercut with a total production of more than 5 Mt. The planning-operational parameters and the relevant design aspects of the ‘pre-undercut’ variant are evaluated in this paper. They include the following points: extraction and undercutting rates, draw bell geometry, allowable distances, support, ore passes, production level layout and undercut height. 1.
Mining Engineer, Codelco Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile. E-mail: [email protected]
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The same basic concepts apply to both variants. The main difference is the sequence of each of the operational elements. The conventional panel caving method has been widely analysed in the technical literature (Rojas and Cavieres, 1993). The following sequence of activities is implemented (Figure 1a):
• • • •
development of tunnels on each level (1), drawbell opening (2), undercut blasting (3), and extraction (4).
The activity sequence has been modified for the pre-undercut variant (Rojas, Molina and Cavieres, in press). The undercut is excavated first and the production level is developed subsequently within a relaxed zone. Thus, the new sequence is as follows (Figure 1b):
• • • • •
development of the undercut level tunnels (1), undercutting blasting (2), development of the production level tunnels (3), drawbell opening (4), and extraction (5).
It can be see in Figure 1a, that the first activity of a conventional undercutting method is to develop the drifts corresponding to the production level and the undercut level. In contrast, the production level drifts are developed following the blasting of the undercut level in the pre-undercut variant. The main challenges associated with this variant involved the undercutting. A flat, low height undercut was decided upon (narrow undercut). The drawbells can be developed only from the production level underneath and over a greater height, connecting to the pre-excavated undercut zone. The described sequence requires that working zones are defined according to the specific operations at each stage of development, construction and drawbell excavations (Figure 1b). The development of the production level under a pre-excavated undercut area is the essential difference in the pre-undercut variant applied to the Esmeralda sector. The other levels are developed in the same way as with conventional panel caving. Some studies have identified a significant stress decrease under a previously undercut area. The pre-undercut variant employs this destressing effect to protect the production level. In this way, a high standard production level is achieved, with improved working conditions and lower risks. In the conventional variant, the production level is prepared in advance of the moving undercutting front and so it is adversely affected by the passing of the abutment stress zone. Improved conditions in the production level is the main advantage of the pre-undercut variant due to better rock mass conditions resulting from the development of the production openings in a destressed area (ie as defined by working zones: pre-mining, transition and relaxed, Rojas, Molina and Cavieres, in press).
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(1a)
CONVENTIONAL PANEL CAVING PRODUCTION ZONE
PREPARATION ZONE 70 m. 3
1 2 1
4
(1b)
PRE UNDECUT UNDERCUT ZONE
PRODUCTION ZONE
80 m. 57.5 m.
PREPARATION ZONE 22.5 m.
2
1
4 5
3 PREPARATION ZONE FIG 1 - Conventional panel caving and pre-undercut activity sequences.
Induced seismicity The induced seismicity affecting the Esmeralda pre-undercut variant is interpreted in similar terms to those presented in the paper titled ‘Control on induced seismicity at El Teniente Mine’ also presented at MassMin 2000. Firstly, as the production level is developed underneath the undercut area, the broken material within the narrow undercut acts as a limited shield against the radiated energy from seismic events. This is a favourable condition which results in less damage to the production level due to induced seismicity.
ESMERALDA SECTOR DESCRIPTIONS Geology and stress condition The main lithological units are a dark massive subvolcanic andesite, highly fractured and hosting the main part of the mineralisation, an hydrothermal breccia complex (igneous and anhidrite breccia) and some intrusive bodies corresponding to diorites and latites intersecting the andesites (Rojas, Molina and Cavieres, in press).
Mine design The design and layout geometry incorporated several aspects aiming to improve the rock mass mechanical behaviour and stability. The main aspects included were the drift orientation with respect to the stresses and the structural condition of the rock mass.
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The distance between the floor of the undercut level and the production level is 18 m. This gives a crown pillar thickness of 14.4 m. The spacing of the undercut level and production level drifts is 15 m and 30 m, respectively (Rojas, Molina and Cavieres, in press).
Production level support The support is implemented as development support and permanent support. The development support is installed just after the development of the drifts. To complete the designed support, the final support is installed 15 m behind the development face. Table 1 gives the support details for both stages. TABLE 1 Rock support production level. Description Development: Fully grouted rebar (φ 22 mm), chain-link mesh (10 006), pattern 0.9 × 1.0 m (long 2.3 m), shotcrete 10 cm Permanent: Fully grouted long cables bolts (φ 15.2 mm), (intersections drifts/crosscuts) Steel sets and birdcage cables (drawpoint) Confining cables (zunchos) and straps (wall pillars)
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Planning and production parameters
Stress measurement
Extraction rate The extraction rates are defined so as to try to control the rate of breakage of the solid column. As a result, a more favourable seismic response is expected. Empirical evidence has shown that continuous caving of the full column height occurs when the height of collapse above the undercut back reaches 30 per cent of the height of the solid column, Table 2.
Thirty eight stress measurement cells (hollow inclusion) were installed on different levels of the Esmeralda sector. Of these, ten per cent were installed in the advanced access on the production level (Rock engineering group, 1999).
Tomography Four production level pillars and three undercut level pillars were monitored to identify the rock mass conditions.
Borehole camera
TABLE 2 Extraction rates.
A borehole camera was used to verify the rockmass condition of the same pillars through 2.5″ to 6″ diameter boreholes.
Primary ore extraction ranges (%)
Initial caving (m/d)
Steady caving (m/d)
0-5
0.05
0.10
5 - 10
0.07
0.13
10 - 15
0.08
0.15
15 - 20
0.10
0.17
20 - 25
0.13
0.20
25 - 30
0.16
0.24
TDR (Time Domain Reflectometer) Four TDR sites were installed, with 600 metres of cable for monitoring the response of the cave back to undercutting and extraction (Rock engineering group, 1999).
Damage surveys These were carried out to determine the causes of the damage, associated with the rock mass condition and the support response.
RESULTS
Undercutting rate A maximum allowable area for the undercut has been defined according to the same concepts. The maximum area is 36 000 m2/year, pro rata a similar monthly rates, with an undercutting height of 3.6 m (Planning group, 1999).
Permissible distances The permissible distances are based on the different rockmass conditions (Rojas, Molina and Cavieres, in press), and they correspond to:
• undercut front/extraction distance: 80.0 m; • undercut front/development distance: 22.5 m.
Stress conditions The stress condition in the production level in relation to the undercut front is presented in Figure 2, using the data collected from 11 stress monitoring cells. The major principal stress (σ1) has a maximum value of 90 MPa, decreasing to less than 20 MPa below the undercut area (Rock engineering group, 2000). The maximum shear stress component Figure 3, increases from the in situ value of 10 MPa to a peak of 26 MPa close behind the undercut front. Then it decreases to 4 MPa below the undercut area. The maximum shear stress is defined as: δ = (σ1 − σ3)/2
The relations between the development-construction-extraction limits are a function of the way in which the area is prepared.
Mine design Production An average production of 12 500 tons/day has been reached during the period January - May 2000, with a peak of 15 000 tpd and an available area of 45 000 m2.
Monitoring program A monitoring program was implemented in order to measure the rockmass response to the pre-undercut variant. This program included:
A global seismic instrumentation Four new seismic stations located in the Esmeralda sector were included in the mine-wide ISS seismic system, increasing the total number of stations to 25.
A local seismic system A local seismic system (eight stations with accelerometers as sensors) was installed in the initial Esmeralda area.
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The mine design (layouts and geometry) of the pre-undercut variant is substantially similar to the conventional panel cave variant. The exceptions are the drawbell construction (underneath the undercut area with complete construction from the production level) and the undercut height (narrow).
Tomography This technique has been used in the conventional caving sectors (Teniente 4 South and Teniente sub 6), and the pre-undercut sector (Esmeralda). It has identified a range of different rock mass conditions for similar situations. The images in Figure 4a), show pillar tomography results under a different development sequence (Rock engineering group, 1998, 2000).
Borehole camera Several borehole camera surveys were made in order to complete the tomographic information regarding the rockmass condition. They confirm the different rockmass conditions under similar situations. The images in Figure 4 b), present the pillar survey results under different development sequences.
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Stress [Mpa]
FIG 2 - Stress changes in relation to the undercutting front (adjusted graphics).
FIG 3 - Maximum shear stress changes in relation to undercutting front.
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THE PRE-UNDERCUT CAVING METHOD AT THE EL TENIENTE MINE, CODELCO CHILE
FIG 4 - Panel caving variants comparitive damages.
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• the tomography results show higher values of rock mass
Damage surveys The surveys carried out on critical infrastructure, ore passes, pillars, drawpoints and support elements have identified the following.
P-wave velocities for the pre-undercut variant than in the conventional variant corresponding to better rock mass conditions;
• the borehole camera images show minor deformation in the pillars using the pre-undercut variant; and
Ore pass system Different degrees of deterioration of the ore pass system have been shown from the damage surveys to the ore pass system due to over breaking (Figure 4c).
• the damage surveys identify better stability conditions for the ore passes and drawpoints with the pre-undercut variant.
CONCLUSIONS
• In Teniente 4 South sector, the analysis of the ore pass information indicates that the onset of over break occurs after an average value of 175 kt of ore has been transferred. The stability of the production level requires that the damaged ore passes be repaired or abandoned.
• Up to the present date, the ore transferred through the ore passes at Esmeralda has varied from 160 to 700 kt, without any over breaking being identified. These ore passes, without branches, go directly to the transport level in a different layout to the Teniente 4 South ore passes.
Pillars The information collected from the instrumentation, field data and the observation of pillar behaviour is noticeably different between the two variants. In the pre-undercut case, the rock mass condition is good whilst for a conventional undercut, the rock mass condition is poor, requiring repairs and even causing collapses.
The pre-undercut variant allows the production level to be developed in a destressed environment, resulting in the following benefits: greater availability, better utilisation and lower associated costs due to less repair and the loss of the production area. Currently, two aspects are the subject for improvement: the undercutting and its interaction between development activities on the production level. In relation to, the size of the undercut level pillars is being reviewed, trying to keep an appropriate size of cut beneath the column whilst minimising the extraction of broken rock. Regarding the development activities, work in underway to achieve better coordination and scheduling.
ACKNOWLEDGEMENTS The authors thank the Mine Planning Superintendency, El Teniente Division, CODELCO Chile, for authorisation to publish this paper.
Drawpoints
REFERENCES
The damage surveys show different levels of deterioration for the drawpoints and brow wear between the two variants (Rojas, Cuevas and Barrera, 1992).
Quantitatively • 99 per cent of the Esmeralda drawpoints do not show any damage associated with mining (drawbell blasting and extraction), for a range of extracted tonnage from 1000 to 39 000 tonnes (Rock engineering group, 1999).
• The greatest drawpoint damage has been observed for extractions ranging from 1000 - 30 000 tonnes in the Teniente 4 South and Teniente Sub6 sectors which employ a conventional undercut.
Support No support deterioration has been observed in the production level for the pre-undercut variant due to the low stress environment compared to the conventional variant. This is accompanied by less damage to the rock mass due to the development of this level underneath the undercut area. The geological discontinuities are disturbed to a limited degree, which permits minor water infiltration and less significant support damage.
DISCUSSION OF THE RESULTS The basis of the design of the pre-undercut variant - the stress decrease underneath the undercut area - has been confirmed by the results of the monitoring program. A significant improvement in stability conditions has been observed on the production level.
266
Geomechanics group, 1993 and 1994. Geomecánica conceptual proyecto Esmeralda and Análisis geomecánico ingenieria basica proyecto Esmeralda, El Teniente Division - Codelco Chile, Internal report. Planning group, 1999. Caso base vigente 2000, unidad de gestión autónoma mina-concentrador, mina El Teniente - Codelco Chile internal report. Rock engineering group, 1998. Evaluación geotécnica métodos de socavación previa y avanzada mina El Teniente. A Karzulovic and Assoc report for El Teniente Division - Codelco Chile. Rock engineering group, 1998. Tomografía sísmica de pilares en el nivel de producción del sector sub 6 Geoexploraciones SA, report for El Teniente Division - Codelco Chile Rock engineering group, 1999. Distancias permisibles panel caving con socavacion avanzada proyecto Reservas Norte, A Karzulovic and Assoc report for El Teniente Division - Codelco Chile. Rock engineering group, 1999. Evaluacion de daños en puntos de extracción mina Esmeralda, El Teniente Division - Codelco Chile, Internal report. Rock engineering group, 1999. Resultados de instrumentación mina Esmeralda, El Teniente Division - Codelco Chile, Internal report. Rock engineering group, 2000. Estudio de tomografía sísmica en pilares sector esmeralda y ten-4 sur, Geoexploraciones SA report for El Teniente Division - Codelco Chile. Rock engineering group, 2000. Postevaluación variante hundimiento previo mina Esmeralda, El Teniente Division - Codelco Chile, Internal report. Rojas, E and Cavieres, P, 1993. Hundimiento avanzado: una variante, al método de explotación de hundimiento por paneles en mina El Teniente – Codelco Chile, Volumen II (Proceedings 44-IIMCH Convention: Rancagua). Rojas, E, Cuevas J and Barrera V, 1992. Analysis of the wear in drawpoint brows at the Teniente Mine - Codelco Chile, in MASSMIN 92, pp 303-310. Rojas, E, Molina, R and Cavieres, P, in press. Pre-undercut caving in the Teniente Mine - Codelco Chile, SME Handbook 2001.
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Esmeralda Mine Exploitation Project M Barraza1 and P Crorkan1 ABSTRACT Within the framework in which future exploitations for Codelco-Chile División El Teniente are situated, regarding the deepening of its reserves and the geotechnical problems that affect their recovery, it becomes strategic to evolve with new variants, which make deep mining possible. Codelco Chile’s División El Teniente is progressing within its long-term planning the exploitation of a sector called Esmeralda, which includes its removal through the gravitational caving method (panel caving). It is planned to use pre-caving, which basically consists of developing and undermining the undercut level, without having developed the infrastructure of the production and transport levels. This is carried out in order to achieve a better redistribution of the stresses, thus obtaining higher quality in the infrastructure development of these levels as well as less risk. The extraction reserves for this area, estimated through the opportunity cost selection criteria, were determined at 348 million tons with an average grade of 1.014 per cent Cu. The project carried out geological and geotechnical studies in order to prevent the geomechanical risks of the sector to be exploited, which is a site conformed totally by primary ore. Lithologically, it occurs mainly in Andesite, located at the western limit by Brecha Braden Pipe. There are also varieties of porphyries and hydrothermal breaches of anhydrite. The Esmeralda Mine production program will increase from 4000 tpd (average) in 1998 to its 45 000 tpd rate of production estimated for the year 2005. The approved investment for the period 1995 to 1999 is $US 150 million, out of a total $US 213 million for the project. At feasibility study level and before the project, it showed a Net Present Value of $US 342 million, an Internal Rate of Return of 37 per cent and an IVAN indicator of 2.3 for a ten per cent discount rate and a copper price of 100 US cents/lb.
INTRODUCTION Currently, División El Teniente Mine-Concentrator phase produces 98 000 tpd of ore, from which 22 000 tpd are treated at Sewell mill and 76 000 tpd at Colón concentrator. Within the long-term planning context, the Esmeralda Project planned the beginning of its production for the second semester of 1997 with 250 tpd, and it must reach a 45 000 tpd production rate starting from the year 2005. Esmeralda Project is located at the centre of the El Teniente deposit, bounded in the north by El Teniente Sub 6, and under the productive area Teniente 4 South. Production level average elevation is 2192 m above sea level (Figure 1).
PROJECT DESCRIPTION
which allowed the elaboration of the production process and its evolution to a 45 000 tpd rate. The most relevant results of these aspects are the following:
Geology and geotechnics Geology The Esmeralda Project area is fully located in primary rock, except for a small body at the North-East end, where there is secondary rock at Teniente 4 South and Teniente 5 levels. Lithologically, it is mainly composed of Andesite rock, located at the western limit by Brecha Braden Pipe. There are also some porphyries and anhydrite hydrothermal breaches, especially at its northern and southern boundaries (Figure 2).
Structural geology The Esmeralda Project area is crossed by important structural domains, thus defining geotechnical units that have a different mining behaviour. The prevailing fault systems have a northeastern direction (N60°E) and northwestern (N60°W), with subvertical dips. Characteristic of these areas is a high relative frequency of small veins, which cause a notorious heterogeneity and structural anisotropy of the rock mass.
Fragmentation prediction An important aspect of the geotechnical works performed, are the fragmentation studies of which the main result is the granulometric fragmentation prediction, obtaining primary and secondary fragmentation curves for the different geological units. These results were the foundation for the material handling design from the production level to the transport level at Ten 8, as well as for the sizing of the transfer infrastructure and equipment size requirement. These curves were validated after the beginning of the extraction process (Figure 3).
Geotechnical characterisation This was used in geomechanical studies, with application in the support designs and modelling studies to determine mining sequence and undercut area. The most relevant parameters to characterise and classify the geotechnical quality of the rock mass were fracture frequency (FF/m3) and structure condition. For the classification of rock mass and the determination of the fault criteria, the following classifications were applied:
During its basic engineering phase, the Esmeralda Exploitation Project contemplated the development of specific studies in key issues such as:
1.
Q
2.
RMR Laubscher Rock Mass Rating, and
• • • • • •
geological,
3.
m, s
geotechnical,
From this, the following properties and parameters were obtained for the different rock types present in the area to be exploited (Tables 1 and 2):
1.
Mining Engineer, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
geomechanical aspects,
Tunnel Quality Index, Barton et al,
Hoek and Brown Rupture Criterion.
exploitation system mining design,
Geomechanics
cost estimation, and development and production programs.
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Geomechanical features The location of the Project area and the point defined to initiate production had to consider as a relevant feature a 140 m solid rock column to be exploited, which was located under a higher
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M BARRAZA and P CRORKAN
FIG 1 - Production areas, El Teniente mine.
TABLE 1 Rock properties. Rock type Andesites
Breaches
Diorite
268
TABLE 2 Rock parameter.
Index properties
Parameter
Andesites
Diorite
RQD
: 75 - 100 (%)
Pe (ton/m3)
2.70
2.70
2.70
FF
: 0 - 4 (fract/m)
E (GPa)
36 - 38
55
40 - 49
0.18
0.15
0.15 - 0.18
RMRL
: 53 - 68
v
RQD
: 75 - 100 (%)
σci (MPa)
FF
: 1 - 3 (fract/m)
Mb
RMRL
: 58 - 68
S
RQD
: 90 - 100 (%)
FF
: 0 - 2 (fract/m)
RMRL
: 65 - 75
Pe E V σci Mb S
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Breaches
87 - 104
130
100 - 120
8.7
14.3
8.5 - 13.2
0.4
0.4
: Specific weight (ton/m3) : Module deformability, E (Gpa) : Poisson Rate : Hoek and Brown Criterion (Mpa) : Hoek and Brown Criterion : Hoek and Brown Criterion
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ESMERALDA MINE EXPLOITATION PROJECT
FIG 2 - Geologic map, Emeralda Mine.
FIG 3 - Fragmentation curve.
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M BARRAZA and P CRORKAN
area already exploited (Teniente 4 Shadow Zone, Figure 1). The analysis and measurements indicated that the project is located in an area with a medium seismic risk level and stress levels as shown in Table 3.
• creation of a ‘slot’ to ease caving at the beginning of extraction (pre-caving slot);
• modular panel caving method with extraction of ore with LHD equipment;
• proven extraction layout, similar to the one used at Teniente 4 South production level;
TABLE 3 Stress condition.
• gallery arrangement at production level according to Teniente extraction layout (60° Drift-Drawbell);
Stress
Magnitude (Mpa)
Azimuth (°)
Inclination (°)
Major
46
312
-34
• intermediate transfer system from the production level to
Intermediate
35
50
-11
transport level through short orepasses and the elimination of the ‘Y’ type intersection of orepasses, thus avoiding the excessive over excavation of orepasses in primary rock;
Minor
16
156
-53
• main transfer system and infrastructure located in the Formación Braden (Pipe) on highly stable rock; and
Note: positive magnitude is compression, negative inclination is under the horizontal.
• accesses in the production level will be located under
As it is acknowledged the panel caving method is based on causing a force imbalance in the rock mass, so it produces the progressive rupture of the rock through undercutting. The fracturing has relevance on the surroundings, both on the rock quality deterioration and on the concentration of induced stresses. At Division El Teniente, the main consequence of this fracture is the ‘seismic activity’, which as a result releases energy which can produce serious damages to the galleries. This is an inherent condition of the rock and of the exploitation method since an activity tending to fracture the rock is developed.
Exploitation method
EXPLOITATION DESIGN AND OPERATION SEQUENCE After years of experience at El Teniente in the exploitation of primary rock, using the conventional panel caving method, Esmeralda Project has introduced important modifications to the typical exploitation design in order to improve the stability conditions during the exploitation stage (Figure 4). Among the main modifications, we can highlight:
• height of primary ore column to be exploited, less than 150 m;
• • • • • • •
height of undercut level cut (4 m); distance between galleries at undercut level 15 m; undercut level drilling design; beginning of the cave with slot; construction of production level under relaxed zone; monitoring of the caving progress; and
Design criteria The conceptual basis and general design criteria for the Esmeralda Project were the following:
• the area to be exploited must not produce interferences with the exploitation Plan for Teniente Sub 6;
• the base elevation of the undercut level is 12 m below Teniente Sub 5 level and the transport level is located at the same level as Teniente 6 level;
• primary rock caving in virgin areas must not be started - they must be connected to existing caving;
• the ‘caving faces’ lengths must tend to avoid stress
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The selected method was panel caving, with analysis of the advanced and pre-caving variants, which define the sequence of development and construction. The exploitation with advanced and pre-caving variants represents an improvement in the distribution of stresses at the production level regarding the existing situation in a conventional application, minimising changes in the tensional state of galleries (Figure 5).
• The undercut level is located 18 m above the production level, with a 15 m gallery separation excavated parallel to production level drifts. The gallery section is 3.6 × 3.6 m and a flat undercutting will be performed at a height of 3.6 m. The drill holes for the undercutting have a diameter of 2.5″ and are drilled at 60° to the galleries axis. Thus, the maximum drill hole length is 8 m (Figures 6 and 7).
• Swell removal from the galleries, to allow the next explosion and to detect the formation of possible pillars, will be performed with 3.5 yd3 remote controlled LHD equipment.
• The ore extraction system at the production level operates using 7.3 yd3 diesel LHD equipment. Oversized ore will be reduced in size through secondary blasting at the extraction point.
• The LHD dumps to an intermediate transfer system with short 30 m orepasses that unload at the haulage level into rail wagons. The tipping points have 40″ grizzlies with remote controlled semi-stationary rock-breakers to control fragmentation size.
• Intermediate transportation at Teniente 6 will be carried out through an eight-wagon railway train, each wagon with a 50 ton nominal capacity, from the loading chutes located at the different cross-cuts to the dumping stations at the main transfer system, situated in the Braden pipe.
local and global mine seismic monitoring.
concentrations;
undermined area to assure its stability.
• The main transfer system to Teniente 8 level consists of three 5 m diameter shafts (OP22, OP23 and OP24), which together, will allow production to go beyond 45 000 tpd. The transfer system comprises three dumping stations with storing hoppers for 7500 t and loading chutes at Teniente 8 level.
• The transport system at Teniente 8 level included the extension of the existing railroad in order to integrate the new OP22, OP23 and OP24 shafts, and to operate at this level with 18 wagon trains for 80 tons each (Figure 8).
• The estimated initial area to induce the caving, based on the hydraulic radius criteria, with MRMR 50, resulted in a 25 m
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ESMERALDA MINE EXPLOITATION PROJECT
hydraulic radius. The existence of the slot defines a caved face adjacent to the first undermined areas, allowing the decrease of the hydraulic radius to 20 m. The latter, defines an initial extraction area of 10 400 m2 with a 25 000 m2 undermined area of advanced caving. In order to achieve the 45 000 tpd production rate for the year 2005, an average extraction rate of 0.44 t/m2/day will be required over an estimated area of 102 151 m2.
PRODUCTION PLANNING Extractable reserves correspond to the ore to be extracted from the mine to be processed at the Colón concentrator. The determination of the quantity and quality of the reserves depend on the current technical-economical context and on the Divisional goals. The following methodology, parameters and criteria were used to estimate the reserves:
FIG 4 - Development and construction sequence for conventional, advanced and pre-undercut panel caving.
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FIG 5 - Stress redistribution scheme for conventional panel caving and pre-caving.
Definition of extraction method
Economic model
During the execution of the conceptual engineering, and based on a pre-established block model, the most convenient method was estimated regarding the economic performance associated to a cut cost of $US 0.64/lb fine Cu. With the current characteristics and operational conditions of the Esmeralda Project, the area defined is 640 070 m2.
The technical and economical information that permits evaluation of each block of the model is the following (Table 4): TABLE 4 Technical and economical information.
Dilution of the in situ reserves
Concentrator recovery
86.4 %
The estimation of dilution follows the application of the Laubscher volumetric model, through algorithms developed within the División. The beginning of dilution entry was considered at 43 per cent of the extraction height, for the area north of N200 coordinate, due to the lower height of the column. For the remaining area, 47 per cent extraction was used. In situ rock column heights fluctuate between 113 and 224 m, prevailing a 164 m height under the T-4 South exploitation area. Diluting material has a 0.6 per cent Cu grade and 2.0 t/m3 density.
Concentrate grade
28.58 %
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Preparation cost
528.0 $US/m2
Mine operation cost
2.58 $US/m2
Concentrate cost
2.80 $US/m2
Drying/filtering cost
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3.71 $US/t-conc
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ESMERALDA MINE EXPLOITATION PROJECT
FIG 6 - Drilling and blasting scheme, medium pillar ‘JW’ (locally termed ‘John Wayne method’).
The reserves are assessed according to an eight per cent moisture copper concentrate, estimating earnings with a sale price (FOB-port) of 70.35 cents/lb, equivalent to $US 1.00/lb fine commercial copper after international processing fees have been discounted.
MassMin 2000
Definition of the extraction sequence This is defined as a balance between the economical need of exploiting areas with better grades during the first periods and technical restrictions associated with:
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FIG 7 - Drilling and blasting scheme, complete pillar JW.
• • • •
type of advanced caving method;
Reserve selection criterion
geomechanical risk;
In order to maximise the economic benefits of the Project, the opportunity cost criterion was used to select the extractable reserves of the area. This refers to the cost associated to the postponement of current exploitation benefits through the extraction of better and high grades and it is estimated through computing algorithms designed by the División, assessed as a ten per cent discount rate. As a result, the column height is obtained for each column of the block model, and consequently the cut off grade, which maximises the Project’s Net Present Value (VAN).
orientation of the caving face against main structures; and operational factors regarding the interaction with other productive areas.
Thus, exploitation began in the Northeast area of Esmeralda, and then it moved on to the West. The sequence continued with successive moves to the N-S and E-W (Figure 9).
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FIG 8 - Esmeralda Mine ore flow scheme.
• the incorporation of maximum new area of 36 000 m2 per
Smoothening Extraction heights during the reserve selection stage combined with opportunity cost are derived from a mathematical solution. They adjust to the operational restrictions involved in the extraction process. The company tries to model the operative feasibility of the Project’s exploitation, modifying, if necessary, the extraction heights based on the following criteria:
• Minimum extraction height • Maximum extraction height • Adjacent columns height
year; and
• the variable production rate of 250 tpd in 1997 to 4000 tpd in 1998, increasing from that year at an annual rate of 6000 tpd until reaching a rate of 45 000 tpd by the year 2005. The production plan obtained, until reaching the rate, is shown in Table 5.
0.5 times in situ height 1.5 times in situ height
20 m As a result, extractable reserves were obtained, which were included in the production plan, that is, 348.5 million tons of ore with 1.014 per cent Cu.
PRODUCTION PROGRAM
ECONOMIC INDICATORS AND PRODUCTIVITY Productivity and costs It is estimated that 690 employees will be required to operate at the rate of 45 000 tpd, including administration, maintenance and preparation. Productivity will add up to 65.2 t/man/day, with a total mine operation cost of $US 2.85/t.
Economic evaluation
Production plan Corresponds to the extraction simulation of the extractable reserve for the given sequence. The following operational restrictions are imposed:
• the variable extraction rate, 0.1 to 0.5 t/m2/day with respect
The following table shows the economic indicators before taxes, considering the base case technical parameters defined for the project, and based on CODELCO’s regulations (Conservative parameters), with a 12 per cent discount rate and currency corresponding to January 1995 (Table 6).
to the extraction percentage;
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Purchasing
TABLE 5 Production plan. Year
Production (tpd)
Average grade (% CuT)
1997
250
1.277
1998
4.000
1.403
1999
10.000
1.335
2000
16.000
1.378
2001
22.000
1.466
2002
28.000
1.473
2003
34.000
1.413
2004
40.000
1.319
2005
45.000
1.279
Purchasing mainly comprises the purchase of equipment and supplies for the works, such as metal structures, and rail laying and trolley material. In the case of equipment, and in order to maintain the 45 000 tpd rate planned for the Project, the following quantities and types have been set (Table 7 and Table 8): TABLE 7 Equipment. No
Development Equipment
3
3.5 yd3 Scooptrams, remote control
4
Advance jumbos
2
Bolting jumbos
2
Blind raise Drifters, 1.5 m diameter
TABLE 6 Economic indicators.
TABLE 8 Equipment.
Indicators with base technical parameters Conditions
VAN ($US M)
TIR (%)
IVAN
No
Production Equipment
4
SW Radial jumbos
Base Cu Price
: 100
450
40
2.8
2
Explosive loading trucks
Sens Cu Price
: 115
618
46
3.8
27
Diesel 6 yd3 scooptrams
: 85
267
32
1.6
34
Semi-mobile breakers
8
Short fork lift truck Secondary reduction jumbos
Base Cu Price
: 100
Sens + 10 % INV
: +5 Cost
411
37
2.3
9
Sens + 10 % INV
: +5 Cost
490
44
3.3
3
High reach drill rigs
7
50 t production locomotives
1.9
44
50 t production mine cars
2.7
1
Service locomotive
1
Cleaning locomotive
Indicators with conservative technical parameters Base Cu Price Sens. Cu Price
: 100 : 115 : 85
302 448
33 39
144
25
0.9
1
Railroad cleaner
: +5 Cost
263
30
1.5
3
Railway Wagons cleaning equipment
: +5 Cost
342
37
2.3
3
Various support wagons
1
Railway ditch cleaning equipment
Base Cu Price
: 100
Sens + 10 % INV Sens + 10 % INV
INVESTMENT MANAGEMENT AND START-UP Codelco’s directors approved the budget required for the start-up of the Project in January 1995, taking into account the support developed from basic engineering and the favourable economic indicators of the project.
Engineering As first relevant activity for the execution of the Project, and the activity planning engineering to be carried out in detail was developed, which mainly includes the execution of accesses, developing tunnels for the various levels and maintenance infrastructure, equipment and supplies purchasing, and the corresponding construction and assembling works. The activity planning within the updated program context considered engineering, purchasing and construction, which made it possible to carry out the start-up ten months earlier with respect to the initial program, that is 1 July 1998, the date initially set. Production began during the second semester of 1997.
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On the other hand, the main equipment the Project plans to install for production support and service are the following (Table 9): Within the equipment purchasing context, we can highlight the evaluation of offers for fans V-54 (Extractor) and V-63 (Injector), which are associated to electrical equipment such as motors, cabinets and frequency drivers.
Construction Construction includes the development of tunnels and mining civil works and infrastructure that make possible to materialise and place the Project into production. It sets the future execution stage for constructions such as extraction points, roads, dumping points, railroad chutes, trolley, dumping stations, stockpiling hoppers, workshops, offices and others. Likewise, the assembly and start-up of the electrical, electronic and mechanical systems required for the Project are performed.
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Investments
TABLE 9 Equipment. No
Main Equipment
1
Main SSZEE 2 × 7.5
According to the works schedule to be performed and its distribution over time, the investment needed to materialise the Project within the period 1995-1999, in currency corresponding to September 1995, amounts to $KU 189.626. Table 10 shows the annual investment detail indicating the percentages of the main items.
1
OMVA 33.0113.8 KV
1
Secondary SSIEE 3 × 1500 KVA 13.810.6 KV
1
Fans 1400 IAP ea (V:54, 55, 63 and 64)
ACKNOWLEDGEMENTS
1
Production control system
1
Breaker and sprayer remote control system
1
Automatic signaling command FFCC TEN-6
The authors thank the Mine 2000 Superintendency, El Teniente Division, CODELCO Chile, for the authorisation the publish this paper.
1
Automatic signaling command FFCC TEN-6
4000 m
Railroads FFCC 80ib TEN-6 and Trolley 650 VCC
2000 m
Railroads FFCC 1321b TEN-6 and Trolley 650 VCC
3
Compressors 960 cfm FAD 3 × 110 KW
2
Compressors 300 cfm FAD 2 × 30 KW
1
Compressor 100 cfm FAD 1 × 22
5
Rectifiers 750 KW ea CC FFCC system
2
Rectifiers 1.000 KW ea CC FFCC system
1
Workshop LHD, Jumbos, Breakers Blind Hole (4500 m2)
1
Shed for wagons and locomotives (1000 m2 11 m high)
1
Petrol station for 64.000 ton capacity equipment
REFERENCES Barraza, M, Bustamante, S, Crorkan, P and San Martín, J, 2000. Esmeralda Mine planning task group. Hustrulid, W, 2000. Recommendations regarding Pre-undercutting in Esmeralda Sector. Lasagna, G and Vargas, J, 1999. Year 2005 production capacity. López, S and Molina, R, 2000. Post evaluation pre-caving variant at Esmeralda Mine. Vargas, J and Montecino M, 1997. Esmeralda exploitation project, División El Teniente, a challenge for the 21st century.
TABLE 10 Annual investment. Annual Investment Cash Flow (k$US) Description
1995
1996
1997
1998
1999
Total
%
Engineering*
2.471
4.264
2.805
1.6161
1.127
12.283
6.5
Purchasing
4162
8.017
4.310
4.009
3.811
30.563
16.1
Construction
2.064
34.751
32.309
30.620
27.036
146.780
77.4
Total
24.951
47.032
39.424
46.245
31.974
189.626
100
* it includes management, internal engineering, engineering and inspection
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‘An Underground Air Blast’ — Codelco Chile - Division Salvador R de Nicola Escobar1 and M Fishwick Tapia2 INTRODUCTION This paper was prepared and translated from the final report of the investigation into the Codelco Chile Salvador Division’s air blast which occurred in the Inca-West area. In this section of the block cave mine, a stable arch formed covering an area of approximately 15 000 m2. This zone subsequently collapsed producing a massive air blast. The collapse followed a series of planned extraction activities and occurred on 5 December 1999. This paper discusses the cave back arching event and the measures that were taken in order to induce the controlled collapse. This was to minimise risk to personnel, equipment and installations. The results obtained to date are presented.
BACKGROUND INFORMATION The region where the stable arch formed was part of the Inca -West mining area of the Codelco Salvador Division underground operation. The mining method used is block caving and the production level design layout is based on the Teniente-LHD layout. The ‘stable arch’ formed in the South West section of the current mining area and covered mainly the primary andesite rock type with a mineralised column height that varied between 180 and 210 metres. The undercut level was at an average depth of 700 m below surface. Above the column height is the crater which formed as a result of previous mining. Extractable mineral reserves of the area correspond to 66 per cent of the total long-term mineral reserves of the mine. The target production for the years 1999 and 2000 are 2.3 and 2.5 million tonnes respectively, with an average grade of 0.5 per cent Cu in both periods. Figure 1 shows the location of the Inca - West area of the underground mine and the area of current mining.
EXTRACTION METHOD In the Inca West area, where the stable arch formed, the extraction method was panel caving, with an extraction grid of 15 × 15 m (225 m2 per draw point). The reduced levels of the production and caving levels were 2454 and 2470 metres above sea level respectively. Drawpoint loading is with six and seven cubic yard loaders and the ore is dumped into ore passes with grate (grizzly) openings of 34″ × 62″. The rock breaker (reduction) level is located at the 2438 RL. The ore is then transferred to storage silos that unload to trains on the main haulage level. These trains transport the ore to the primary crusher located on the surface. Table 1 shows a summary of some of the basic data for the Inca West panel cave. The predominant lithology corresponds to a biotised primary andesite, which is intruded by an approximately 20 m wide veinlike form of silicified primary andesite, crossing from east to west all of the ore column in the south section of the undercut. This more competent intrusive effectively separates the cave into two distinct areas where cavability may differ (see Figure 2).
1.
Chief Mine Planning, Subgerencia Mina-Concentradora, Codelco Chile, División Salvador.
2.
Deputy Manager, Mina – Concentradora División Salvador.
MassMin 2000
TABLE 1 Summary of some basic data for the Inca West Panel cave. Kton Kton Kton Kton
113 836 51 713 12 531 49 592
% %
0.647 0.024
Life of Mine Extraction rate for period 2000 - 2006 Extraction rate for period 2007 - 2014 % Primary ore 2000 - 2006 % Primary ore 2007 - 2014
Years Kton/year Kton/year % %
15 5056 9806 76.83 47.23
Extraction grid Distance between drawpoints Distance between cross cuts Production level RL Caving level RL Rock breaker (reduction) level RL Ventilation level RL Height of undercut Drill hole diameter Length of ore passes Opening of ore pass grate Opening of rock breaker grate Crusher opening (Setting)
m² m ml m.a.s.l m.a.s.l m.a.s.l m.a.s.l m inch m inch × inch inch × m inch
15 × 15 = 225 15 30 2454 2470 2438 2446 3.0 2 ½″ 12 34 × 62 13 × 3.0 6
Reserves Primary ore with Anhidrita Primary ore without Anhidrita Secondary ore Average grade Cut Average grade Mot
Dimensions of workings Working Xc drive Production drawbells UCD Xc Ventilation Xc Reduction (rock breaker) Ventilation shafts Free face shafts (Long = 11.2 m) Ore pass Total lineal metres Caving (Year) 1995 1996 1997 1998 1999 Total caved by 1999
Section 4.3 × 3.8 m 4.0 × 3.5 m 3.0 × 3.0 m 3.5 × 3.5 m 3.5 × 3.5 m diam = 2.0 m diam = 1.0 m diam = 2.5 m 14 237 ml Caving (m²) 11 438 1660 3614 3568 6363 26 643
Mining activities began in 1994, the plan included a caving sequence from the south towards the north-west for the following year and an undercut footprint of 11 439 m2 to begin production. Following this, and in agreement with the established plans, the undercut footprint was progressively increased reaching 20 280 m2 in July 1999. The total ore extracted was approximately 3.2 Mt. At the end of 1998 a Jaw crusher was commissioned in the intermediate ‘reduction’ level, replacing the rock breakers. This enabled an increase in the extraction rate from 0.12 to 0.28 t/m2/day. As a consequence of this, a number of draw points
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began to be depleted of ore particularly in the extreme south–east section of the panel. These subsequently become empty and the cave back could be seen from the drawpoint positions. This situation increased progressively to the rest of the production area, resulting in a total of 32 opened draw points and covering an area of approximately 15 000 m2 by July of 1999. All activities in the area were stopped to analyze the phenomenon given the risk of a sudden collapse, which could cause an air blast or piston effect with undetermined consequences to people, equipment and installations. In addition, the ‘stable arch’ would not allow the successful completion or achievement of the established production targets. Similar experiences around the world have had catastrophic consequences, due to the high speed or velocity of displacement air through mine workings. Below is a summary of some of the geotechnical information associated the Inca West Area:
Stress components in MPa σN-S
σE-W
σS-Vert
σE-N
σE-Vert
σN-Vert
17.5
14.12
18.51
0.02
1.67
3.84
Principal stresses in MPa σ1
Αz
I1
σ2
Az
I2
σ3
Az
I3
22.08
14.2
49.1
15.10
134.1
23.3
12.95
239.3
31.4
Condition of in situ stresses around caved area
MRMR (Laubscher)
Location
: Drive #22, production level
Coordinates West
: -7792.001
North
: 19557.274
RL
: 2457.015 m
Date
See caving chart below.
: 25 August 1999
Depth of measurement : 11.53 m
Stress Environment (in situ) (based on 1995 measurements)
Bearing
: 298.57°
Dip
: -1.79
Location
: drive #18, production level
Rock type
: Primary Andesite
Coordinates West
: -7855.000
Poisson ratio
: 0.29
North
: 19490.306
Young’s modulus
: 62.97 GPa (desde test biaxial)
RL
: 2454 m
Date
Stress components in MPa
: 30 November 1995
Depth of measurement : 8.34 m
σN-S
σE-W
σS-Vert
σE-N
σE-Vert
σN-Vert
Bearing
: 225°
23.05
33.06
39.13
-6.10
2.32
0.05
Dip
: -4°
Rock type
: Primary Andesite
Poisson ratio
: 0.33
Young’s modulus
: 81.73 GPa (from biaxial test)
Principal stresses in MPa σ1
Az
I1
σ2
Az
I2
σ3
Az
i3
40.20
289.2
-63.9
34.94
117.3
-25.9
20.10
25.7
-3.2
CAVING CHART AFTER LAUBSCHER LAUBSCHER 1990
100
STABLE
90
TRANSITION
80
LAUBSCHER 1995
70
60 MR MR
UNSTABLE
50
40 EXTENSION PROGRAM (RH) = 44
30
UNDERCUT DECEMBER1999 (RH) = 38
20 TYPICAL TENIENTE CURVE
UNDERCUT JULY 1999 (RH) = 34
10
0 0
10
20
30
40
50
60
70
(RH) = RADIO HIDRAULICO = AREA/PERIMETRO (mt.)
280
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‘AN UNDERGROUND AIR BLAST’ — CODELCO CHILE - DIVISION SALVADOR
UBICACIÓN INCA OESTE EN MINA SUBTERRÁNEA Inca Oeste.
200 2
200 6
200 1 200 5
200 0 200 5
200 4
200 4 201 2 201 1
201 1 201 0
200 9 200 7 200 6
200 8
200 3
200 2
200 1
200 5
200 0 200 3
200 4 201 1
201 0
200 6 200 9
200 2
200 1
200 8 199 9
200 5 200 7
200 4
200 3
Zona desarrollada Inca Oeste.
FIG 1 - Location of the Inca West area in the underground mine.
FIG 2 - The andesite intrusive feature which divides the mining area into two distinct zones.
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THE CAUSE
SOLUTION
The cause of the arching can be partly explained by the geotechnical characteristics of the rock mass in the opened area and the conditions of ‘clamping’, which generated a stable zone that did not allow the propagation of caving at a rate sufficiently higher than the planned extraction rate, thus creating a zone of stable arching.
Given the nature of the caving problem and the analysis of all the available information, there were two possible solutions:
THE PROBLEM The problems associated with the formations of the stable arch, was the development of a cavity and the subsequent possibility of a sudden collapse of the corresponding cave back which would result in the displacement of a large volume of air travelling at high speeds. This had the potential to injure people and damage, equipment and installations.
ACTIONS UNDERTAKEN A working group, comprised of personnel from CODELCO and external consultants was formed. The objective of this working group was to rehabilitate the area thereby minimising the risk to people, equipment and installations. As an immediate measure, all activities and personnel access to the area were stopped whilst follow up actions were considered. The actions undertaken to achieve the main objective are summarised as follows:
data collection, numerical modelling, lithological and fragmentation analysis, and productive capacity of the area.
Dimensioning of cavity • • • •
With respect to the vertical rate of propagation of the cave, it was proposed to maintain a strict control through geotechnical and seismic instrumentation, so that the extraction rate at the draw points was less or at most equal to this propagation rate. (estimated to be below 0.07 ton/m2/day). The above satisfies one of the issues, to initiate the collapse of the stable arch.
RISK MINIMISATION
1.
maintain draw points and drawpoint drifts full or ore;
2.
develop an emergency procedure;
3.
maintain a permanent seismic monitoring system; and
4.
have constant personnel communication.
THE RESULTS
cavity measurement with laser equipment, and seismic reflection.
closure of draw points, contingency strategy for extraction, stress monitoring, emergency procedures Inca West, and communication to personnel.
• increase of undercut footprint, and • ore extraction. Actions after the collapse
282
To increase the open area towards the North and West, ie zone of primary biotised andesite, with the view to move away from the apparent equilibrium condition of the rock mass, postponing in time the caving of the clamped and most competent zone, whilst another more technically attractive and economic method to induce caving was evaluated (see Figure 3).
exploration drilling,
Operational
• • • • • • • •
2.
installation of micro-seismic system,
Mitigation of operational risks • • • • •
To induce continuos caving through drilling and blasting. This alternative was considered slow (estimated to take more than ten months), appropriate equipment was not available, high in cost and had an uncertain outcome.
The second issue addressed was that any sudden collapse should not cause injuries to personnel and damage to equipment and installations. For that, preventative measures were adopted to minimise the consequences of the air blast. These were:
Analysis and investigations • • • •
1.
frequent aerial observations of surface crater, aerophotogrametric survey of surface crater, lithological and fragmentation analysis, numerical modelling, pulling strategy, increase of opened area, geomechanical monitoring and control, and
The additional caving of 6039 m2 and the controlled ore extraction of 230 000 tonnes, carried out during the months of August until the 5 December, meant an increase of the fracturing process of the rock mass, mainly in the zone of primary biotised andesite. This resulted in two collapses. The first occurred in the North-West zone on 5 December, when an area of approximately 4000 m2 collapsed which daylighted in the crater and followed by an air blast with significant displacement of air and dust throughout the underground mine. This forced the total evacuation and stoppage of operations for a period of 24 hours. On 8 December a second collapse occurred, also in the area of biotised andesite. Although the magnitude of this event was similar to the previous one, the effects of the air blast were less due to the lesser volume of air displaced, given that the stope was semi-full. The physical evidence of cave propagation through the full column height was evidenced by subsidence of the crater, significant variations of the stress condition around the undercut footprint, cutting of TDR cables and changes in the location or concentration of seismic activity. The most unfavourable estimation was that the collapsed area would be around 4000 m2. Owing to the height to the cave back in the collapsed zone (>140 m), the lateral rilling of material increased the quantity of ore above the extraction points located in the uncaved region, thus increasing the cushion of ore in those draw points as shown in Figure 4.
emergency procedures modification.
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‘AN UNDERGROUND AIR BLAST’ — CODELCO CHILE - DIVISION SALVADOR
S-
50
°0
20
0'
.3
6. -W
57
m
40
@
m
60
Ho °0
ri
z.
12
FIGURA Nº3: SOLUCIÓN, AMPLIACIÓN DE ÁREA ABIERTA .4
0m
0' 20
Ho
.3
Sri
69
m
z. S@
60
°0
20 50
60
°0
°0
0'
0'
.3
-W
34
m
@
60
°0
0'
20
.3
S-
46
m
@
50
60
°0
°0
0'
TROYA-3A
-E
4
5
SECTOR INCA CENTRAL OESTE ESCALA
MAYO 30, 1999.
7
WSO/MVM.
8
7850.00W
TROYA-29
SECTOR HUNDIDO EN PRODUCCIÓN
6
1 : 2000
SECTOR OESTE ELEV. 2486.233m
3 4
C
E .V
NT
.
N-
EL
54
.
2
00
'-
6 44
E
.0
0m
20 21 22
9 10
P.7 11
7893.75W
TROYA-31
12 13 14
5 11
13
10
9
TROYA-7
TROYA-5
TROYA-3B 2 3
TROYA-3
0'
-E
0'
18
17
El. 2434.00
12
15
14
6
EL. 2434.00m7
19620.00
8
19
16
76.4
20 9
1m
Hori
z. N-70
10
°00'
11
-W
12 86.5
16
7760.00
7850.00 '-
E
PI
45
18
17
20
31
E-
28
30
24
N-
52
°
' 00
26
19 20
RAMPA 19528.00 @ +14%
27 23
22
21
22 18
21
AV. 19521.50N
29
-E
C. Vent. EL. 2446.00m
TROYA-15
@ -1 4%
17
19500.00 EL. 2446.00m
24 19486.436
24 25 23
9
22
8
21 20
AI RE
7
AREA SOCAVADA ACTUAL
S30 00 'W
11 10
19
QU
20
2446.00m
16
21
19
16
TROYA-13
25
E-
00
QU
18
5m
15
FR ES CO
iz. 8m. Hor
75.073m.@-10% = 542'38"EL.
15
14
7918.480 SUB-ESTACION ELECTRICA
RAMPA 19500.00
12
7928.687
141.01
TROYA-17
21.266m. Directo N
11
'20"-W
PI
N-
8030.00W
8120.00W
AREA A HUNDIR PARA TROYA-19 13 PROVOCAR HUNDIMIENTO
TROYA-21
S-7950
7940.00W
17
7715.00
15
7696.25
El.2446.00
14
7651.25
AREA A HUNDIR FUTURA
13
19
6 10 12
11 19360N
13
15
14
11
9
16
12
EL. 2438.53m
18
13
17
18
14 19
AIRE FRESCO
20
15
21 16
17
RAMPA 19368.00 75.073m. @-10%=5°42'38" EL. 2455.20m
17
18
19
20
21
22
7800W
AREA VIRGEN
19368.00 EL. 2446.00m
19360N
8000W
16
EL. 2446.00m
22
ZONA EMPOTRADA A CERRO VIRGEN
FIG 3 - The solution to increase mining area to induce caving.
THE AIR BLAST When the first collapse occurred, the collapsed ore displaced the air in the cavity which flowed in between the broken material, sealing the draw points and the extraction drives towards the production level, without generating appreciable damage. The height of broken ore above the drawpoints when the first collapse was approximately 13.5 metres. Afterwards the displaced air converged at great speed towards the ventilation level workings, where it caused the most significant damage to infrastructure and injuries to personnel who were evacuating the area at the time. The air speed in the main access was estimated to be greater than 500 km/hr. The main consequences were:
• an injured worker with a fractured left wrist due to a loss of footing;
• 17 workers with minor injuries (acoustic trauma and foreign objects in eyes);
• a 4WD vehicle located near the main access was overturned; and
• various damage to ventilation gates. The second collapse was seismically similar to the first one, however this event had no negative consequences of any kind, due to the cushion of ore already present on top the pillars from the first collapse, which better contained the piston effect.
MassMin 2000
The preventative measures adopted minimized consequences of the air blast and these included:
the
• maintaining draw point drifts and draw points full and covered with ore material,
• maintaining an emergency procedure, • maintaining permanent seismic monitoring, and • constant communication with personnel. In retrospect however, the emergency procedure which required personnel to be evacuated through main access tunnels needs re-evaluation. In the main access tunnels the air acquired greater speed and there were injuries to personnel and minor damage to equipment and installations. However, the negative consequences would have undoubtedly been greater if actions to mitigate against a potential airblast had not been taken. The geotechnical modelling was demonstrated to have a good predictive capability. This allowed safety measures to be put into place before the imminent air blast. After the collapses, numerical modelling was also to simulate possible future caving conditions, with the view to expand the area in connection with the crater. The idea was to make all the caving fronts uniform. It was always anticipated that caving of the full column height would occur in the weakest zone. This was based on the frequency of seismic activity in the area (see Figure 5), changes in fragmentation at draw points and continuity of production at these points.
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R DE NICOLA ESCOBAR and M FISHWICK TAPIA
PERFIL ESTIMADO DE PROPAGACIÓN DE CAVING
Cráter
Techo cavidad
Talud de escurrimiento
Zona empotrada
FIG 4 - Section showing the approximated cave propagation, clamped region, rill angle and back of cavity.
Seismic activity in the extreme South-East was almost non existent, suggesting a slow vertical propagation of caving. After the collapse, ie between the months of December and February, there was a significant increase in the seismic activity within the stable arch region, which alluded to the fracturing of the rock mass, fundamentally supported by the increase in opened area and productivity pull (see Figure 6). The seismic monitoring also showed that a reduction in the extraction rate also reduced the level of seismic activity, or in other words, when there is no extraction, no fracturing is produced in the rock mass. Figure 7 describes the sequence of events.
CURRENT CONDITIONS AND IMPACT ON MINE PLAN The section of the mine affected by the collapses is now producing and the production rate is approximately 1800 t per shift. Geotechnical monitoring is being continued and it is showing no signs that would indicate potential problems with the continuity of caving and therefore risk to personnel, equipment and installations. As a cautionary measure however there will be no extraction from the draw points located in the extreme South-East. This situation will continue to be studied.
284
With the advance of caving towards to the North, a connection will be made at the end of the third quarter of this year (2000) with the caving of the Inca central region. This will then ensure caving of the whole West section, with the exception of the South-East zone. This is in spite of the fact that numerical modelling predicts that this South-East zone will achieve continuos caving. There is also no evidence suggesting that the mine-wide propagation of caving will be impeded. In addition, the increase in caving area means a greater number of available draw points. The Inca West region may return to the normal levels production levels planned for 2000, ie 2000 t per shift, 145 available draw points and extraction rate of less than 0.25 t/m2/day). As a result of the ‘air blast’, the mining plans for the years 1999 and 2000 were revised to reduced productions of 1 354 000 and 1 306 000 t respectively with an average grade of 0.5 per cent cu/T. The increased pull of remaining ore with better grade (0.8 per cent Cu/T) from the Inca Central section and the recovery of ore columns placed in crossed collapses in the same region, have compensated for the reduced production of the Inca West region.
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‘AN UNDERGROUND AIR BLAST’ — CODELCO CHILE - DIVISION SALVADOR
FIGURA Nº5: REGISTROS SÍSMICOS (21/07 AL 05/12) Límite de hundimiento al momento de colapsos 26.319 m2, Rh = 38,6 m Zona de concentración actividad sísmica.
Área estimada involucrada en colapsos > 4.000 m2.
Zona sin extracción. No registra actividad sísmica.
FIG 5 - Registered seismic events (21 July 1999 to 5 December 1999).
FIGURA Nº6: REGISTROS SÍSMICOS (09/12/99 AL 29/02/00) Límite de hundimiento actual. 28.557 m2, Rh = 42 m
Área cavimg inicial
Zona de propagación actividad sísmica.
FIG 6 - Registered seismic events (9 December 1999 to 29 February 2000).
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R DE NICOLA ESCOBAR and M FISHWICK TAPIA
Roca in situ Roca in situ aire
aire
Aumento de área abierta
Situación inicial
Debilitamiento de losa de Roca in situ
Conexión a cerro hundido de niveles superiores.
aire
aire
Propagación del caving en altura
Colapso. Generación de caving
FIG 7 - Sequence of collapse events/air blast.
Figure 8 describes the impact to production due to this contingency plan from the base case 2000 and shows that the area will converge in production with what has been planned for the year 2001. The advancement in open area with respect to the base case 2000 to solve the problem, has only altered the schedule of the year 2000, with caving being inferior in the following five years.
CONCLUSIONS AND COMMENTS Caving of the full column height mainly occurred in the North-West region where there was a major presence of primary biotised andesite, with ore movement towards the central and South-East section filling the cavity following a natural riling angle. It was recommended to continue with the extraction to expand the area of continuous caving. Together with this, a controlled extraction was maintained. Undeniably, the apparent volume of air that was present in the ‘opened stope’ was much less than that present when the stable arch was detected. Estimates indicated that of the initial 500 000 m3, today there is only 50 000 m3 of air. The collapses had no catastrophic consequences due to the fact that the control measures, although somewhat incomplete, were properly implemented. The losses that occurred were minimal.
286
The experience obtained from the air blast has indicated that under no circumstances, given the possibility of a collapse, personnel evacuation should be carried out through the main access routes. It is recommended to install well implemented refuge chambers in blind workings (ie with telephone, compressed air, first aid kits, etc). If personnel are not capable of reaching these refuges then they must go to any blind working and locate themselves closest to the ground. Numerical modelling showed a good predictive capability. From this, it was possible to carry out an almost precise estimation of what extreme natural conditions of instability caving would occur. For that it was necessary to have a thorough previous knowledge of all of the geotechnical parameters and geological characteristics of the rock mass, and with these as the basis of input information, configure a correct numerical model of such characteristics. It was recommended to continue with the instrumentation which included, seismic monitoring, TDR and stresses, since theses measurements allow to further understand, interpret and control the evolution of caving in the area. This must be carried until there is certainty about the conditions of the zone that has not caved. After the collapse, a drawing rate strategy was established with the aim to support the evolution of caving towards the non-caving zone and thus regulate the drawing rate from the caved zone. In addition, this strategy defined the drawing rate criteria under which certain draw points can be incorporated in
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MassMin 2000
‘AN UNDERGROUND AIR BLAST’ — CODELCO CHILE - DIVISION SALVADOR
P r o d u c c ió n S e c t o r I n c a O e s t e S u r 3 .0 0 0
2 .5 0 0
Ktons
2 .0 0 0
1 .5 0 0
1 .0 0 0
500
0 19 9 9
2000
2001
2002
2003
2004
2005
A ños
Program a
Program a Real D if e r e n c ia
A ño KT ons KT ons KT ons
1999 2 .3 0 0 946 - 1 .3 5 4
2000 2 .4 6 0 1 .1 5 4 - 1 .3 0 6
Real
2001 1 .9 8 0 1 .9 8 0 0
2002 2 .3 8 8 2 .3 8 8 0
2003 1 .9 5 2 1 .9 5 2 0
2004 1 .9 9 9 1 .9 9 9 0
2005 2 .0 0 7 2 .0 0 7 0
FIG 8 - Impact on production.
sections presumed to be under no continuos caving and that are expected to be in a caving regime and also precisely where there has been a positive evolution of rock mass failure according to the interpretations from the seismic information. It is fundamental for any new block cave to estimate the speed of propagation of caving so that an extraction rate which does not exceed this speed of propagation can be defined, with the aim to avoid the generation of an air cavity that represents a potential risk to personnel, equipment and installations. In addition, an on-line production control system must be adopted which guarantees a rigorous control of the drawing rate from draw points. After the air blast, the area of Inca West has progressively been incrementing its production, earmarked within the plan of normalization of operations, which basically consists of reaching the extraction levels expected in accordance with the mining plans (approximately 2500 t per shift). It is expected that in January of 2001 the production target of the base case 2000 will be reached. Caving is not only limited to the breakage of the base of the block (undercut), but also to the characteristics of the lithological units present. Undertaking a back analysis of what occurred the following must be considered when starting a new mining block:
• For every new area that is incorporated to production using the caving method, and with new variables as it is the case of
MassMin 2000
primary ore, there must be a thorough knowledge of the geological and geotechnical characteristics, since these are the basis to a correct engineering design.
• Similarly, it should always be taken into consideration that a clamped zone in an area of virgin primary ore will require a larger open area (footprint) to achieve continuos caving, unless palliative measures of inducement or slotting are undertaken.
• To comply with production plans there must be a greater number of draw points available, considering a low rate of propagation of caving which will depend upon the local geological and geotechnical conditions, without even considering the operational variables such as secondary blasting or layout of the mining area to reduce the interference between the distinct operational units. Instrumentation, monitoring and geotechnical analysis is fundamental to maintain control over the evolution of caving at height, supporting and allowing the correct decision-making to avoid operational risks that can alter the continuity of the production process. When a region is already caved and a new area is to be opened, efforts should be aimed at commencing the undercut from the zone already caved. This favours the rapid propagation of caving. If technically and economically this is not justified, what has been indicated in the points mentioned above should be considered.
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ACKNOWLEDGEMENTS The authors of the report wish to thank the Management of the Codelco Chile Salvador Division for allowing the release and publication of this document. This paper is a translation of a final investigation report of the collapse and subsequent airblast.
REFERENCES Analisis Situacion Inca Oeste, Modelo Flac 3D, Ingenieria De Rocas Limitada (in house report).
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Estudio De Hundibilidad Inca Oeste, Codelco Chile, Division Salvador, Subgerencia Minco (in house report). Estudio Geofisico, Hundimiento Inca Oeste, Reflexion Sismica Geoexploraciones SA (in house report). Informes Internos De: Geologia, Geomecanica Y De Control De Riesgos. Codelco Chile, Division Salvador, Subgerencia Minco (in house report). Sismicidad Inducida En El Sector Inca Oeste, Division Salvador. Dunlop, R and Rojas, E, Codelco Chile, Division El Teniente (in house report).
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Freeport Indonesia’s Deep Ore Zone Mine J Barber1, L Thomas2 and T Casten3 ABSTRACT Freeport Indonesia’s DOZ (Deep Ore Zone) mine is scheduled to commence production in the second half of 2000. The DOZ will produce 25 000 tpd using mechanised block cave mining methods and will have a number of major improvements when compared to previous block cave mines in the district. The DOZ will have an independent and improved ventilation system, a flexible and efficient truck haulage system, and a more reliable crushing and conveying system than in previous underground mines in the district. This paper will describe the mine plan, haulage, and ventilation systems, and compare them to those used in other operations in the district.
INTRODUCTION PT Freeport Indonesia operates a copper and gold mining complex in the Ertsberg Mining District in the province of West Papua, Indonesia (see Figure 1). The Ertsberg District is located in the Sudirman Mountains at elevations from 3000 to 4500 metres above sea level. The topography is extremely rugged and rainfall in the mine area averages 3000 mm (118 in) per year. Freeport began production in the district in 1972 when the mill began processing ore from the Ertsberg open pit. Underground mining began in 1980 when the GBT (Gunung Bijih Timur – Ertsberg East) started production using block caving methods.
In 1993 the first of several studies was completed that indicated that portions of the DOZ could be successfully and economically mined using block caving methods. Limited development was started in 1995. Full-scale development and construction commenced in late-1997 and are on-going. The crusher and conveyor system was commissioned in the first quarter of 2000. Portions of the ventilation system are completed, with two new fans in service and a third recently commissioned. Undercutting and production are scheduled to commence in the second half of 2000. The DOZ will reach its full production rate of 25 000 tpd in 2003. A study is currently underway to determine if an ultimate production rate of 35 000 tpd is feasible.
GEOLOGY AND ORE RESERVES Geology The DOZ orebody is situated in the lower portion of the Ertsberg East Skarn System (EESS). The EESS is hosted by Tertiary age carbonates that have been altered to calcium-magnesium silicate skarn. The EESS is an essentially vertical tabular body with a vertical extent in excess of 1200 metres, a strike length of over 1000 metres and an average width of 200 metres. The northeast (hangingwall) contact of the EESS is a skarn reaction front in sudden contact with barren marble. This contact coincides with a zone of localised faulting and brecciation. The EESS is bounded to the southwest (footwall) by the Ertsberg Diorite intrusive. Moving across the strike of the orebody from the footwall to the hangingwall, the specific rock units encountered in the DOZ are:
• Ertsberg Diorite. Generally a hard, competent rock unit with good ground conditions. Proximal to the skarn contact, the diorite has been locally altered and mineralised. Where appropriate, these mineralised areas have been included in the DOZ reserves.
• Forsterite skarn. A massive unit adjacent to the Ertsberg Diorite contact, averaging 0.5 per cent Cu. Generally a hard competent rock unit with good ground conditions.
• Magnetite-Forsterite skarn. Grades vary between 0.5 - 2.0 FIG 1 - Ertsberg mining district location map.
The GBT reached a maximum production rate of 28 000 tpd in 1991, and was exhausted in 1994. The IOZ (Intermediate Ore Zone) was brought into production in 1994, also using block caving methods, with a design rate of 10 000 tpd. The IOZ is currently producing at a rate of 19 000 tpd. The DOZ (Deep Ore Zone) was discovered in the mid-1980s by deep drilling from the GBT. The GBT, IOZ, and DOZ deposits are all situated in the EESS (Ertsberg East Skarn System). 1.
General Superindendent, Engineering and Planning, Underground Mines, PT Freeport Indonesia, PO Box 434, Tembagapura Irian Jaya 98100, Indonesia.
2
Senior Manager, Underground Mines, PT Freeport Indonesia, PO Box 434, Tembagapura Irian Jaya 98100, Indonesia.
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Chief Engineer, DOZ Mine, PT Freeport Indonesia, PO Box 434, Tembagapura Irian Jaya 98100, Indonesia.
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per cent Cu. Often finely bedded. Generally a hard competent rock unit with good ground conditions but with localised zones exhibiting poor ground conditions
• Massive Magnetite. Occurs mainly along the marble contact. Often strongly bedded. High-grade ore with grades ranging from two to ten per cent Cu. Generally a hard competent rock unit with good ground conditions but with localised zones exhibiting poor ground conditions.
• DOZ Breccia. A lenticular zone which plunges westerly across the lower half of the DOZ, cross-cutting all other units. Ore grades tend to be >2 per cent Cu and locally are greater than four per cent Cu. Almost without exception, ground conditions in this unit are very poor with a history of failure.
• Dolomite-marble. Alteration extends 250 - 300 metres from the skarn into the hangingwall and is generally barren of mineralisation. Rock quality and ground conditions are highly variable and locally very poor proximal to the skarn/marble contact. Figure 2 shows the geology at the DOZ extraction level.
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FIG 2 - Geological plan of DOZ extraction level.
FIG 3 - Geological section EESS.
Ore reserves Current reserves (diluted) for the DOZ block cave orebody are 131 million tonnes grading 1.06 per cent Cu, 0.81 g/t Au, and 7.49 g/t Ag, or 1.66 per cent Cu equivalent, utilising a cut-off grade of 0.90 per cent Cu equivalent.
Hydrogeology The underground mines at PTFI have a history of wet muck runs in the production areas. Wet muck runs as large as 2000 cubic metres have occurred from individual drawpoints. These runs are a serious safety hazard and impediment to production. In order to minimise the problems of dealing with wet muck in the DOZ Mine, an extensive dewatering program is underway. This will reduce the amount of wet muck which reports to the drawpoints in the DOZ. Water enters the cave through rain falling onto the subsidence zone, surface drainage into the subsidence zone, and ground water flowing into the subsidence zone. The water moves down through the cave and creates saturated conditions in the drawbells, which leads to the wet muck runs. Three main sources of ground water that must be intercepted have been identified. 1.
Hanging wall water that is impounded behind a sandstone aquaclude (Sirga Sandstone). As the cave cracks intercept the Sirga SS, the impounded water flows into the cave. The impoundment is continually recharged by rainwater from the surface (see Figure 3).
2.
Long-strike water that flows into the cave along contacts and relict bedding in the skarn formations and along the skarn/diorite contacts. Some of the long strike water is perched water and some is recharge from the surface.
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3.
Recently, major ground water flows in a permeable zone in the Ertsberg Diorite have been identified. This may be a major source of the long-strike water. Further investigation and drilling is ongoing.
This situation is complicated by a series of high-angle cross-strike faults that provide some communication between the above water sources and the cave. The model is further complicated by a number of karst features in the hanging wall. Dewatering strategy is based on draining impounded water from behind the Sirga SS so that the phreatic surface falls below the elevation at which the cave cracks intercept the Sirga SS. This is being accomplished by means of a drill gallery developed in the hanging wall outside the predicted ultimate crack line of the DOZ cave. Diamond drilling from the gallery pierce the Sirga, allowing the impounded water to drain off through the drill holes. Continuous flows from the Sirga drainage drilling are predicted to be on the order of 315 l/s (5000 USgpm). Long-strike water will be intercepted outside the predicted ultimate DOZ cave crack by drilling from development in the hanging and footwall parallel to strike. The long-strike development will be on a number of different levels, utilising existing workings as jumping off points. This method has been successfully used in other areas of the EESS. Continuous flows from the long-strike drilling are also predicted to be on the order of 315 l/s (5000 US gpm).
Geotechnical Geotechnical engineering and planning are based on diamond drill holes drilled between 1988 and 1998. Holes drilled prior to 1994 were logged for RQD only, the newer holes were logged for
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fracture frequency and RMR analysis. As development work has exposed rock in the orebody, scanline and cell mapping have improved our knowledge of the ground conditions. Starting at the hangingwall and proceeding to the footwall, ground conditions go from very poor to very good. Rock strengths (UCS) vary from a high of 219 MPa (31 000 psi) in some massive magnetite to less than 10 MPa (1400 psi) in the DOZ breccia. RMR varies from a low of 25 in the poorest areas to a high of 65 in the most competent ground. Cavability was evaluated using a combination of RQD, RMR, and MRMR data. The hydraulic radius needed to sustain the cave ranges from ten metres in the DOZ Breccia and other areas of poor ground to 30 metres in the forsterite skarn. The total DOZ footprint has a hydraulic radius in excess of 60 metres, so cave propagation will not be a problem. Fragmentation forecasts predicted that less than 30 per cent of the orebody will exceed 1.0 metre. Comminution within the draw column will further reduce the material size reporting to the drawpoints.
FIG 5 - Advanced undercut.
Extraction level design MINING METHOD The mining method chosen for the DOZ Mine is mechanised block caving. The basic method is similar to the methods successfully used at the GBT and IOZ mines.
General description The production (extraction) level of the DOZ will ultimately be 900 metres long and average more than 200 metres wide, with the widest location being 350 metres wide. The maximum height of draw will be 350 metres. Production panels will be transverse to the orebody on 30 metre centres. The undercut level will be 20 metres above the extraction level. Drawpoints will be on 18 metre centres along the panel drifts, yielding a draw column footprint of 15 × 18 metres. Undercutting will begin in the centre of the orebody and progress eastward. This takes advantage of the extremely weak ground conditions and the higher than average grades in the DOZ Breccia. After the IOZ has been depleted, undercutting in the DOZ will be advanced westward, underneath the IOZ. Undercutting will begin utilising the conventional post undercutting methods employed at GBT and IOZ (see Figure 4). This method was chosen because it is familiar to PTFI operations crews and can be readily implemented at the DOZ.
Drawpoint layout Draw point layout is the herringbone style. This was chosen because it allows us to use central muck raises, which is a major change from the IOZ and GBT where grizzlies were located between panels at the north and south fringe drift, outside of the orebody. This change will reduce average tramming distance by 40 metres and maintain preferred mucker orientation for most drawpoints. Right and left drawpoints in each panel are staggered to minimise open span and to reduce exposure of workers in a drawpoint to muck slides from opposite drawpoints.
Drawpoint excavation and support Drawpoints are excavated with long-term ground support and stability in mind. Each round is line drilled and trim blasted to minimise pillar damage. After each round is advanced, the heading receives 125 mm of shotcrete as primary ground support before the heading is advanced another round. After the drawpoint has been excavated one round beyond the design lintel location, permanent ground support (3.5 metre grouted rebar and monolithic concrete) is installed. This is the first time this procedure has been used at PTFI; we are pleased with the results and have started using it in other areas that previously would have been supported with timber or steel sets. One lesson learned from the IOZ is that the quality of monolithic concrete is critical to the success of the method. A rigid quality control system is used to assure that the concrete meets the design specifications. All concrete is mixed in a batch plant equipped with recording weightometers. QC engineers and technicians work all shifts to inspect the reinforcement and formwork, to measure concrete slump as it is placed, to make sure that correct vibration is used in placing the concrete, to make sure there are no cold joints, and, in general, to prevent the concrete crews from taking shortcuts. The QC engineers have the authority to stop a pour if warranted and to reject loads of concrete if they do not meet specifications.
Panel ventilation
FIG 4 - Conventional undercut.
Management has been investigating the use of alternate undercutting methods and have recently made the decision to convert to an advanced undercut method. The proposed ‘standard’ advanced undercut is shown in Figure 5.
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Extraction level design of the DOZ Mine departs from previous PTFI block cave (see Figure 6).
The panel ventilation system implemented at the DOZ is designed so that all personnel working in a panel will be in fresh air. Ventilation in the panels is by means of fresh air delivered by both the north and south fringe drifts and exhausted through exhaust raises in the center of each panels. The exhaust raises
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FIG 6 - Typical panel layouts, IOZ and DOZ Mines.
discharge on a dedicated exhaust gallery which leads directly to the exhaust mains (see Figures 6 and 7). This means that two muckers can operate at the same time in a single panel, each in its own split of air.
The system used at DOZ uses a combination of trucks and chutes to deliver the ore to the crusher.
Grizzly muck handling Muck from the central grizzly in each panel is stored in a muck raise, 4 m (13 ft) diameter, and 40 - 50 metres (130 - 165 ft) long. Average raise capacity is 1000 tonnes (1100 st). All panel muck raises bottom at chutes on the 3076 haulage level. This level is a limited access, one way traffic, racetrack type truck loop with a chute for each panel muck raise (see Figure 8). All roadways will be paved. 50 tonne (55 st) capacity trucks haul ore from the chutes to the direct dump crusher station.
FIG 7 - Typical cross-section, DOZ Mine.
This is a significant improvement over the methods used in the GBT and IOZ where fresh air is delivered at the south fringe and exhausted from the north fringe. The older method means that anyone working on the north side of the mine is always working in exhaust air and that it is almost impossible to provide enough ventilation to operate two muckers in a single panel. FIG 8 - Truck haulage level.
Haulage system The haulage system built for the DOZ Mine is a major departure from the haulage systems used at the GBT and IOZ Mines. The haulage system used at GBT and IOZ consists of a pocket under each grizzly on the north and south fringes of the orebody. Each pocket discharges on to a coarse ore conveyor, which dumps into a 1067 mm × 1219 mm (42″ × 48″) jaw crusher. Crusher discharge is conveyed to the surface.
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Crushing and conveying The haul trucks will dump directly into the Fuller-Traylor 1372 mm × 1956 mm (54 in × 77 in) gyratory crusher installed just below the haulage level. Discharge from the crusher falls into a 1800 tonne (1985 st) live capacity ore bin. The bottom of the bin is equipped with a Jaques 1829 mm (72 in) × 11 metre (36 ft) apron feeder which pulls the crushed ore from the bin and
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discharges it onto a 1524 mm (60 in), 3500 tph (3860 stph) conveyor system which delivers the ore to the mill.
Crusher station design The final excavation design can be seen in Figure 9. The excavation is based on four main levels; the 3076/L Truck Haulage Level which has the crusher assembly, truck dumping and crusher operations chamber; the 3060/L Eccentric Maintenance Level which gives access to the crusher lubrication system, eccentric maintenance cart and the top of the ore bin; the 3034/L Top Feeder Access Level; and the 3022/L Apron Feeder and D505 Conveyor Drift. A manway, cable and services raise was included for access between the Eccentric Level, the mid bin nuclear detection level and the 505 Conveyor level.
Ore bin The ultimate ore bin height was fixed by existing drifting that was to be used for the conveying system and by the crusher chamber location above. Ore bin live capacity is 1800 tons. The crusher discharges into an 8.0 m long, 6.0 m diameter transfer raise to the ore bin. The apron feeder at the bottom of the ore bin is offset from the crusher feed by 5.3 m to create a large dead bed on one wall of the ore bin. Past experience with the underground ore types and existing ore flow systems has shown that when ore is allowed to fall directly onto the feeder it can often pack solid and will not flow into the feeder opening naturally.
Nuclear detection level 3076/L truck haulage level Several factors influenced the ultimate size of the top crusher chamber, however, the driving factor was the size of the gyratory crusher itself. Enough height was needed to install a 55 tonne capacity overhead crane to assemble the crusher and maintain it over mine life. The main crusher assembly and truck dumping chamber was excavated 11.5 m wide, 12.2 m high and 34.5 m long. The maximum spans used in designing the ground support were 24 m. Smaller excavations lead off the main chamber for truck access and egress, operational control and maintenance and lower level access for men and materials. The truck dump areas required enough brow clearance to enable the 50 ton trucks to dump.
3060/L eccentric maintenance level This level required careful excavation of the crusher foundation. In addition stubs were required to house lubrication pumps, the main hydraulics and the top of the cable/manway raise. The level also gives access to the top of the ore bin.
The DOZ ore is commingled with other ores on the mine ore flow system. It is necessary that the DOZ ore flow system be integrated into the overall system PLC control network. Part of the control network are three nuclear bin level detectors in the DOZ ore bin, a high level alarm, a mid-level detector and a low-level alarm. The mid-level detector required the installation of a sublevel and manway access raise. The raise also serves to hold electric cables and services. It is vertical, 2.0 m by 2.0 m in cross section and 38 m long. The sublevel access to the ore bin and nuclear detection device is 18 m from the bottom of the raise. The second access to the Nuclear sublevel is a 25 per cent incline.
Apron feeder and Conveyor 505 The feeder elevation was established as low as possible to maximise bin capacity. The limiting factors were the necessity of flat loading onto Conveyor 505 and meeting the transfer point between Conveyors 505 and 506 that was set by existing drifting.
FIG 9 - Cross-section through crusher station and ore bin.
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System flexibility
Current system
The haulage system as installed at the DOZ is more flexible than the ones installed for IOZ and GBT. The new system can be readily expanded or changed to suit modifications in the mine plan or the discovery of additional reserves. It can also easily support additional production from other areas. A further advantage of the DOZ haulage system is that it is sufficiently far below the extraction level that it will not be effected by mining induced stresses and should be useable long after the DOZ is exhausted.
Improvements to the cave management system have been made in two areas:
Ventilation systems The DOZ Mine ventilation system includes a number of improvements over the ventilation systems built for the IOZ and GBT Mines. All shops, storage facilities, conveyor ways, crusher stations, ore dumps, compressor stations, powder magazines, etc are ventilated from a fresh airway directly to dedicated exhaust airway. All conveyor transfers and feeder stations are equipped with dust collection hoods and ducting to deliver the dust directly to dedicated exhaust airways. These two changes in design philosophy will result in a major improvement in the quality of the ventilation in the production areas. The DOZ cave line will destroy the existing main fan installations. To ensure adequate mine ventilation throughout the life of the DOZ and beyond, two new ventilation shafts 800 metres deep and seven metres in diameter are being excavated. Each shaft will be equipped with two centrifugal fans of 750 kW (1000 hp) each. In addition, existing raises to the idle DOM Mine have been converted into exhaust airways and equipped with a 450 kW (600 hp) centrifugal fan. Overall ventilation of the DOZ will be significantly better than that of the IOZ. Total ventilation throughput of the IOZ is currently 425 m3/s (900 kcfm) for a production rate of 18 000 tpd, giving a ventilation rate of 24 m3/kt (50 cfm/ton). The DOZ will have throughput of 920 m3/s (1950 kcfm) for a production rate of 25 000 tpd, giving a ventilation rate of 37 m3/kt (78 cfm/ton).
• automation of data management; and • dispatch system. Improvements to the data management system have been based on the development of a database, called Ubase, to manage all drawpoint data. Ubase contains information on each drawpoint concerning daily draw order, actual reported draw, sample data for Cu and Au, drawpoint status, drawpoint condition, wet muck status, and initial drawpoint reserve. Ubase also records shift and daily production data and conveyor belt weightometer data. Ubase provides a single platform for all operational data. Data is input only one time, the laboratory directly reports assay results over the Internet; these procedures minimises transposition errors. All daily, weekly, and monthly reports are generated from Ubase. Drawpoint history reports are much more readily run, which makes it much easier for the cave management engineers to analyse drawpoint grade trends, hangup frequencies, etc. Modular Mining Dispatch system was commissioned in December 1999. The system provides real-time reporting of mucking activity to the dispatcher and allows the dispatcher to give draw orders to the mucker operators. The dispatcher can react to changing conditions in the production area (hang-ups, closed grizzlies, BO equipment, etc) and issue changes to draw orders as needed. We have already seen improvements in the accuracy of production reports from the field and in compliance with draw orders. Direct data transfer from the Dispatch Database to Ubase has reduced data processing time, transposition errors, and given us the ability to update draw orders on a daily basis. The dispatch system was initially installed to track production muckers only; it is being expanded to include jumbos used for secondary blasting and other service equipment in the production areas. Dispatch has been installed and commissioned in the DOZ.
SUMMARY
Cave management Original system The cave management/draw control system used at the IOZ was developed at the GBT. Paper draw orders were issued twice per week to the production superintendent. The draw order told the operations supervisor how many buckets of muck to pull from each individual drawpoint on each shift. The shift supervisors issued the paper draw sheets to the mucker operators. The operators were to pull the muck as ordered and record the number of buckets from each drawpoint. It was the responsibility of supervision to correct the shift draw orders to account for hung up drawpoints, panel repair, or other operations problems. The primary method of checking the accuracy of the reports was reconciliation against belt scale tonnage reports. This allowed a good check of the total bucket count, but did not help check against the draw point report. Spot checks and field bucket counts showed that draw report accuracy was not good. The production reporting and recording system was primarily a manual system, utilising spreadsheets to manage the data. This was labour intensive and prone to errors during data manipulation and transfer. The difficulties with field reporting and office data manipulation prompted a review of available technology. As a result, major modifications have been made to the cave management/draw control systems.
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The DOZ design team had the benefit of building on the experiences and successes of the GBT and IOZ mines; but did not restrict themselves to copying existing designs. They investigated new methods and techniques and utilised them where they were found to be beneficial. Freeport Indonesia is confident that the DOZ Mine will be a reliable and economic producer.
BIBLIOGRAPHY Barber et al, 2000. Development of the DOZ Mine at Freeport Indonesia, SME. Casten et al, 2000. Excavation Design and Ground Support of the Gyratory Crusher Installation at the DOZ Mine, PT Freeport Indonesia, SME. Coutts et al, 1999. Geology of the Deep Ore Zone, Ertsberg East Skarn System, Irian Jaya, in Proceedings PACRIM ’99, pp 539-547 (The Australasian Institute of Mining and Metallurgy: Melbourne). Hubert et al, 2000. Tele – Operation Projects at PT Freeport Indonesia, manuscript. PT Freeport Indonesia, 1998, Feasibility Study of the DOZ Block Cave, Internal Document, 19980727.
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Excavation Design and Ground Support of the Gyratory Crusher Installation at the DOZ Mine, PT Freeport Indonesia T Casten1, R Golden2, A Mulyadi3 and J Barber4 ABSTRACT The Deep Ore Zone (DOZ) underground mine will be a 25 000 tpd mechanised block cave operation starting production in 2000. The ore handling system consists of a direct truck dumping 54 × 74 gyratory crusher, which feeds into an ore bin. An apron feeder below the bin is used to load a conveyor belt. This paper describes the design, sequencing and methodology of the crusher station, ore bin and feeder chamber excavations. The Geotechnical investigation and analysis of the site are also discussed as well as the design, installation and quality control of the long-term ground support methods.
INTRODUCTION The Deep Ore Zone (DOZ) is the next underground mine project at PT Freeport’s East Ertsberg Skarn System. It is located in the province of Irian Jaya, Indonesia. Current reserves stand at 131 million tons at a copper equivalent grade of 1.66 per cent. The equivalency grade includes copper, gold and silver values. The deposit is suitable for block caving operations. There are additional reserves located below the block cave horizon that may be mined in the future using existing access, facilities and ore flow systems. The mine will begin production in 2000 and ramp up to a maximum production rate of 25 000 tpd by 2003. Current mine life predictions are in the order of 17 years. A study is currently underway to investigate the feasibility of an increased maximum production rate of 35 000 tpd. An essential part of achieving such high production rates is the ore flow system. After considering the existing ore flow systems being used underground and the expected fragmentation in the DOZ it was decided to install a direct dump 54 × 77 gyratory crusher to service the mine. The excavation design, methodology and ground support required to install this size of crusher was a new challenge to the PT Freeport Underground Mines Division.
ORE FLOW SYSTEM SELECTION PT Freeport currently operates another block cave; the Intermediate Ore Zone (IOZ) mine situated 350 m above the DOZ mine. The IOZ uses a system whereby the LHD’s dump the ore at the fringes of the panels, where it is broken over a 0.4 m to 0.5 m grizzly opening by a rock breaker. There are a series of these stations located on the periphery of the orebody. Each grizzly then transfers the ore to a conveyor belt via a feeder and into one of two 42 × 48 jaw crushers. The use of rock breakers to reduce the ore to a size where it can be transported to the crusher is fairly inefficient, as it requires the installation, operation and maintenance of an in-line secondary breaking system. 1.
Chief Engineer, DOZ Mine, PT Freeport Indonesia.
2.
Chief Geotechnical Engineer, PT Freeport Indonesia.
3.
Senior Engineer, PT Freeport Indonesia.
4.
General Superintendent Engineering, PT Freeport Indonesia.
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The DOZ adopted a coarse ore system where no rockbreakers are employed on the extraction level. The basic concept is to send as much coarse sized ore as possible to the gyratory crusher for a one-time reduction to a conveyable size fraction. The typical DOZ panel design has a transfer raise centralised in the middle of the panel as opposed to the fringes for more efficient LHD tramming. The LHD’s dump into a 4.0 m diameter raise bored transfer with a 1.0 m grizzly opening. The transfer raise design is typically 45 - 50 m long with a dip of 70 - 75 degrees. The transfer raise drops the ore down to the Truck Haulage Level where loading is achieved using large hydraulic powered lip and chain gate type chutes controlled from the truck cab. The trucks direct dump at the crusher where the ore is broken to 0.2 m and is dropped into an ore bin of 1800 tons live storage. The bin is located above a variable speed apron feeder that loads the crushed rock onto the first of a series of 1.5 m wide conveyor belts for eventual transfer to the mill stockpile over 2.0 km away.
SITE LOCATION AND INVESTIGATION Once the type of ore flow system was selected, a location for the crusher was required. Several factors narrowed down this location to a relatively small area of the mine. A location centralised in the orebody, to improve truck haulage efficiencies, was a major factor. In addition, it was preferable to use as much existing development as possible for the conveying system and for access to the crusher. Considering the size and complexity of the excavation required, the ground conditions of the proposed site were also crucial. The central section of the orebody is located in relatively poor ground, making it a readily cavable area but creating a problem for the long-term support of large openings. An embayment of diorite into the southern central section of the orebody was located. Diorite is the best quality rock to be found in the DOZ area. This was also in the site set by the other criteria mentioned above. Diamond drilling through this site was undertaken from a nearby drift to provide geotechnical data, in addition to mapping the existing surrounding development. Based on these positive results and the relatively centralised location and existing development available, the current site was selected (Figure 1). shows the crusher location, Truck Haulage Level, the proximity to existing access ramps, development and the orebody footprint.
EXCAVATION DESIGN The final excavation design can be seen in Figure 2, which shows a longitudinal section. The excavation is based on four main levels; the 3076/L Truck Haulage Level which has the crusher assembly, truck dumping and crusher operations chamber, the 3060/L Eccentric Maintenance Level which gives access to the crusher lubrication system, eccentric maintenance cart and the top of the ore bin, the 3034/L Top Feeder Access Level and the 3022/L Apron Feeder and D505 Conveyor Drift. A manway, cable and services raise was also designed for access between the Eccentric Level, the mid bin nuclear detection level and the 505 Conveyor level.
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200m
N
Crusher Location
Ore Body Outline
FIG 1 - Crusher and truck haulage level location.
The mine design was done with using Maptek’s Vulcan Software package running on a series of IRIX based Silicon Graphics workstations. The ability to easily view cut sections and modify the proposed openings in 3D made the excavation and ground support design process simpler and quicker than typical two dimensional drafting methods. The basis of the excavation design is described in the sections below for each of the four levels previously mentioned. This was a highly iterative process between the mechanical, mining, ore flow and geotechnical groups involved.
3076/L truck haulage level Several factors influenced the ultimate size of the top crusher chamber, however, the driving factor was the size of the gyratory crusher itself. Enough height was needed to install a 55-ton capacity overhead crane to assemble the crusher and maintain it over mine life. Plate 1 shows the crusher main shaft in a horizontal position prior to being installed. This single component of the gyratory crusher is 6.5 m long and weighs 48 tons. The main crusher assembly and truck dumping chamber was excavated 11.5 m wide, 12.2 m high and 34.5 m long. The maximum spans used in designing the ground support were in the order of 24 m. Smaller excavations lead off the main chamber for truck access and egress, operational control and maintenance and lower level access for men and materials. These smaller chambers required brow transitions from the higher main chamber. The truck dump areas required enough brow clearance to enable the 50 ton trucks to dump at over 60 degrees.
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3060/L eccentric maintenance level This level required careful excavation of the crusher foundation. In addition stubs were required to house lubrication pumps, the main hydraulics and the top of the cable/manway raise. The level also gives access to the top of the ore bin which at this elevation was designed at 6.0 m in diameter.
Ore bin The ultimate ore bin height was fixed by existing drifting that was to be used for the conveying system and by the crusher chamber location above. Another design criterion was from Oreflow Operations who required a minimum of one operating hour of material storage in the bin. The open pit and underground mines produce in the range of 230 000 tpd, the majority coming from the Grasberg Open Pit mine. The Grasberg Mine employs a series of surface and underground conveyor belts to transport the ore to the mill. The DOZ oreflow system feeds onto this larger ore handling system at a midway point so space must be reserved for this ore on the belts. In case of a loss of production from the DOZ, the ore bin will sustain optimum belt loading while the feed rates from the Grasberg Mine can be adjusted to match the shortfall. The crusher discharges into an 8.0 m long, 6.0 m diameter transfer raise to the Ore Bin. The apron feeder intake is offset from the crusher feed by 5.3 m to create a large dead bed on one wall of the Ore Bin. Past experience with the underground ore types and existing ore flow systems has shown that where ore is allowed to fall and impact directly onto the feeder then it can often pack solid and will not flow into the feeder opening naturally.
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EXCAVATION DESIGN AND GROUND SUPPORT OF THE GYRATORY CRUSHER INSTALLATION
Scale
N
10 m
Spare Main Shaft
3076/L Truck Dump Level
Initial Access Drift into Main Chamber for Upper Lift
55 Ton Crane
Rock Breaker
Gyratory Crusher
3060/L Eccentric Maintenence Level Cable/ManwayRaise Access Ramp 3042/L Ore Bin Access/Detection Sub-Level Dead Bed
3032/L Top Feeder Access Level
Conveyor 505 3022/L Conveyor 505 and Feeder Level
FIG 2 - Long-section of crusher excavation, ore bin, feeder and conveyor 505.
Apron feeder and conveyor 505 In order to increase the ore bin size to its maximum the feeder elevation was dropped as low as possible. The limiting factor here was the necessity of flat loading onto Conveyor 505 and meeting the transfer point between Conveyors 505 and 506 that was set by existing drifting.
6.5m
GROUND SUPPORT DESIGN The site was investigated by drilling three diamond drill holes down the crusher axis, cell mapping of geologic structures and by classifying the rock mass using the Norwegian Geotechnical Institutes (NGI) Q system. Rock strength properties from previous drilling in diorite were compiled as shown in Table 1. The diamond drilling confirmed that there were no major geologic structures or zones of weak ground that would require re-siting the facility.
FIG 3 - Crusher main shaft.
TABLE 1 Rock properties used for design basis.
Nuclear detection level As mentioned above, the DOZ is part of a larger system of ore flow from the mine and keeping the level of ore in the bin at a relatively consistent height was determined to be an important factor for long-term operations. In order to achieve this control, an ore level detection device was required at three locations. A high level alarm, a mid-level detector and a low-level alarm. A mine standard is the use of nuclear source and detection devices. The high and low level alarms could be installed within the existing drifting. The mid-level detector required the installation of a sublevel and a manway access raise. The raise also had to hold electric cables and services. It was designed vertically at 2.0 m by 2.0 m in cross section and 38 m length. The sublevel access to the Ore Bin and nuclear detection device was located 18 m from the bottom of the raise. In order to allow mobile equipment to gain access to the Ore Bin and the sublevel the other access was designed as a 25 per cent incline.
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Rock properties
Average
Max
UCS (MPa)
111
197
Min 66
Density (tons/m3)
2.78
3.01
2.63
Tensile Strength (MPa)
11.7
14.9
5.5
Analysis using Q (Table 2) indicated that excavation support was feasible and would require systematic bolting and shotcrete. Analysis using the modified Mathew’s method confirmed this result. Owing to the complexity of the excavation every wall and back had a unique orientation and dimension. Ground support was tailored to every section. Stereo plots of geologic structures were created from the cell mapping data. Ground support for walls and
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T CASTEN, R GOLDEN, A MULYADI and J BARBER
TABLE 2 Q rock classification derived from local mapping (Barton, Lien and Lunde, 1974).
Standard Length and Spacing Colors
NGI Tunnelling Quality Index (Q) Parameters Rock Quality Designation (RQD):
E = 6.0
Joint Roughness Number (Jr):
B = 3.0
Joint Alteration Number (Jn):
B = 1.0
Joint Water Factor (Jw):
A = 1.0
Stress Reduction Factor (SRF)1:
L = 5.0
NGI Index Value: (Fair)
Q=8 ESR = 1.6
Design Span (m)
24.0
Equivalent Dimension (De)
15.0
NGI Support Recommendation Category3:
19
1. SRF – Mild Rock Burst Potential 2. ESR – Permanent Mine Opening 3. NGI Support Recommendation – ‘Bolts (tension grouted) on 1 to 1.5 m centres with mesh reinforced shotcrete 5 to 10 cm thick. Tensioned cable anchors are often used to supplement bolt support pressures. Typical spacing 2 to 4 m’.
brows was designed by analysing the stability of rock blocks and wedges created by the intersection of geologic structures with the free face. For walls where no adversely oriented structures were anticipated, support recommendations from Q were used. Support was designed for each back using a beam concept. Resulting bolt lengths varied from 3 m to 8 m, with bolt spacing in a range from 1.0 to 1.8 metres. Additional ground support was later designed to accommodate large static and dynamic loads applied by the crusher to the crusher hopper and foundation areas. The planned ground support was assumed to be 20 mm grouted thread bars with yield strength of 17 tons. The design assumed that the grout would provide an anchorage strength of 10 tons/m of embedded length. In reality, pull-out tests conducted on thread bar embedded in 0.2 m to 0.5 m of grout provided anchorage strengths in excess of 20 ton/metre after 14 days (Rachmad, 1998). These estimates allowed a preliminary design of the total support requirements to be made prior to mining into the area in question. Material quantities for bolts, grout and equipment with long lead times could then be estimated and ordered. For the entire excavation approximately 3600 bolts were required for a total of 18 400 m of installed 20 mm threadbar bolts in 45 mm and 52 mm drilled holes. The average bolt length was 5.1 m with a range between 3 m to 8 m depending on location. In order to simplify the ground support process and ensure that the correct length and orientation of bolts was achieved a series of drawings were created for distribution to the operators underground. Each drawing showed the excavation outline and two bands of colour representing bolt length and spacing. The drawings also showed specific information for bolt installation angles, maximum spans, and spacing. An example of this is shown in Figure 4, which shows a cross-section of the excavation, standard length and spacing colours and the typical bolt information and examples of ideal bolt installation. The bolt design was varied based on joint set orientation. This can be seen when comparing the two sides of the excavation shown in the figure.
298
Specific Information on Dip, Bolt Length, Spacing, Bolt Orientation and Span
80
Joint Set Number (Jn):
Excavation Support Ratio2:
11.5m
FIG 4 - Cross-section of the upper crusher chamber showing bolt length, spacing and orientations.
Grouting was performed using a toe-collar method and a very stiff grout mix, with a water to cement ratio of 0.3 - 0.35. The bolts were inserted into the hole after the grout was placed. A galvanised welded wire fabric mesh with 100 mm by 100 mm openings was used with the threaded bar. This mesh was pinned tight to the back and ribs between the grouted bolts using 2.1 m galvanised split set bolts to prepare for shotcrete application. Two GP 2000 grout pumps were purchased from Master Builder Technologies (MBT) Australia for the grouting activities. A grouting additive from MBT was also used. The ‘Flowcable’ additive helped to achieve high early strengths and reduce shrinkage. Shotcrete was applied to the majority of the excavation over the top of the grouted threadbar and mesh. The thickness varied between 5 to 15 cm. The batching of the shotcrete was performed on surface and trucked underground over 2.0 km in agitator trucks that held 3.5 cubic metres. A plasticiser and accelerator were added at the pump, prior to spraying. Facilities to test shotcrete panels were unavailable on site, however, cylinders were taken for uniaxial compressive strength (UCS) testing. After seven days the average UCS results were 34.6 MPa and after 28 days they were 45.6 MPa. This exceeded the designed specification of 41.4 MPa (Latief, 1998). Owing to the complexity of this project a field engineer was assigned full-time to the excavation. His duties included checking to ensure that the correct bolting specifications for the particular area were being used and working directly with the miners where field modifications had to be made to bolting patterns.
DEVELOPMENT METHODOLOGY The excavation described in the previous section was staged to ensure that the spans opened up remained manageable and access to the larger chambers was maintained as required. Owing to the complexity of the excavation and the nature of the ground support required in this area the mine contracted the work out to PT Redpath Indonesia, an on-site contractor. As of writing all of the excavation and support activities have been completed except for the ore bin support, which is still underway. A description of the excavation methodology and support sequencing done to-date follows.
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EXCAVATION DESIGN AND GROUND SUPPORT OF THE GYRATORY CRUSHER INSTALLATION
The overall strategy was to excavate and support the upper crusher sections using contractors while simultaneously excavating the feeder and conveyor areas below using Freeport personnel. The Ore Bin would be excavated last. This schedule allowed the Construction Department to start as early as possible on the crusher, conveyor and feeder installations. Pre-split blasting was used throughout the excavation to reduce overbreak and blast damage. In order to access the top section of the 12.2 m high Crusher Chamber this area was done in two lifts. A 13 per cent ramp was driven on the same bearing as one of the truck access drifts so that the main chamber was initially accessed 6.0 m from the final back. This allowed the standard scissor lifts, jumbos and loaders to be used. This access drift and the Upper Lift can be seen in Figure 2. Grouted thread bar bolting and shotcrete were installed as the upper chamber progressed from the North to the South. At the same time the central access from the South to North was being driven, 6.2 m below the top lift floor. A breakthrough from the upper lift to the lower was made. This improved ventilation in these fairly large openings. Once the long-term support was completed in the upper chamber, including a series of pad eyes required for construction activities later on, the upper section was retreat benched into the lower section. The Upper Feeder Access drift was driven and, simultaneously the lower section of the Ore Bin was excavated and supported. The sill pillar between the 505 Conveyor drift and the Feeder was then excavated to final size and supported. Development in this area was then placed on hold until the feeder foundation steel and Ore Bin bottom concrete were installed by the Construction Department. Once this was complete, mining activities recommenced as the Construction personnel were protected by a cap of steel and concrete above them. The next stage in the excavation was the 25 per cent inclined access ramp up to the Nuclear level detection drift. A steep ramp was employed here instead of a raise so that mobile equipment access could be maintained for development of the whole sublevel. Once the sublevel elevation was achieved, a drift was driven directly through the proposed Ore Bin to the other side of the sublevel and continued onto the cable/manway raise.
CONCLUSION The future of the DOZ mine relies on the quality of the Ore Flow system that has been constructed. In December the crusher will be commissioned and the ore flow system will be active. Although full production is not expected until mid-2000 the DOZ development activities produce over 500 000 tons of material annually that requires removal to the surface. The ore flow system will used to send the majority of the development material to the stockpile and will reduce the truck haulage costs. The excavation and support for the Upper Crusher Chamber described above began on 19 October 1998 and was completed by 25 March 1999. The lower section, including access, sumps, conveyor drifts, feeder and the ore bin access excavation were driven over different periods, with interruptions due to scheduled Construction activities. Initial development began on 1 June 1998
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and it is predicted that the crusher will be operational by 15 December 1999. A total of 2478 m of equivalent standard drift was driven during this time. Involving all of the groups involved with this project in the concept, design and implementation stage has given the project a lot of ‘buy in’ from the different areas of the mine. On-site suppliers were included in the design phase and a lot of good ideas were generated from these discussions. Having a dedicated field engineer on this project was a tremendous benefit to the eventual quality and speed of the excavation. Minor issues could often be resolved in the field allowing work to continue. A full time Survey crew could not be dedicated to the excavation due to other work commitments. Having a consistent survey control and feedback would have been an advantage.
ACKNOWLEDGEMENTS Numerous people were involved in the success of this project and credit needs to be given to several individuals. Kamaruddin, Martin Thomas and Farhat Shahab of PT Freeport Engineering for the mechanical, civil and foundation work. Roger Cayouette, and Bruce Mennie, PT Redpath Indonesia, should be recognised for the professionalism shown by their crews and for the time saving ideas produced during the excavation process. Joko Basyuni, Superintendent of Development, PT Freeport Indonesia for his continued support during the development phase. Harry Mostert and Warren Mahoney, Master Builder Technologies Australia, for their input and ongoing support of the grouting and shotcreting process. David Starke, Ani Arnall Australia, for his input into the ground support process and assistance in material selection, installation methodology and testing. Dave Nicholas, Call and Nicholas Incorporated for his assistance in the overall design of the ground support.
REFERENCES Barton, N, Lien, R and Lunde, J, 1974. Analysis of rock mass quality and support practice in tunneling, and a guide for estimating support requirements. Golden, R, (Chief Geotechnical Engineer), 1998. PT Freeport, Internal Report Preliminary Design to Support the DOZ Crusher Excavation. Golden, R, (Chief Geotechnical Engineer), 1998. PT Freeport, Internal Report Revised Design to Support the DOZ Crusher, Orebin and Feeder Chamber Excavations. Gallagher, J and Teushcer, J S, 1998. Henderson Mine: Preparing for the Future, in Proceedings SME Conference, Orlando, Florida, p 5. Keskimaki, K, 1998. Personal Communication, Henderson Mine, Climax Molybdenum Company. Nazaruddin, L, (Senior Quality Control Engineer, PT Freeport), 1998. Internal Report Compressive Strength Test Report, p 15. Rachmad, L, (Senior Rock Mechanics Engineer, PT Freeport), 1998. Internal Report Rock Properties for DOZ Crusher Design Check, p 2.
Brisbane, Qld, 29 October - 2 November 2000
299
Commissioning of Two 750 kW Centrifugal Fans at PT Freeport Indonesia’s Deep Ore Zone Mine F Calizaya1, T Casten2 and K Karmawan3
This system includes a set of dedicated drifts, shafts and raises developed to handle the total quantity of air. The drifts are located on five levels: 3971/L (top of vent shafts), 3646/L (intermediate accesses), 3186/L (bottom of vent shafts), 3140/L (exhaust galleries), and 2976/L (conveyor drifts). Except for the 3646/L access ways, all drifts were developed to a finished cross-section of at least 5 m x 5 m. The access drifts (3 m x 4 m) are two old conveyor drifts that are now used for ventilation. The 3971/L drifts are the main exhaust ways driven at the bottom of the Dom copper deposit (DOM). They connect the DOZ vent shafts to the surface. They are located 860 m above the intake drifts. The inner walls of these drifts were reinforced and shotcreted to minimise the resistance to airflow. The shafts were developed in two stages, first by boring to 2.74 m in diameter then slashing to seven metres. During the development, the shafts intercepted blocky ground that required special support. Grouted threadbar and wire mesh were used to support them. Figure 1 shows a line diagram of major openings and the manner in which the contaminated air is exhausted and discharged to the surface. It also shows the location of the
LEGEND Centrifugal Fan Airlock Door
LEVEL
Bulkhead Airflow Direction
Shaft 2
3971
R
Regulator
VR
Vent Raise
XC
Cross cut
3646 Shaft 4
The Deep Ore Zone (DOZ) is a new ore deposit added to the PTFI mine reserves recently. The deposit is located in the Province of Irian-Jaya, Indonesia. It includes over 131 million tons of copper, gold and silver ore at 1.66 per cent copper equivalent grade. Based on the ore reserves, the rock/ore characteristics and the geologic conditions of the surrounding rocks, the deposit will be mined using a block caving method. At present, the mine is being developed on five levels by conventional means of drilling, blasting and mucking. Numerous headings are developed on each level. From each heading, the broken rock/ore is hauled by means of diesel-powered units, which require large volumes of dilution air. The fresh air is supplied through the DOZ intakes and the Mill Level Adits (MLA). The contaminated air is directed to two vent raises: Borehole #3 and the DOZ Exhaust Raise. Both raises are connected to surface and equipped with 450 kW fans. Based on the company’s medium term plan, the production stage is to start in the fourth-quarter of year 2000 at 2000 tpd and will ramp up to 25 000 tpd by year 2003. The airflow requirements for this mine were determined based on the Indonesian Mine Regulations, the MSHA 30 CFR 57 regulations and the experience gained from similar activities in the Intermediate Ore Zone (IOZ) block cave mine. When applicable the following design factors were used a minimum air velocity of 0.76 m/s (150 fpm) and a minimum air quantity of 7.9 m3/s per 100 kW (125 cfm/bhp). To produce 25 000 tpd of ore and maintain safe and healthy working conditions, the total airflow requirement was estimated at 780 m3/s (1 652 000 cfm). This is strictly fresh air and does not include the leakage flowrate. For the projected airway and leakage path conditions, it is estimated that the total quantity exhausted by the main fans will increase to 996 m3/s (2.1 million-cfm). This quantity was reviewed by an independent auditor and found to be reasonable (MVS, 1999). To supply this quantity, the vent system will require five centrifugal fans: one 450 kW fan (currently used in IOZ) and four 750 kW fans. This
MINE EXHAUST SYSTEM
Shaft 1
INTRODUCTION
paper summarises the basic principles involved in the installation, testing and the subsequent commissioning of two of these fans.
Shaft 3
ABSTRACT The PT Freeport Indonesia‘s Deep Ore Zone (DOZ) copper deposit is located in the province of Irian-Jaya, Indonesia. Currently, a block cave mine is being developed around this deposit. Production is scheduled to start the fourth-quarter of year 2000 and to ramp up to 25 000 tpd by year 3 2003. The air quantity requirement for this mine is equal to 780 m /s. To supply this quantity, the mine will require five centrifugal fans (one 450 and four 750 kW fans). This study summarises the technical aspects related to the installation and commissioning of two of these fans. The fans were installed on the surface in parallel arrangement with two 3 others. Each has a design capacity of 240 m /s of air at 2.2 kPa of static pressure. The study presents the installation details from the underground shaft excavation and portal development through fan testing and commissioning. It also presents the results of a monitoring system that was installed to determine the fan duties.
3186
DOZ Intakes
1
Chief Ventilation Engineer, PT Freeport Indonesia.
2.
Chief Engineer, DOZ Mine, PT Freeport Indonesia.
3.
Senior Ventilation Engineer, PT Freeport Indonesia.
Exhaust Level
3140
FIG 1 - DOZ mine exhaust system.
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F CALIZAYA, T CASTEN and K KARMAWAN
exhaust fans required for this project. Figure 2 shows a plan view of the DOZ exhaust system on the 3971 level. It shows the general arrangement of the main fans and their position in relation to the exhaust drifts and shafts. This paper deals with the installation and commissioning of two main fans (12 and 15). Fans 13 and 14 will be commissioned in year 2001. Figure 3 shows a view of fans #11, 12 and 15. It also shows the future site of Fans #13 and #15 at the DOZ exhaust portals.
INSTALLATION OF TWO 750 kW CENTRIFUGAL FANS These are two identical WBF3200 Howden Sirocco fans of single inlet type. They are located at the DOZ Exhaust Portals at an elevation of about 3971 m above sea level. The installation for each includes the construction of an attachment chamber, the erection of fan and motor foundations, the installation of a 750 kW fan and its subsequent testing and commissioning. The technical drawings were provided by the vendor (Howden, 1997; PTFI, 1998) and the installation jobs were carried out by PT Freeport’s Construction Department. Figure 4 shows the longitudinal and end views for one of these fans. It shows the main fan components and the required foundations. It also shows the position of this fan in relation to the mine portal.
Fan components Each fan consists of an inlet duct, a casing, an impeller assembly, an expansion cone and a 750 kW induction motor (Howden, 1997). The inlet duct includes a square-circular transition piece and a pair of self-closing doors at the inlet end and a radial vane controller (RVC) at the outlet end. A flexible joint is used to connect the inlet duct to the RVC assembly. The RVC is used to regulate the fan output. It is operated by means of an electric AUMAR actuator. The fan casing is made of fabricated steel plate
DOM Service Drift
VR 1
VR 2
Fan 11 DOM Exhaust 12 14
Shaft 2 Exhaust Drift 1
13 15
Shaft 1
Exhaust Drift 2
Fig.2a. DOM Exhaust System - Plan View
Fan 12
Evase Inlet Duct Self Closing Doors
Portal
750 kW Motor Impeller
RVC
Fan 14
Drift 1
25 o
Airlock Doors
Attachment Chamber Fig.2b. Fan Installation Details
Fig 2 - DOM fan system layout.
FIG 3 - DOZ mine exhaust system - DOM portals, Fans 11, 12, 13, 14 and 15.
302
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COMMISSIONING OF TWO 750 kW CENTRIFUGAL FANS
Evase’ Flexible Joint Self-Closing Doors
Rotation
Bearings
RVC
Motor
Portal Motor Plinth
Longitudinal View
Drive End View
FIG 4 - 750 kW fan details – longitudinal and end views.
reinforced with stiffeners. It is fabricated in two sections, the lower and upper casing halves. An expansion cone (evasé) is bolted to the casing discharge flange. The impeller assembly consists of an impeller, a shaft, two bearings and a coupling. The impeller is the rotating portion of the fan. It consists of an inlet ring, a hub and 11 blades. The fan shaft supports and rotates the impeller and is connected to the motor by a flexible coupling. The shaft is supported by two grease lubricated spherical roller bearings, the drive-end and the non-drive end. The drive end bearing is fitted with locating rings to prevent axial movement. The non-drive end bearing is free to move axially to accommodate shaft expansion. A foot-mounted 750 kW-12 pole induction motor drives the fan. This is a totally enclosed air-cooled motor. It produces a constant speed in the fan shaft to which it is connected by a Fenner Wellman Bibby type coupling. The fan motor starter is of the autotransformer type, designed to accelerate the motor at 80 per cent terminal voltage. Table 1 shows the technical specifications for one of these fans.
TABLE 1 750 kW centrifugal fan specifications. Fan Details Type and Size
Centrifugal, WBF3200SIRVC
Impeller
Eleven Backward aerofoil bladed
Shaft
Machined from normalised carbon steel
Inlet
Circular, single axial duct
Effective diam
3.2 m
Control
RVC with Electric AUMA Actuator
Bearing
SKF grease lubricated, spherical roller type with adapter sleeves
Coupling
Fenner-Wellman Bibby type 2140 H
Self-closing Doors Two semi-circular doors fitted in inlet duct Main Drive
Installation sequence In each case, the fan installation started with the civil foundation work. This included the construction of an attachment chamber and the foundations for the self-closing doors, inlet duct, fan casing and motor. Next, the lower casing, the motor base plate and the shaft/impeller assembly were mounted. Then, the radial vane control, upper casing, inlet duct, and the motor components were installed. The work was completed with the installation of an electric AUMA actuator, the discharge duct and two airlock doors. A great deal of time was spent on the fan/motor shaft alignment to bring this to within the 0.001 - 0.002 mm range (allowable limit: 0.05 mm). From previous installations it was found that shaft misalignment could result in excessive vibration, shaft fatigue, bearing wear and increased power consumption.
Technical Data and Specifications
Monitors: Pressure
TECO 750 kW electric induction motor, 3 phase, 60 HZ, 4160 V.
Flowrate
Rosemount 1151DP. Static pressure range: 0 - 3 kPa range Rosemount 3051CD. Quantity range: 0 - 500 m3/s
Bearing Temp
RTD Type sensor
Vibration
Schenck type VS-068 connected to model 920 vibration monitor
Fan Duty: Inlet Volume
240 m3/s
Static Press.
2.2 kPa
Fan/motor Speed
590 rpm
Testing and commissioning of Fan 12 At start-up, some difficulties were experienced with the Siemens vacuum contactor, which required re-alignment of the coil and armature and adjustment of the vacuum bottle gap. This was rectified, the run-up time set to 40 s and the motor coupled to the fan. Owing to the low fault level at the DOM substation, the supply voltage dip was excessive, resulting in a longer acceleration time than is desirable. When this problem was solved, the fan functioned satisfactorily for more than one hour.
MassMin 2000
It was switched off due to obstructions in the intake drift. When the drift was cleared, the fan was tested successfully with the RVC aperture at 15 per cent open (Harrold and Snaith, 1999). When the second portal was expedited and the 3971 drifts connected to the fan inlet (partial load), the fan was tested with the RVC aperture at 100 per cent open. At start-up, the fan ran smoothly. However, half an hour into the start, the fan showed
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F CALIZAYA, T CASTEN and K KARMAWAN
substantial vibration pulses in the discharge duct. The pulses were emanating from the discharge evasé and the fan casing. The fan had to be stopped and the vendor contacted. Howden Sirocco recommended installing a set of channel stiffeners onto the evasé. This reduced the vibration in the discharge duct, but the problem remained in the casing. To overcome the problem, two horizontal channel stiffeners were welded across the full width of the fan casing (drive end only). Following this modification, the fan was tested again on 12 May 1999 and left running overnight with the RVC at 100 per cent open. Table 2 shows the results achieved. An evaluation of these results shows that all measurements were below the critical alarm levels (5.5 mm/s for vibration and 85oC for temperature). The fan was commissioned at full load on 3 February 2000. For these conditions, the fan duty was equal to 280 m3/s at 2.5 kPa of static pressure.
was heard from both bearings. The fan was shut down immediately. Now, it was evident that the bearings had been damaged due to the ingress of water through the grease nipple holes, which were left unplugged for more than two months. At this point it was decided to replace the bearings. This job took several weeks to complete. Following the bearing replacement, the 3971/L drift was expedited and the fan tested at partial load. This time, the bearing temperatures increased slowly reaching a maximum of 69oC ten hours after the start and stabilising around 64oC. Twenty hours into the start, the following results were recorded:
• Bearing Temperatures; Fan End: 51oC; Motor End: 64oC; • Bearing Vibration: 0.88 mm/s; and • Flowrate: 349 m3/s at 1.83 kPa of static pressure; Input current: 131 Amp. Based on the above results (all below the critical alarm levels), the fan was commissioned on 30 May.
TABLE 2 750 kW fan testing results. Parameter
1 hr into Start
22 hr into Start
Fan Bearing Vibration Horizontal, mm/s Vertical, mm/s Axial, mm/s
DE1 0.25 0.30 0.15
IE2 0.85 0.35 0.25
Motor Bearing Vibration Horizontal, mm/s Vertical, mm/s Axial, mm/s
DE1 0.12 0.05 0.03
NE3 0.16 0.06 0.06
Motor Speed, rpm
596
596
Stator Temperature Sta 1, oC Sta 2, oC Sta 3, oC
41 40 40
48 46 47
Bearing Temperature Motor End, oC Fan End, oC
22 16
45 35
Fan Duty Static Pressure, kPa Flow Rate, m3/s Input current, Amp Fan efficiency, %
0.94 320 113 72
0.95 321 113 72
Normal Air Data Barometric pressure Dry bulb temperature Wet bulb temperature Air density
63.8 kPa 9oC 7oC 0.78 kg/m3
1.
Each fan is equipped with two shaft bearing resistance temperature detectors (RTD’s) and one Schenck 920 fan vibration monitor. Each is equipped with a digital display monitor and two alarm and trip set points. The fan duties are monitored by means of Rosemount pressure and volume flow transducers. These are also equipped with digital readouts. The monitors are mounted into the fan starter cubicle. The output signals are transmitted via a PLC system to the Mine Oreflow Computer Control Room. The data is then processed and displayed on remote screens. The fan is monitored continuously at the Control room where the operator is trained to respond to emergency cases. The animation part is written in Wonderware software and linked to the Freeport Network System. Figure 5 shows a sample screen display of the fan monitoring system. Eventually this data screen will be available to all users of the Freeport Intranet.
CONCLUSIONS AND DISCUSSIONS
Drive End;
2.
Impeller End;
3.
Non-drive End
Testing and commissioning of Fan 15 Siemens vacuum contactor problems were again experienced at this fan on 3 February 2000. The 3TL810 ‘run’ contactor would not energise at start. The contactor magnets had to be realigned to solve this problem. Then, the changeover time was set at 47 s to regulate the starting current. Following this, the motor was coupled to the fan and the fan operated with the RVC at 100 per cent open for one hour. During that time it was observed that the bearing temperatures were rising more rapidly than normal. The fan was stopped and the bearings inspected. The clearances were rechecked and the multi-purpose grease changed to Mobilith SHC100 Synthetic grease. The change did not improve the results significantly. The bearing temperatures reached 80oC in less than two hours. This time, an audible rolling scraping sound
304
Fan monitoring and instrumentation
During the fan installation, it is essential to have the fan and motor shafts aligned properly. Misalignment can result in excessive vibration, bearing wear, shaft fatigue, and increased power consumption. Every effort must be made to achieve alignment readings of less than 0.05 mm. In this case, both angular and offset measurements fell between 0.001 - 0.002 mm. The bearing temperatures are key indicators of the fan health. If the bearings are installed properly, ie with the right clearance and settings, the bearing temperatures should never exceed the alarm level of 85oC. In this case, the bearing temperatures for fans 12 and 15 stabilised at about 47oC and 64oC respectively. The commissioning of Fan 12 was delayed twice due to an excessive vibration on the casing and evasé. The installation of horizontal and vertical stiffeners solved the problem. The utilisation of stiffeners is crucial to the fan design, especially if the system is to handle large volumes of air. The commissioning of Fan 15 was simpler than that of Fan 12 except for some bearing problems. The unplugged greasing holes allowed the ingress of water to the bearing housings, thus causing internal damage. The bearings had to be replaced to solve the problem. When the return airways were expedited and Fan 12 powered, the fan operating point was equal to 280 m3/s at 2.5 kPa of static pressure. This point is within the elevation adjusted fan curve. When the new exhaust shafts are commissioned and Fan 12 operated at full load, it is predicted that each fan will exhaust approximately 250 m3/s of air at 2.3 kPa of static pressure.
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COMMISSIONING OF TWO 750 kW CENTRIFUGAL FANS
120
478 22
27
0.86 mm/sec
1170
100%
FIG 5 - Screen display of the fan monitoring system.
ACKNOWLEDGEMENTS The authors gratefully acknowledge the cooperation of the following people: J Herr, K Bok, B Sihombing, D Somad and W Kawengian for their assistance during the installation and commissioning of both fans. The participation of John Snaith and David Harrold from Howden Sirocco throughout the fan installation is also acknowledged.
REFERENCES
Howden, 1997. Description, Installation, Operating and Maintenance Instruction Manual, DOC C0345-01-OMI, Howden Sirocco Pty, Limited, 97-103 Pacific Highway, North Sydney NSW 2060, Australia, pp 1-1, 5-15. MVS, 1999. Ventilation System Review of PT Freeport Indonesia Company’s DOZ Underground Mine, Mine Ventilation Services, Inc, 4946 E Yale Ave, Fresno CA 93727, 13 p. PTFI, 1998. Mine Documentation Control, Engineering Department, DOM 1000 HP Howden Fan Drawings, PTFI, Technical Services, Drawing Register, Tembagapura, Irian-Jaya, Indonesia, Drawing No 03450000- 03452801, 19 p.
Harrold, D and Snaith, J K, 1999. Commissioning Vent Fan – Maintenance and Spares Audit Report, Contract Ref CS 3254, Howden Sirocco Pty Limited, 97-103 Pacific Highway, North Sydney NSW 2060, Australia, 10 p.
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A Case History of the Crusher Level Development at Henderson M F Callahan1, K W Keskimaki2 and W D Rech3 ABSTRACT As an integral part of the conversion from rail haulage to conveyor haulage at Henderson, over six kilometres of drifting and several large complex openings were mined. The deepest of these areas was nearly two kilometres beneath the surface, bringing all of the challenges of rock stress and heat that come with mining at these depths. Though many mines operate at these depths, the development of the infrastructure necessary to support a 36 000 tonne per day caving operation brought some unique challenges to the development mining activity. The issues handled include: finding experienced staff for a ‘fast track’ development program; rock bursts and ground conditions; inflow of 50 degree centigrade water; acceleration of the project completion; large diameter raise boring and difficult muck handling.
INTRODUCTION The Henderson Operation has been a producer of high-quality molybdenum concentrate for the past 24 years. The rail ore haulage system, used since the start of production, was in a state of severe obsolescence and was experiencing declines in performance. Rail haulage system maintenance costs were increasing and the condition of the equipment posed a threat of 1.
Mining Coordinator, Henderson Operations, Climax Molybdenum CO, PO Box 68, Empire CO 80438, USA. E-mail: [email protected]
2.
Technical Services Manager, Henderson Operations, Climax Molybdenum CO, PO Box 68, Empire CO 80438,USA.
3.
General Manager, Henderson Operations, Climax Molybdenum CO, PO Box 68, Empire CO 80438,USA.
serious production interruption or shutdown of the mine. After several months of evaluation, it was recommended that in order to continue economically viable production, the existing locomotive fleet needed replacement. In August of 1996, Henderson initiated a project to replace the old 24-kilometre rail ore haulage system with an underground crush and convey system. A portion of this project was the mining of 6454 metres of declines and accesses for a new crusher level and excavation of over 28 700 cubic metres for the underground crusher and conveyor transfer station. Mining started in September of 1996. All mining and excavating was complete in April 1999. Figure 1 is a general schematic of the new crush and convey system and the associated mining required for the project.
HENDERSON OPERATIONS SCOPE OF WORK Decline mining Three main declines were mined to reach the crusher level. DC1 decline for access from the current production level and to establish a development truck dump feed to the old train system, LA decline for access from the main shaft to the top of the crusher and PC1 decline for access to the reclaim of the crusher and installation of the first of the underground conveyors. In addition, drifting around the crusher dump was required for transport of the ore to the crusher. Figure 2 is an overview of the decline mining required to reach the new crusher and for installation of the first conveyor belt.
FIG 1 - Crush and convey schematic.
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LA DECLINE
CRUSHER
#2 SHAFT PC1 DECLINE PC1/PC2 TRANSFER STATION
DC1 DECLINE
PC2 CONVEYOR (17KM)
#2 SHAFT DC1 DECLINE
LA DECLINE
PC1/PC2 TRANSFER STATION
PC1 DECLINE
CRUSHER
FIG 2 - Decline mining lay-out
Excavations Two large excavations and several smaller excavations were necessary for the project. An 8200 cubic metre opening for the underground conveyor transfer station and a 20 500 cubic metre opening for the crusher complex. Smaller muck transfer,
pumping, explosives and caps, storage and truck loading cutouts were also mined. Figures 3 and 4 show the detail layouts of the PC1/PC2 transfer station and the crusher. Figure 5 photo demonstrates the scale of the 20-metre by 18-metre by 26-metre high underground conveyor transfer station.
FIG 3 - PC1/PC2 Transfer station lay-out.
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TRUCK DUMP
7065 HAUL LEVEL CENTER LOADING CHUTE
CRUSHER DECLINE
72 TONNE HUAL TRUCK CRUSHER
7017 LUBE ROOM
ACCELERATOR BELT
APRON FEEDER
PC1 DECLINE
PC1/PC2 TRANSFER STA 1.3 KM BELT TURN-OVER
6930 RECLAIM GALLERY
6920 DRAIN SUMP
FIG 4 - Crusher lay-out.
FIG 5 - Drilling bottom bench in PC1/PC2 transfer station.
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Rock bursts and extraordinary ground support
CHALLENGES
Dealing with rock bursts and poor ground conditions increased required mining time and costs. There were four major events, including three rock bursts, and several smaller areas of poor ground. When ground problems were encountered, the mining process was changed to: blast a short round; wait 12 or 24 hours; support heading with welded wire fabric and long bolts; then apply shotcrete. A new shotcrete batch was developed with Henderson personnel and outside vendors to increase application thickness, with less than eight per cent rebound and 3000 psi strength or higher. As a result, the ability to support the large excavations and poor ground condition areas was improved. Mining in all headings was cautious after the first incident. Figure 8 shows the locations of each incident.
Staffing and efficiencies The largest mine crew sizes were 12 people each, with four crews rotating in order to cover 24 hours a day and seven days per week. The total staff (including maintenance, concrete/shotcrete, supplies and electrical) reached a maximum of 93 people. Hiring qualified underground personnel (or even skilled laborers) was difficult and training was an on-going issue. Total turnover for the mine crews averaged five percent per month. Achieving forecasted rates of mining and excavations was hampered by the ground problems, water inflows and raise boring problems. Over 80 per cent of the critical path was in the form of decline mining. Each delay in cycles led to a corresponding reduction in efficiency, and, in fact, all water problems immediately impacted progress. Only with constant reaction to changing conditions were the mine personnel able to complete the mining with the staffing available. Safety was always the top priority in all phases of the project, and proper training and hazard recognition skills were constantly emphasised. One lost time incident and nine minor injuries occurred, which is low for a large project of this type. In Figures 6 and 7, each graph displays a corresponding decrease in development rate as difficulties are encountered. These rates were tracked for weekly meetings in order to promptly devise methods to increase development rates without shorting the safety or quality of the mining produced.
1.
Ground repair in PC1/PC2 substation area and access – After starting development into the PC1/PC2 transfer station, sloughing of the ribs occurred along the access decline and a cut-out designed for installation of the transfer station substation. A program to clean up, re-bolt, install mats and re-shotcrete the areas were executed in March and April 1998.
2.
Rock burst in PC1 drift – An area of approximately 30 metres by five metres high was displaced when a rock burst occurred in PC1 drift just below PC1/PC2 Transfer station in April 1998. The drift was cleaned up, re-bolted, mats installed, re-shotcreted and a retaining wall built and pumped. The heading was shutdown for the entire month of April for repair. Prior to repair, core holes were drilled to study and plan for the remainder of the mining. Henderson
Rock Burst and Raise Boring Problems
Hiring, Training and Water Problems
Average Meters / Day
10.0
8.0
Rock Burst in Crusher ramp
12.0
6.0
4.0
2.0
Average Actual Meters / Day
Mar-99
Jan-99
Nov-98
Sep-98
Jul-98
May-98
Mar-98
Jan-98
Nov-97
Sep-97
Jul-97
May-97
Mar-97
Jan-97
Sep-96
0.0
Nov-96
Forecast Meters / Day
Fig 6 - Decline mining rates.
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120.0 Average Cubic Meters / Day
Rock Burst at Crusher ramp
Ground repair and Rock Burst at Transfer station
140.0
100.0
Crusher Excavations Bolting
160.0
80.0
60.0
40.0
Apr-99
Mar-99
Feb-99
Jan-99
Dec-98
Nov-98
Oct-98
Sep-98
Aug-98
Jun-98
May-98
Apr-98
Mar-98
Feb-98
Jan-98
0.0
Jul-98
Average Actual Cubic Meters / Day Forecast Cubic Meters / Day
20.0
FIG 7 - Excavation mining rates.
#2 SHAFT
1
LA DECLINE
DC1 DECLINE
3
2 PC1/PC2 TRANSFER STATION
PC1 DECLINE
4
5
6
CRUSHER
FIG 8 - Map of difficult ground support locations.
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staff and consultants developed a three-dimensional geomechanical stress model. Then outside geotechnical experts were consulted on a plan for repair and advance. Mining was re-started in May 1998, with the process of shooting a three-metre round, wait 24 hours, bolt and mat the heading, then shotcrete. The heading was completed to within 20 metres of the Transfer station, and the remainder was mined from the Transfer station. Figure 9 is a photo showing the displacement of ground on the right rib of PC1 drift below the transfer station and Figure 10 shows the structure within this area. Owing to these two incidents, an exacting bolting plan was developed for the PC1/PC2 Transfer station area. A series of resin grouted hollow core bolts were installed in a 1.5 x 1.5 metre pattern in two-metre, three-metre, 4.5-metre or six-metre lengths dependent on local rock structure. Figure 11 illustrates a typical ground support plan. 3.
4.
LA decline geology – In May 1998, extraordinary ground support was necessary for the LA decline intake raise cutout and 45 metres down the decline due to contact with a structure carrying high grade molybdenum. Support consisted of two-metre long split bolts, two-metre long resin grouted rebar bolts, 100 mm x 100 mm x #4 Ga. mats and re-shotcreting. Rock conditions did not improve until the decline reached the 7065 level. Ground repair in Crusher top access – Faults and structures led to brow collapse as mining progressed into the Crusher top in September 1998. Clean up of the rock and three-metre or six-metre resin grouted hollow core bolts were installed prior to continuing the excavation of the Crusher top.
5.
Rock burst on the Crusher decline – In late October 1998, two rock bursts occurred in the Crusher decline just below the Crusher top. Again, Henderson and outside consultants developed a plan for repair and advance. Steps included clean up, re-bolt, install mats and re-shotcrete in order to advance the Crusher decline drift. The heading was shutdown for a portion of November 1998 for repair. Installing welded wire mats and waiting either 12 hours or 24 hours after blasting prior to entering each heading became the normal process. All drifting near the crusher was mined using caution and the drifts were not designed and mined until the geology was mapped. All of this led to delays in the mining schedule.
After this incident, a micro seismic monitoring system was installed in the areas currently developed. Measuring the release of stresses within the rock around the crusher area gave operations supervisors and miners data to estimate the potential of any other events occurring. Figure 12 is a photo showing the five-metre by five-metre Crusher decline nearly filled with rubble from the October rock burst.
Water handling Since 80 per cent of critical path mining was in the form of declines, dewatering was a critical element to the overall project. Several preliminary hydrogeological studies were done prior to and during the decline mining. Original estimates showed an average water flow of 150 to 200 l/m per heading, with a total average of 460 l/m. This analysis was based on only two core holes drilled from 7500 level to the Crusher. Based on these studies, mine de-watering wells were drilled ahead of the
FIG 9 - Rock burst in upper PC1 drift.
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PC1/PC2 TRANSFER STATION
FIG 10 - Structure in PC1/PC2 transfer station and upper PC1 drift.
FIG 11- Typical custom ground support needed for large excavations.
headings to reduce the amount of water encountered at the face. Flow meters were installed on the de-watering (weep) lines to monitor water flows as the mining advanced. But by July 1997, total flows had reached 1500 to 2000 l/m of 50 degree C water. Also, a tremendous amount of silt and fines were carried to the
MassMin 2000
face during mucking operations, reducing the pump capacities and wearing the pump impellers faster than expected. Pump performance was also poor due to electrical and pump re-build problems. The following water handling program was developed in response to these conditions:
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FIG 12 - Rock burst in crusher decline.
FIG 13 - Face jumbo in decline.
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Each heading was equipped with either a 12-kw or 22-kw submersible pump, pumping to a 22-kw or 45-kw pump no more than 100 metres away. This pump would feed a series of 45-kw booster pumps pumping to a large pumping station. A dedicated person was assigned to monitoring of the pumps for pressures and amps. A pump replacement and re-build program was instituted based on this monitoring. This person also checked for electrical problems and planned for pumping power to be available on a timely basis. Lease of a back-up diesel pump with sufficient head to replace a pump in case of failure. Purchase of an engineered recessed filter plate press to remove the majority of our higher than anticipated solids, reducing the possibility of main pump failures and decline mining flooding. On-going maintenance of the drift roads and ditches to minimise the amount of silt reaching the pumps.
Drifting for the LA Down decline was proposed to be more of a ‘straight path’ to the top of the crusher. This entailed changing the decline from a 7.0 per cent grade to a 12.5 per cent grade. This grade change eliminated approximately 210 metres of drifting on the critical path. In addition, 220 metres of drifting deemed not necessary to the project was deferred until a later year. Acquiring two articulated 36 tonne development haul trucks allowed the elimination of two planned muck loading stations for an additional reduction of 120 metres of drifting, again on the critical path. Also, the articulated haul trucks allowed the drifting to be reduced to five-metre wide by five-metre high, instead of 5.5-metre wide by 5.5-metre high. This reduction lowered the quantity of muck, and helped ease problems with muck handling. Also added to the design were pump cutouts for LA decline and for PC1 decline. Not only were these used for staging booster pumps to help de-water the face, but also were used to drill de-watering holes parallel to the drives for de-watering and structure mapping.
Acceleration of project schedule In August of 1997, the decision was made to accelerate the entire project by ten months. It was soon realised that we were not going to be able to accelerate the mining this much without changes to scope or resources. A new schedule was instituted in October 1997. This included a new decline design with less mining, a decrease in non-truck haul drift width, addition of new mining equipment, increase in staffing and an increase in electrical and maintenance support. Figure 13 is a photo of a new two-boom heading jumbo in a decline.
Large diameter raise boring Boring of a large (3.7 m) diameter intake raise feeding fresh air from the vent level to LA decline was difficult, and slowed down the mining in the critical LA decline. Blocky and hard rock was encountered while boring. This caused the reamer head to oscillate breaking two stems and two stabilizers. The reamer head dropped four times and the hole finished one month late. The effective delay to the decline mining was two weeks in February 1998. For future large diameter holes the raisebore contractor used 327 mm diameter drill rod rather than 286 mm, decreasing the chances of drill component failure. Also, for the
Fig 14 - End dump 36-tonne truck.
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remaining large diameter raises, a reamer more suited to our rock type was leased in order to speed up the reaming rate. A total of 1090 metres of raises were then bored with only minor problems.
Summary of difficulties impact
Difficult muck handling Eighty per cent of the decline mining was steep grade (from 7.0 per cent to 14.5 per cent). All muck generated by the mining and excavation was required to travel up these drifts to a single dump point. This resulted in traffic problems, double and triple handling of muck and longer cycle times than estimated. The original system used two six-tonne loaders, one articulated 27-tonne truck and two end dump 36-tonne trucks. The photo in Figure 14 show one of the original end dump 36-tonne trucks dumping. Muck handling was improved by adding one 15-tonne loader, two 12-tonne loaders and two highly mobile 36-tonne articulated trucks. All muck handling had to be scheduled with the highest priority.
SUMMARY OF IMPACT Table 1 lists the schedule and cost impact of each problem encountered. Originally, mining was forecasted to require 164 weeks and cost $197(US) per meter. When the project was accelerated ten months, the mining schedule was shortened a corresponding 42 weeks and an additional $45(US) per metre was added to help achieve this earlier finish. The difficulties added an additional 15 weeks and $34(US) per metre for a total of 137 weeks at a cost of $276(US) per metre.
CONCLUSIONS Despite the best studies and plans, challenges appear when undertaking a mining project of this magnitude. Owing to staffing problems, rock bursts and poor ground conditions, higher than anticipated water inflows, schedule changes, raise boring problems and trying muck handling, mining and excavations were completed 15 weeks behind an accelerated schedule. But, with careful monitoring of conditions and rapid response to these challenges, delays to the full project were non-existent. With all of these challenges to mining overcome, the 27 week improvement to the original schedule was critical to the overall project success. Mining costs represented 14 per cent of the project and with the early completion helped contribute to a six per cent savings on the entire project.
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TABLE 1 Effect of problems on schedule. Schedule in weeks
Cost per metre (US)
Original Estimate
164
$197
Acceleration of Project
-42
$45
Final Estimate
122
$242
Staffing and Training
5
$18
Rock burst and Water Handling
6
$10
Raisebore Delays
2
$5
Muck Handling
2
$1
Impact of . . .
Total Increase
15
$34
Actual
137
$276
REFERENCES Barfoot, G and Keskimaki, K W, 1998. The Henderson coarse ore conveying system, SME Annual Meeting, Orlando, Fl. Keskimaki, K W and Jensen, E B, 1992. Advances in equipment technology at the Henderson Mine, MassMin 92, Johannesburg, pp 345-350, SAIMM. Keskimaki, K W, 1996. Productivity gains at the Henderson Mine, MINExpo 96, Las Vegas, NV. Rech, W D and Lorig, L, 1992. Predictive numerical stress analysis of panel caving at the Henderson Mine, MassMin 92, Johannesburg, pp 55-62, SAIMM. Rech, W D, Jensen, E B, Hauk, G and Stewart, D, 1992. The application of geostatistical software to the management of panel caving operations, MassMin 92, Johannesburg, pp 275-281, SAIMM.
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The New Henderson Mine Truck Haulage System — The Last Step to a Totally Trackless Mine W D Tyler1, K W Keskimaki2 and D S Stewart3 ABSTRACT In 1999, the Henderson Mine commissioned the Henderson 2000 project - a new truck haulage level, a new underground crusher and a 24 kilometre conveyor system to replace a 23-year old rail haulage system. Henderson has been able to continuously improve in almost all aspects of its rubber-tired mining. The rail haulage system represented one of the last barriers to continuous improvement to the overall plant performance. Many of the lessons learned in other areas of the mine were utilised to design the truck level. This paper chronicles the benchmarking, equipment selection and design of the level. Areas covered will include haulage trucks, level layout, ore chute design, traffic control, road design and crusher layout.
INTRODUCTION The rail system in 1995 was the last step at Henderson to complete the changeover to a completely trackless mining system. The haulage system, although it had undergone several upgrades in its 23-year history, required a major renovation to enable it to complete hauling the reserves above the 7500 level. In addition, the haulage system would have to be modified to allow it to haul the ore located beneath the 7500 level. Declining performance and obsolescence of the ore transportation system at Henderson led to benchmarking and haulage alternative studies. The modernisation of the rail haulage system was estimated to have required an equivalent amount of capital to the final choice of the crush and convey system. The mining required to provide access to the lower level for the existing haulage system would have required a five year lead time.
THE RAIL HAULAGE SYSTEM During the past 20 years the ore has been transported from the underground operation to the mill by way of a 24-kilometre long rail haulage system. This system consisted of a 16-kilometre, three per cent upgrade tunnel, a 7.2-kilometre, one per cent grade surface segment and the 7500 gathering level underground at 2290 metres in elevation. This double track system in conjunction with a fleet of 32, 50-tonne locomotives and 220, 18-tonne ore cars was the only link between the mine and the mill, which are 100 kilometres apart by highway (Figure 1). The railroad system was extremely vulnerable to production interruptions, due to broken rail, derailed rolling stock, or train wrecks (Barfoot and Keskimaki, 1998). In addition to the aging rolling stock and track structure, the power distribution and control systems, as well as the trolley, also required upgrading. The other significant cost issues were elimination of PCB transformers, installation of a constant tension trolley on the surface (due to wide temperature variations), and a new control and data acquisition system for the six direct current rectifier stations in the system. 1.
Mine Technical Superintendent, Henderson Mine, PO Box 68, Empire CO 80438, USA.
2.
Technical Services Manager, Henderson Mine, PO Box 68, Empire CO 80438, USA.
3.
Mine Manager, Henderson Mine, PO Box 68, Empire CO 80438, USA.
MassMin 2000
FIG 1 - Henderson surface haulage system.
BENCHMARKING FOR A NEW SYSTEM Henderson saw many operations that have superior ore transportation systems as compared to the antiquated train haulage system previously utilised. In evaluating the Henderson 2000 project (Figure 2), the many potential improvements to ore transportation system included the following: 1.
centre-loading pneumatic chutes remotely controlled by the driver from the truck cab;
2.
large underground haulage trucks;
3.
truck gathering to move ore from the ore pass to the crusher over high quality roadways with excellent drainage;
4.
direct truck dumping to underground gyratory crushing; and
5.
ore conveyors.
Truck gathering system The ore transportation system begins with a truck gathering system on the 7065 level. A fleet of four side-dumping 72 tonne and two rear-dumping 36 tonne haul trucks is loaded using centre-loading, cab controlled, pneumatically powered chutes. Figure 3 and 4 shows the end section of the trucks in the haulage drifts. The rubber-tired equipment allows much more flexibility in mine layout than the track haulage equipment used previously (Gallagher and Teuscher, 1998). The haul trucks operate on a relatively flat grade with an average tram of 900 metres to dump at a centrally located crusher. The side-dump trucks are not capable of ramp operations as only two of the five axles are driving axles and they are equipped with 325 kilowatt engines. However, individual 72 tonne haulage trucks are capable of hauling up to 1000 tonnes per hour with cycle times as low as four minutes. Figure 5 shows the 7065 Haulage Loop layout in plan. The rear dump 36 tonne trucks were chosen to assist with haulage of muck from development loading stations during the mining for the new lower level. In addition, they will assist when needed in haulage from the production ore passes.
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FIG 2 - Henderson 2000 ore delivery system.
WHEEL LOAD = 13,600 k
EMPTY
787mm
3302mm
EMPTY @ TOP OF CAB
3581mm
4877mm
3906mm
5486mm
FIG 3 - Rear-dump truck and drift cross-section.
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305mm
3775mm
4868mm
794mm
3899mm 5486mm
FIG 4 - Side-dump truck and drift cross-section.
33 CH
UTE 34 CHUTE
ACCESS RAMP
44 CHUTE
TRUCK SHOP
CRUSHER
66 CHUTE
FIG 5 - Haulage loop layout.
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FIG 6 - 72-tonne truck dumping into crusher.
include use of bi-directional radial tires that are not limited in rotation direction. This will allow tire rotation so that the other side of the tire can be worn. Finally, door and box location sensors on the box are being moved to less vulnerable locations.
CRUSHING AND CONVEYING The original crushing system was located at the end of the haulage track system. As part of the upgrade, a used 54 × 84 Fuller-Traylor crusher was reconditioned. This crushing unit has a throughput of 2300 tonnes per hour. The crushed material is fed onto a 1.5 metre wide by 25 metre long Accelerator belt. The intent of the Accelerator belt is to increase the speed of the highly abrasive ore before it is loaded onto the main conveyors and to act as a sacrificial belt in the event that steel should cause a longitudinal belt cut at the transfer. Another important feature
R=
12 m
2438
6098mm
Rigid frame haulage trucks were chosen for several reasons after benchmarking articulated trucks and tractor/trailer combinations used in the underground mining environment. The articulated trucks typically found in underground environments were limited in capacity and had shorter lives. At the time of placing the order for the 72 tonne trucks, the largest articulated truck had a capacity of 60 tonnes or 17 per cent less. In open pit environments, rigid frame trucks have achieved frame lives in excess of 75 000 hours. Henderson believes that the rigid frames should have a useful life five to ten times longer than an articulated truck. Tractor/trailer combinations have greater capacities than the 72 tonne truck in a similar cross-section. However, the rigid frame truck has a shorter turning radius. Benchmarking showed that trailers have problems with tire wear, and kingpin wear. The rigid frame is capable of faster dump cycles, since no stabilising jacks are required. The 72 tonne truck is also easier to back up into chutes. Two of the four chutes in operation today at Henderson are back-in style chutes. The side-dump truck was also chosen because the dump cycle is about 15 per cent quicker than that of a rear-dump truck. The trucks are easier to spot at the crusher dump pocket using a laser beam to align the trucks (Figure 6). Finally, the trucks are less risky because backing has been minimised. Backing typically results in damage to the truck or the installation. Figure 7 shows a back-in style chute and a typical intersection. The only issues that have been experienced at Henderson with the haul trucks include minor spillage, tire wear and sensor placements on the truck body boxes. The spillage issue relates to the two-part box and the gap between the doors. Most of the spillage problem has been alleviated using rubber skirting on the crusher dump pocket lip to guide spillage into the crusher. Tire wear was anticipated to be a substantial part of the operating cost of the trucks. Options being examined to maximise tire life
6098mm
FIG 7 - Back-in style chute layout.
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of the belt is the 1.5 metre width and slower speed relative to the rest of the system. This facilitates the removal of steel via two magnet belts, one mounted longitudinally and the other mounted transversally. The ore is then transferred onto a three flight conveying system. PC1 is a 1200 metre long conveyor that runs from below the crusher to a transfer station at the end of the PC2 belt. PC2 is the longest conveyor in the system at 16.8 km and passes through the haulage tunnel. The ore is finally transferred onto the PC3 conveyor. PC3 is an overland conveyor of 6.7 km in length. PC2 and PC3 are both 1.2 metres wide.
ANCILLARY SYSTEMS Haul road construction and maintenance Good haul roads are key to creating and maintaining an efficient haulage system. Concrete roadways are used extensively in the production areas to facilitate LHD clean-up of spillage. However, concrete roadways are expensive to install and damaged sections are difficult to remove. Given the past experience, the decision was made to construct roads initially using a nominal 300 millimetre thick, roller-compacted roadbase material. This decision was justified where drainage was adequate and rain shield installed to prevent water puddling on the roadways. Some intersections were installed with concrete roadways to prevent erosion due to the wear of loaded trucks turning. Since the initial construction, a few sections were rebuilt, using geo-cellular polyethylene membrane filled with roadbase. This has prevented the formation of washboard and improved the speed of travel of the trucks. A Caterpillar 120G motor grader and a vibratory compactor are used to repair damaged sections. Roadways at the chutes themselves are concreted to assist in cleanup and provide a consistent chute-to-truck loading distance.
High compressive strength concrete (34.5 mPa) is used in the roadway and is poured using laser leveling to ensure a consistent loading surface (Figure 8).
Traffic control system The traffic control system is designed to provide safe movement of vehicles on the 7065 haulage level. This tool is needed to allow Henderson to maximise the production from the haulage level and accommodate the volume of traffic required to continue the development of the shop areas and haul roads as well as traffic related to the maintenance and clean up of the chute areas. Pedestrian traffic is not accommodated or allowed during haulage operations. The development of the 7065 haulage level will be an on-going project for several years. One-way traffic on parts of the haulage loop will not be established until some time in the future. Haulage vehicles need to produce between 18 000 and 36 000 tonnes per day. This is the equivalent of 250 to 500 haul truck cycles per day. This has a truck arriving at the crusher every two to three minutes during peak production periods. The system is also designed to accommodate changing traffic patterns as more of the haulage level is developed over the next 15 years. Traffic on one haulage segment has to handle two-way traffic on a roadway only wide enough to handle one-way traffic. This was handled using a timed system, like the traffic light-controlled intersections in municipalities. The traffic control equipment selected comes from suppliers for municipal traffic control. The system is designed using off-the-shelf components that are readily available. A local traffic engineer and supplier were selected to help design and configure the system. Instrumentation staff at the mine performed the actual installation and programmed the software that controls the system. 243
9m
m
ORE PASS EXHAUST RAISE
121
8m
m
508mm
5486mm
9449mm
CENTER LOADING CHUTE
6096mm
15559mm
4874mm
COMPACTED ROAD BASE ON GEOTEXTILE COVERED DRAIN LAYER HAULAGE DRIFT
CENTRE LOADING CHUTE
FIG 8 - Centre loading chute layout.
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The alternative was to design and build a Programmable Logic Controller (PLC)-based system. That system, while infinitely programmable, was deemed less desirable due to the dependence on an outside programmer. The equipment associated with PLC systems is also more sensitive to variations in the ambient temperature and humidity. The total cost of the traffic control system was just under $100 000. The total amount included all equipment and a spare controller that can be used for training. The spare controller will eventually control intersection lights as the level is developed. The cost also includes Henderson costs for installing the cabling and sensors.
Chute monitoring A laser distance measuring system is used to indicate the minimum or maximum level of ore in an ore pass or bin. The lasers have the ability to operate in dusty or humid environments without producing erroneous measurements. The lasers are connected to the mine PLC system to convert readings into meaningful information for the haulage dispatcher and to eliminate false readings caused by dust created when ore is dumped down a partially empty ore pass (Figure 9). The lasers do not measure distance down the ore pass, but confirm the presence or absence of muck at a given level of the ore pass. The laser is aimed horizontally across an ore pass intercept and allows the indication of ore in the intercept. The primary reason for installing the laser was to prevent the chutes from being pulled empty. The lasers have been generally successful, with the exception of PLC errors and deadbed material blocking the laser beam.
FIG 9 - Ore pass laser installation.
The communications to the PLC system on the network is tested every 20 seconds. Error detection had to be installed to ensure that the readings were valid and recent. If the distance doesn’t update on the PLC data highway, the value will not change. The PLC communications protocol used is a 4-20 milliamp current loop. The PLC was programmed to display the interpretation of the bin level using the laser distance indication and some logic to mask confusing readings. For example, the laser is not impervious to dust, but dust created by dropping muck down a
FIG 10 - Typical ore pass intersection layout showing laser.
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partially empty ore pass will dissipate within 50 to 60 seconds. No interpretations are changed during that time. The ‘ringing’ of distances in dust is actually used as an indication by the haulage dispatchers that the ore pass level is at the intercept on one of the legs (Nelson and Harney, 1995). The build-up of deadbed was not anticipated, but was caused by the angle of repose being flattened out by a large amount of water coming down some of the ore passes. The last ore bin constructed has the bulkhead at the edge of a 65° slope into the bin to aid in clean out of the deadbed. In addition, the bin has a timbered window that could allow a small LHD to be used to remove the deadbed. The laser is mounted in a cast opening of the intercept bulkhead. The laser can shoot distances over the top of the bin. When the minimum distance is registered, the bin is full. If the maximum distance is registered, the bin will be near empty. Tests are on going for the use of a weighing system on the chute assembly. A non-linear response is seen to the amount of material added to the ore pass and bin, however a setpoint should be able to be established as a minimum amount of ore in the system. This will aid in using the full capacity of the storage bin, without pulling the chute empty.
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CONCLUSION The replacement of the rail system with a productive truck haulage fleet will allow Henderson to exploit the life-of-mine ore reserves economically. The trucks were chosen for their productivity, to provide long life and accommodate both pull-through and back-in chutes. Good haul roads are important in keeping the trucks productive and in minimising maintenance costs. Traffic control should minimise queuing delays in the haulage system. Instrumentation is a key element in keeping the chutes from being pulled empty. Alternative instrumentation, such as weighing systems is being tested to evaluate their effectiveness in providing full use of an expensive ore pass complex.
REFERENCES Barfoot, G and Keskimaki, K W, 1998. The Henderson Coarse Ore Conveying System, SME Annual Meeting. Orlando, FL. Gallagher, J and Teuscher, J S, 1998. Henderson Mine: Preparing for the Future, SME Annual Meeting. Orlando, FL. Nelson, B V and Harney, D C, 1995. Remote Control and Monitoring at the Henderson Mine, SME Annual Meeting. Denver, CO.
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Application of Block Caving System in the Tongkuangyu Copper Mine Zhou Aimin1 and Song Yongxue2 ABSTRACT During the design and operation stages of a block caving project, a large amount of research and development work has been undertaken in the Tongkuangyu Copper Mine. This paper presents the research achievements from these studies, including a rock mass quality evaluation method, cavability prediction, fragmentation evaluation, caving mechanisms, a computerised production scheduling model and the optimised draw control system. The main features of the block caving system in the Tongkuangyu Copper Mine are also discussed.
INTRODUCTION The Tongkuangyu Copper Mine, located in Yuanqu County, Shanxi Province, China, is a leading operating mine of Zhongtiaoshan Nonferrous Metals Corporation of CCLZ (China Copper Lead Zinc Corporation). The mine was designed for a full capability of 4 Mt/y and it was commissioned briefly and hastily in 1974 using a pillared sublevel caving system. An annual production of 600 000 - 800 000 t was actually achieved prior to the late-1980s. The inability to operate at the designed full capacity was responsible for long-term loss of the mine. In the late-1980s, a technological innovation program on the mining system was launched and the block caving system was applied, which turned the mine operation from financial loss into profit-making. In 1985 - 1986, a Sino-American joint design group accomplished preliminary and detailed designs for block caving of the No 5 orebody. Capital constructions concerning the technological innovation program were accomplished in 1987 1989 and trial production started in 1991, and full production was reached in 1999. The mine is currently the biggest underground metallic mine in China. In order for the program to achieve its expected targets, during the periods of project design, capital construction and trial production, experts from Changsha Institute of Mining Research, Beijing General Research Institute of Mining and Metallurgy, Beijing General Institute of Nonferrous Metals Design and Research, Central South University of Technology and Beijing University of Science and Technology performed a series of experiments and studies on the block caving system. Research achievements have been adopted in the mine design and production, and remarkable application results have been achieved.
MINING TECHNOOGY CONDITIONS Orebody occurrence The Tongkuangyu mineralised zone is located in the middle-to-top of the Tongkuangyu metavolcanic group, part of the Lower Proterozoic Jiang County Group. The outcropped metavolcanic group, descending from old to new, consists of metakali-rich rhyolite, metakalic basic volcanics and 1.
Changsha Institute of Mining Research, 236 Lushan South Road, Changsha, Hunan Province, PR China. E-mail: [email protected]
2.
Zhongtiaoshan Nonferrous Metals Corp, Shanxi, China.
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metatuffaceous semi-pelite. The Tongkuangyu deposit is a complex copper deposit having undergone multiple geologic actions and having multiple geneses. It is embedded in meta-semi-pelite and consists of 113 orebodies, seven of which are suitable (big enough) for commercial mining, and the No 5 orebody is the largest, followed by the No 4 orebody. The No 5 orebody is the major mineralisation at the Tongkuangyu Copper Mine and it accounts for 70 per cent of the total ore reserve. This orebody consists mainly of meta-quartz crystal tuff, meta-quartz porphyry and meta-beschtauite. Hangingwall host rock is sericitic quartzite while the footwall host rock consists of sericitic quartzite and chlorite quartz schist. The orebody is a stratoid lenticular shape and its occurrence is parallel to the host rocks with a dipping angle of 30 - 50o to the north-west. The strike length of the orebody is 980 m and its extension is more than 1000 m. The average thickness is approximately 110 m and the maximum thickness 200 m. The orebody has a geological copper grade of 0.67 per cent, with a small amount of copper contained in host rocks and intercalated rocks.
Structural features of rock mass A well-developed fault structure is found in the mine area and most of the faults were formed in the late stage of regional metamorphism or afterwards. Active movements in the long process, however, have caused two principal faults and a number of developed secondary faults. There are two predominant sets of joints, almost orthogonal to each other. The first set has an average inclination of 320o and a dip angle of 59o and it accounts for 40 per cent of the overall joints. The second set has an average inclination of 135o and a dip angle of 50o and it accounts for 38.5 per cent Random joints account for 21.5 per cent. RQD of the ore-bearing rock mass is an averaged of 73 per cent. Joint spacing is an average of 0.42 m in the first set, 0.73 m in the second and 0.47 m for random joints. Closed joints are dominant, accounting for 77.5 per cent. Surfaces of these joints are closely contacted, almost containing no infill materials. When subjected to stress actions, they can easily be opened. Open joints account for about 21.4 per cent while re-closed joints are very rare, accounting for only 1.3 per cent. Surfaces of joints are continuous and joints with a persistence of more than 2 m account for 49.2 per cent. More than 84 per cent of joint surfaces are rough planar.
In situ stresses and rock properties The mine area is a medium-to-low stress zone, where tectonic stress is predominant. The maximum principal stress in the area is 10 - 14 MPa at 810 m level with an inclination of 60o - 80o and a dip angle of 10o. The ratio of horizontal stress to vertical stress is 1.0 - 2.5. Physical and mechanical properties of the ore-bearing rock mass are shown in Table 1. The uniaxial compressive strength (UCS) tests indicated that most of the specimens failed violently, a sign of brittle failure, while some of them exhibited shearing failure. In tensile strength tests, however, most of the specimens were cleft along their central axis and a few of them were broken along micro-fractures.
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TABLE 1 Physical and mechanical properties of ore-bearing rocks (standard 50 mm specimens). Rock type
UCS MPa
Tensile strength MPa
Young Modules GPa
Poisson’s Ratio
Bulk density tonne/m3
Metaquartz crystal tuff
60 - 130
4 - 13
54.4
0.244
2.720
Metaquartz porphyry
120 - 150
6 - 16
metabeschtauite
120 - 160
Metamorphic basic intrusions
60 - 100
Chlorite quartz schist
80 - 130
Sericitic quartzite
90 - 150
Sericitic quartz schist
100 - 155
Diabase
150 - 220
70.8
0.230
2.848
62.9
0.257
2.687
2 - 10
50.8
0.295
2.987
4-9
52.4
0.280
2.884
5 - 11
58.3
0.215
2.775
50.5
0.270
2.742
73.5
0.260
2.900
CAVABILITY AND FRAGMENTATION Block caving is usually applied to mining of highly fractured orebodies. Once a sufficient area at the bottom of a block is undercut caving starts naturally under gravity and the influence of induced stresses in the cave back. The success of a block caving mine is considered to be greatly dependent on two key factors: cavability and fragmentation of the rock masses. Because of limited expertise available in China, necessary research work such as orebody cavability, fragmentation and caving mechanisms was carried out in the Tongkuangyu Copper Mine by the research institutions mentioned previously.
Rock cavability The orebodies in the Tongkuangyu Copper Mine consist of hard rocks with joints. The cavability could be variable from place to place in the mine. Predicting the cavability is important in the mine design and production scheduling. Owing to variable rock mass conditions in the mine, a method was developed to evaluate rock mass quality. Four parameters were used for the evaluation of rock mass quality and they are aggregative joint spacing, RQD indices, aggregative friction angle and equivalent joint set number. According to the rock mass quality and orebody geometry, the cavability at the No 5 orebody can then be assessed.
Data collection To obtain geological and geotechnical information, more than 5000 m of development was mapped and more than 7800 m of drill holes were logged. Laboratory tests on rock specimens were also undertaken. Relevant properties on the ore-bearing rock mass were obtained from mapping, logging and testing. These include number of joint sets, joint spacing, occurrence, continuity, openness, properties of filling materials, roughness, waviness, intensity of weathering, UCS, C, φ, rock fracture’s subcritical propagation rate, rheological behaviour of rocks and mechanical properties of joint planes.
Data processing The Mukhrejee and Mahtab’s method (Mukhrejee and Mahtab, 1987) was used to determine comprehensive joint parameters. The empirical formula, φ=arctan(18.76+1.28i) (where i is the root mean square slope of joint planes), was used to calculate the aggregative joint friction angles (Pan et al, 1994). RQD and the orientation of major joint sets obtained from mapping and logging were also taken into considered. All data were then analysed and the evaluation indices were determined for all four items, ie aggregative joint spacing, RQD value, aggregative friction angle and equivalent joint set number.
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A sample database for evaluation indices Based on the obtained evaluation indices, the three-dimensional Kriging block method was used to deduce evaluation indices for the location where rock mass data was not available so that a database could be built for the entire orebody.
A model for rock mass quality classification The weighting of various evaluation indices was assessed using the uniformity functional method. A standard mode for rock mass quality classification was then formulated through clustering analysis.
Cavability prediction Through mode identification, the rock mass in the area of interest, defined by No 2 and No 13 exploration lines horizontally and between the 800 m and 900 m levels vertically in the Tongkuangyu No 5 orebody, was classified in Table 2. TABLE 2 Rock mass classification in the No 5 orebody. Classification
Rock mass quality
Percentage
Type I
Very good
0%
Type II
Good
32.38%
Type III
Fair
58.33%
Type IV
Poor
9.29%
Type V
Very poor
0%
The corresponding RQD values for Type II, Type III and Type IV are 83.86 per cent, 58.73 per cent and 26.38 per cent respectively (Tan, 1997). It can been seen that the combined Types III and IV rock mass in the Tongkuangyu No 5 orebody account for over 60 per cent of the total volume, indicating a fair-to-good cavability.
Fragmentation prediction Fragmentation prediction is important in a block caving mine. It has significant influence on production cost and scheduling. The fragmentation is usually classified as primary and secondary fragmentation. Primary fragmentation of a deposit is mainly relevant to weak planes such as joints. Orientation, spacing and persistence of joints are considered when the primary fragmentation is assessed.
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APPLICATION OF BLOCK CAVING SYSTEM IN THE TONGKUANGYU COPPER MINE
However, the secondary fragmentation is the block size of ores drawn out from the draw-points. The induced stresses, due to undercutting and arching, and the crushing effect between rock blocks are taken into account by the secondary fragmentation assessment. Therefore, the secondary fragmentation heavily depends on induced stresses, the draw column height, rock strength, shape of in situ blocks and joint conditions. The Monte Carlo simulation technology and the three-dimensional block size model are utilised for the prediction of both primary and secondary fragmentation in the Tongkuangyu Mine. The prediction indicates a fine fragmentation in the No 5 orebody (Wang et al, 1998). When production began at the drawpoints, the drawpoints fragmentation was monitored regularly. A size distribution of ore blocks at drawpoints was produced. A comparison between the predicted and monitored fragmentation is shown in Figure 1. It can be seen that the monitored drawpoints fragmentation agrees well with the predicted secondary fragmentation. The oversized (more than 1.2 m) ore is six per cent of the total ore mined, compared to the nine per cent of oversize predicted from the fragmentation model.
Caving mechanism Based on the mining conditions at the Tongkuangyu No 5 orebody, research has been undertaken to understand the rules governing the natural block caving process. The research work involves physical model simulation, numerical stress analysis using three-dimensional finite element method and field monitoring of caving propagation. In the physical model tests two 3.0 x 2.0 x 1.45 m (3D) models and two 2.5 x 0.2 x 1.35 m (planar) models were built to investigate the caving mechanisms. Caving propagation monitoring was carried out through drill holes using a cirrus-type detector with open circuit cables (Zhang, 1997).
It is considered that for such a well-jointed hard rock orebody, the natural caving process in the Tongkuangyu Copper Mine can be described below:
• caving of the orebody largely depends on the distribution of joints in the rock mass;
• time for micro-fractures (healed joints) to transit from the subcritical propagation to the high-speed propagation is a function of stress exerted on the micro-fractures; and
• owing to the influence of variable micro-fracture propagation in the rock mass caving occurs periodically. The study also reveals that the natural block caving process can be divided into two stages, ie undercut caving and continuous caving.
Undercut caving Prior to undercutting, the orebody is in its virgin stress status. As the undercut area expands, the virgin stress status is disturbed. When the induced stresses is sufficiently high the orebody starts caving. However, the unstable area is very limited at this stage. When a stable arch or a critical stable state is reached caving will stop. Increasing the undercut area can initiate caving again. The main feature of this stage is that, whenever the undercut area increases, caving will take place. As a feature of undercut caving, the undercut area is important in consideration of capital construction planning and production scheduling. For the Tongkuangyu No 5 orebody, the area for undercut caving to occur is typically 3200 – 5100 m2.
Continuous caving Continuous caving occurs when the undercut area (including the weakened area) increases to such an extent that caving is not dictated by undercutting. At this time, the caving process will no longer depend on the increase in the undercut area. Once there is
FRAGMENTATION
Accumulative %
100.0
Predicted Secondary Fragmentation Predicted Primary Fragmentation Monitored Fragmentation at Drawpoints
80.0 60.0 40.0 20.0 0.0 0.0
0.5
1.0
1.5
2
2.5
Block Size, m FIG 1 - The ore fragmentation at the Tongkuangyu No 5 orebody.
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an open space or air gap created by drawing caving can be maintained, ie caving is controlled by drawing from the extraction levels until caving breaches to the surface. The characteristic parameter indicating the speed of caving is called the continuous caving rate. The continuous caving rate in the Tongkuangyu Copper Mine is approximately 0.375 m per day.
DRAW CONTROL In the Tongkuangyu Copper Mine, the ore removal system consists of scrapers and chutes. About 200 drawpoints are in operation during a normal production period. In order to maintain a steady production and minimise ore loss and dilution, the unequal-homogeneous-drawing principle is used for production scheduling. A command-based draw control system has also been put into use. In this system draw schedules and commands are formulated and programmed in a computer.
Production scheduling Production scheduling involves short-term plans (12 months) and medium- to long-term plans (two - five years). For the short-term plans, draw tonnage is projected on a monthly basis, and for the medium- to long-term plans, draw tonnage is projected on a quarterly basis. Production scheduling is made according to the scheduling model described below (Zhou, 1997): n α ⋅ Qd i − Qpi > 1000 m if Q = ∑ η(α i • Qd i − Qpi ) η = 7 1 if α ⋅ Qd i − Qd p < 1000 I α ≥ 1 i =1
Qf i =
α i • Qd i − Qpi n
∑ (α
j
•Q
i = 1,2, ......n
(1)
(2)
• Qd j − Qp j )
j =1
≤ 30 ≤5 ≤ 50 ≤ 10 ≤ 60 Qpi ≤ 20 s. t. Qfi if ≤ 70 Qdi ≤ 30 ≤ 80 ≤ 40 = 100 > 50
Final control over the drawing process is realised through control commands which, after being optimised on the basis of the 12-month plans, are issued on a daily basis. When determining the optimal drawn-out tonnage of a certain day, the actual daily drawn-out tonnage and the actual working conditions of the day are taken into account. In the Tongkuangyu Mine, based on the ore properties, a series of physical simulation tests were carried out, including a three-dimensional 1:50 single-chute drawing model, a three-dimensional 1:100 multiple-chute drawing model. A computerised simulation drawing test was also undertaken. According to the test results, the drawing process beneath the overburdens was optimised using ‘the least ore-waste contact area’ principle. This principle is based on the fact that when drawing is undertaken under overburdens, ore loss and dilution happens primarily at the ore-waste interface. Therefore, the smaller the contact area, the less the ore loss and dilution. A correlation between the contact area and draw-out tonnage was found through the tests and simulation results in the mine. The optimised draw scheduling was then implemented in the production and, the ore loss and dilution rates were reduced to a minimum. The research results indicate that when the draw column is higher the drawn-out body is such an ellipsoid shape that the upper part of the ellipsoid is bigger and lower part is smaller. This ellipsoid is formed from a butted joint of two ellipsoids with an equal semi-minor axis. Based on this assumption, a formula describing the movements of rock particles in a draw column can be deduced (Zhou, 1998). This formula can be used to describe the movement of any particles in the draw-out body. Therefore, with the aid of numeric methods, a functional relationship is established between the drawn-out tonnage and the ore-waste contact area. Furthermore, ‘the least ore-waste contact area’ can be defined as the target function of the variables for the drawn-out tonnage.
min S = f (x (Qf), y (Qf), z (Qf)) s.t. Qfj = Cj j = 1,2 ...... T< n n
∑ Qf
i
(3)
=Q
i=1
Where, Q~total production in the scheduled period; Qfi~planned drawn-out tonnage in the scheduled period; Qdi~ Drawn-out tonnage; Qpi~drawn-out tonnage; αi~scheduling index of drawpoint, the value of which is similar to that adopted at the Henderson Mine (Dewolfe, 1982), ie αi of the drawpoint immediately above the undercut line is set to zero and starting from the second row, the index ascends at an increment of ten per cent each time until it reaches 100 per cent and thereafter, the value of αi is always set to 1; η~stope production scheduling coefficient; n~number of drawpoints in a stope; m~number of months in a scheduling unit, ie three for medium- to long-term plans and one for yearly plans. For planning medium- to long-term production, ore tonnage and grade are determined according to Laubscher’s grade assessment (Laubscher, 1994). The drawing ellipsoid above a drawpoint is considered to mix with waste rocks, ie dilution entry, during ore drawing. Assuming that all ore above a drawpoint is drawn-out completely, the total quantity of ore/material drawn-out is linearly proportional to the amount of dilution entry due to the mixing between ore and waste.
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Optimisation of drawing control
Where S - overall ore-waste contact area; x(Qf), y(Qf) and z(Qf) spatial status of any rock particle; Qf - tonnage drawn-out at the drawpoint; Q - overall drawn-out tonnage in a panel; and Cj tonnage at a drawpoint where special treatment is required. In the Tongkuangyu Mine, excellent results in ore drawing have been achieved since the draw is controlled through an optimised draw control system. Based on the ore removal records from the 63 scraper drifts at the end of 1999, the overall ore recovery now is up to 97.1 per cent and the ore dilution rates are cut to as low as 8.6 per cent.
LAYOUT OF MINING OPERATIONS The first lift of block caving at the Tongkuangyu Mine is from the 810 m level at the No 5 orebody. The extraction level is developed along the strike of the orebody. Caving production has been continuous along the orebody strike. The main extraction level is located at 810 m level. Owing to the moderately dipping orebody three auxiliary extraction levels are necessary for ore removal at the 833 m, 853 m and 870 m levels. At the 870 m level an observation drift is also developed. A plan view and two section views of the Tongkuangyu block caving are shown in Figure 2.
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APPLICATION OF BLOCK CAVING SYSTEM IN THE TONGKUANGYU COPPER MINE
FIG 2 - A sketch of block caving system at the No 5 orebody in the Tongkuangyu Copper Mine.
A bottom structure is designed for ore removal using a scraper-chute system. A locomotive-based haulage system is deployed at the 810 m haulage level and immediately above the 810 m level is the scraper level for ore removal (Figure 3).
The 810 m level haulage system The 810 m level haulage system offers circular hauling. The overall track length is more than 2200 m and the length of each cross-cut haulage road is 350 - 500 m. Trains are hauled by overhead-wire electric locomotives. There are 15 - 20 trains for transporting ore, waste rocks, concrete, explosives, materials and services. When so many trains are travelling underground, traffic could be very heavy at the haulage level. Ore trains have to be directly
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under chute holes for ore loading. So they must be well co-ordinated with the drawing process. Passage preference is required and reverse movement is allowed for concrete tankers and explosive trains. Therefore, dispatching of the 810 m level haulage system becomes very complicated. To tackle this problem, locomotives are automatically dispatched through a computerised control system. This system can automatically track and monitor the movement of each train while the automatic control semaphore and the electric switch stand will guide trains to their pre-set destinations. The PLC of the system, placed in the haulage dispatch room at the 810 m level, is connected via a local I/O station to a simulated haulage console. The console cannot only dynamically simulate the field situation but also make it possible for the haulage dispatcher to conveniently and flexibly guide the movements of haulage
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A
4
2 6300
1 5000
7
3 10000 39050
6500
6
10000
9 3
5000 2750
5000
70 13
6336 3000
2585
5
8 305
3000
3000
457
8
1830
A£ A
A
1. Cross-cut haulage drift; 2. Service raise; 3. Scraper drift; 4. Scraper chamber; 5. Chute; 6. Return air vent; 7. Cross-cut exhaust drift; 8. Undercut drift; 9. Finger-like raise FIG 3 - Structural parameters of the ore drawing system at the No 5 orebody.
vehicles. The system is equipped with an excellent self-diagnostic capability. Prior to starting it can identify whether the system or the control equipment is under a good condition through CPU scanning.
Scraper-chute ore removal system The scrapers system that was preferred over the LHD units was purely due to the capital budget limitation at the time of planning. At the 813.5 m level, the scraper drifts are developed along the strike of the orebody. In each scraper drift, six ore removal chutes are constructed face to face symmetrically. Each chute is responsible for an ore removal area of 10 x 10 m. As the orebody has a dip angle of about 45, auxiliary level scraper drifts are constructed at the 833 m, 853 m and 870 m levels. The scraper drifts and chute mouths are consolidated with 300 mm high strength concrete.
Undercutting and weakening The major undercut engineering work involves the development of undercut drifts. The major weakening engineering work is the 50 m high boundary slot at the end of the orebody and the weakening raises at the corners of each panel. A diagonal undercutting sequence is from the footwall to the hangingwall and advances in benches along the orebody strike. The main reasons for this arrangement is that the undercut drifts and scraper drifts are deployed along the orebody strike and that most of the minor faults have a dip direction along the strike of the orebody. These closely spaced faults could assist in initiating the caving. Each undercut blast is carried out to form a rectangle shape in plan, with the longitudinal direction of each undercut area perpendicular to the orebody strike. Each undercut blast ring usually covers an area of 3.5 x 20 m. In each undercut drift, no more than three rows of blast holes should be detonated in a single undercut blast. However, blast can be undertaken simultaneously in up to three undercut drifts. With this undercutting sequence, large-scale failure along a fault can be prevented and risk of having coarse fragmentation due to the
330
fault failure reduced. It is also easy to check undercut quality and ensure no pillar is left. In addition, due to frequent change of the undercut front where the stress concentration occurs, the duration for suffering higher stressing on a given area is reduced. Therefore, damages due to this stress concentration can be mitigated at the undercut and ore extraction levels.
THE FUTURE Mining at the 810 level of the No 5 orebody has been completed. Production at the 690 m level in both the No 4 and No 5 orebodies has been commissioned. The scraper system is still being used for ore removal at the 690 m level in the No 5 orebody while 3.0 m3 electric LHD units are being used for the 690 m level in the No 4 orebody. The production at the 690 m level (including No 4 and No 5 orebodies) is maintained at 4 Mt per annum. The mine design is currently being carried out for the level below the 690 m level. The 6.0 m3 electric LHD units will be used for ore removal and conveyor belts for ore transportation. The production is expected to reach 6.0 Mt per annum in 2005 in the Tongkuangyu Copper Mine when 6.0 m3 LHD units are used for the lower levels.
CONCLUSIONS The research program on the block caving system in the Tongkuangyu Copper Mine was supported and funded by the Chinese Government and was designated as one of the key science and technology research projects in the Nation’s Seventh Five Year Plan period. Experts from several research institutions and universities were involved in various research tasks. The research achievements have been applied in the mine’s technical renovation designs and production operations. Further research work is being carried out to continuously improve the mining operations at the mine. With the introduction of LHD units from the levels below the 690 m level (inclusive), the underground working environment will be improved significantly. The higher output of 6.0 Mt per annum from the current 4.0 Mt per annum will reduce the total costs and therefore put the company in a
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APPLICATION OF BLOCK CAVING SYSTEM IN THE TONGKUANGYU COPPER MINE
better financial position. It is expected that application of block caving system will play a more important role in improving mine production in the Tongkuangyu Copper Mine.
REFERENCES Dewolfe, V, 1982. Draw control in principle and practice at Henderson Mine, Design and Operation of Caving and Sublevel Stoping Mines. Laubscher, D H, 1994. Caving mining—the state of the art, Journal of the South African Institute of Mining and Metallurgy, 10:279-293. Mukherjee, A and Mahtab, A, 1987. Size distribution of ore fragments in block caving, the 13th World Mining Congress Thesis, Stockholm Sweden. Pan, C, Li L, Wang, L, Wang, W, and Tan, L, 1994. Rock caving characteristics and caving rules in block caving, Journal of the Central South Institute of Mining and Metallurgy, 25(4):441-445.
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Tan, Guangwei, 1997. A study on the block caving rules of the No 5 Orebody in the Tongkuangyu Copper Mine, Nonferrous Metals, 5:8-12. Wang, L, Pan, C and Tan, G, 1998. Simulation-based ore fragments model and its applications, Nonferrous Metals, 2:6-10. Zhang, Feng, 1997. Monitoring the orebody caving state in a block caving mine, Ferrous Mines, 9:9-12. Zhou, Aimin, 1997. Optimisation and method of drawing control in block caving at the Tongkuangyu Mine, Trans Nonferrous Me Soc China, 17(2):9-13. Zhou, A and Wang, D, 1989. A study and application of ore drawing rules in high level caving system, the Quarterly of Changsha Institute of Mining Research, 9(2):20-27.
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Block Cave Design and Geotechnical Considerations Design of the Second Block Cave at Northparkes E26 Mine
S Duffield
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Modelling and Design of Block Caving at Bingham Canyon
C J Carter and F M Russell
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Considerations for Design of Production Level Drawpoint Layouts for a Deep Block Cave
A R Leach, K Naidoo and P Bartlett
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Noranda’s Approach to Evaluating a Competent Deposit for Caving
S Nickson, A Coulson and J Hussey
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The Past Focuses Support for the Future
A D Wilson
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Rock Mechanics as Applied in Philex Block Cave Operations
R S Dolipas
395
Block Cave Undercutting — Aims, Strategies, Methods and Management
R J Butcher
405
The MRMR Rock Mass Rating Classification System in Mining Practice
J Jakubec and D H Laubscher
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The Role of Mass Concrete in Soft Rock Block Cave Mines
R J Butcher
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Meeting Geotechnical Challenges — A Key to Success for Block Caving Mines
Dianmin Chen
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Block Caving — Controllable Risks and Fatal Flaws
T G Heslop
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Design of the Second Block Cave at Northparkes E26 Mine S Duffield1 ABSTRACT The pressure of low commodity prices forced the design process for the second lift of Northparkes E26 Mine to adopt a continuous improvement philosophy. The emphasis was to minimise capital expenditure while still maintaining the excellent low operating costs achieved on Lift 1. Although the first block cave lift at Northparkes is the most productive underground mine in the world and is held up as a benchmark for the use of technology in mining, areas for improvement were identified. Every area was critically assessed and investigated for the possible application of new techniques and technologies. Key learnings from the construction and production phases of Lift 1 were captured, documented and used as the starting point for the Lift 2 design. This paper shows how and why the design parameters initially used on Lift 1 were changed to become the current E26 Lift 2 development proposal. The main differences between the two block caves are highlighted. The alternative techniques and equipment that were considered are also discussed.
Small Open Pit
Assisted Block Cave
INTRODUCTION Northparkes Mines operates three copper-gold porphyry related deposits located in central New South Wales (Figure 1), identified as E22, E27 and E26. Production comes from the mining of open pits at E22 and E27 together with the underground operation at E26. A further deposit known as E48 is planned to be mined after E26. Northparkes Mines E26 underground mine is Australia’s first mine to employ the block caving method of mining (Figure 2). Construction for the first block cave, known as Lift 1, commenced in October 1993 and reached the milestone of full production in July 1997, three years and nine months later.
E26 LIFT 1
Pre Break
Block Cave 350 m
E26 LIFT 2
Narrow Undercut
FIG 2 - Schematic section through E26.
FIG 1 - Location map.
1.
Senior Mining Engineer, Northparkes Mines, PO Box 995, Parkes NSW 2870. Email: [email protected]
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By designing a low maintenance working environment, operating costs for Lift 1 were minimised and high productivities achieved. Last financial year (1998/99), the E26 underground mine produced 42 600 tonnes of copper/gold ore per underground employee, including contractors. Construction of the second block cave, known as Lift 2, is planned for completion in year 2003/2004, in order to replace underground production as Lift 1 winds down. E26 Lift 2 will produce at a planned rate of five million tonnes per annum over a period of about six years.
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In order to maintain the low operating costs for the second block cave mine, which is located 350 metres below Lift 1, it was essential that the low maintenance philosophy was continued. However, it was also realised that a high capital investment would result in an unattractive total cost per pound of copper and destroy shareholder value. This was particularly relevant given that by the time Lift 2 is planned to be commissioned in 2003/2004, there will be a number of large low-cost copper operations either coming on-stream or expanding their existing capacity.
LIFT 1
THE E26 OREBODY Geology The E26 deposit is associated with a quartz monzonite porphyry that intrudes volcanic rocks. The primary copper sulphide mineralisation occurs in the intrusives and surrounding volcanics. The mineralisation is located mainly in quartz veins and fractures, such that the deposits have a high-grade central core associated with dense stockwork quartz veining, for which grade and density of quartz veining decreases radially outwards into the volcanics (Figure 3). The potentially economic mineralisation at the E26 deposit occurs as an irregular subvertical column, 200 to 300 metres in diameter and is known to extend to over 800 m in depth (Dawson, 1995). The central core is dominated by bornite and lesser chalcocite, enveloped by a chalcopyrite dominant zone and then a pyrite-magnetite zone. The Lift 2 reserve is currently estimated at 24.5 Mt at a grade of 1.21 per cent Cu and 0.47 g/t Au.
Geotechnical The E26 Lift 2 rock mass comprises generally strong rocks, with rock property testwork showing mean intact rock strengths of 80 to 91 MPa for all rock types. However the maximum rock strengths range from 136 MPa in the biotite monzonites up to 227 MPa in the volcanics. The rock mass is generally well jointed with some narrow NW trending faults and shear zones evident. Unique to the E26 area is a post mineralisation event of gypsum that pervaded the local fracture system and contributes strongly to the geotechnical conditions of the deposit. The fracture system for Lift 1 of E26 was dominated by steeply dipping joint sets and the gypsum veining. The rock mass strength was reduced by the jointing and the gypsum veining. The Lift 1 rock mass strength ranged from 50 - 60 MPa. For Lift 2, the rock is generally more competent than Lift1 due to the lack of gypsum veining. For Lift 2, the rock mass ratings (RMR) range from 57 in the biotite monzonites down to 50 in the volcanics. In situ stress measurements were taken to obtain a better understanding of the regional stresses that would influence the stability of the large permanent openings. The anticipated in situ stress on the RL 9450 Extraction Level is: Major principal stress, σ1= 36.0 MPa at bearing 150°, dip = 17° The ratio of horizontal to vertical stress magnitudes is 1.8 on average.
REVIEW OF E26 LIFT 1 PROJECT As a starting point for the E26 Lift 2 design, the design team reviewed Lift 1, as the mine layout has proved to be efficient and flexible at generating high productivity with low operating costs. In order to close out the E26 Lift 1 project, two reviews were undertaken as construction finished. These reviews were seen as essential to capture learnings that could be applied to future projects, especially Lift 2, and were therefore carried out before the pre-feasibility study into Lift 2 started.
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South - North Section Rock Types
Grade %eCu 0.5-0.8
QMP Volcanics
0.8-1.2 1.2-2.5 2.5 +
Biotite Monzonite Diorite FIG 3 - E26 mine geology – Section.
Project management review The objectives of the March 1998 project management review were to summarise the expectations and philosophies that governed the E26 Lift 1 project development. The project outcomes were assessed against the original expectations and the key learnings documented then distributed to other North operations.
Design and construction review This technical and cost review concentrated on the design and construction areas of the Lift 1 project, assessing whether the chosen method, equipment, strategy or system met, exceeded or failed to reach the expectations, both from a technical and commercial viewpoint.
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Some of the specific key learnings which have been or will be incorporated into the design of Lift 2 were:
• allow enough time for each stage of the feasibility study; • involve operations personnel throughout the project; • the final feasibility study must be subjected to a critical review and risk assessment by independent experts, both technically and financially;
• the drawpoint brow is the most important part of the construction. By installing the right lining system first time, maintenance was minimised;
• by pushing the drawpoint spacing to the accepted boundary, a significant gain in LHD productivity was achieved; and
• the Toro 450E LHDs have exceeded the design productivity, due in part to the custom design of the Extraction Level. These conclusions were incorporated into the design philosophy for E26 Lift 2.
STAGES OF DESIGN Previously in 1997 an in-house development concept was formulated for extracting the Lift 2 resource as a 4.0 million tonnes per annum block cave operation, with identical development and infrastructure to Lift 1. All previous conceptual studies assumed a 400 metre extension to the Hoisting Shaft and no investigation into alternative ore handling arrangements had been undertaken up until 1998. In order to evaluate possible mining methods and ore handling systems, the design process went through three distinct stages:
• pre-feasibility study, • focussed investigation (extension to the pre-feasibility study), and
specifically the undercut shape and the use of the inclined undercut. It was predicted that caving will occur at 80 - 95 per cent of the planned undercut dimensions. In addition, the layout of the Lift 1 Extraction Level and exploration development will allow extensive coverage of the Lift 2 orebody for cave monitoring. Contingency planning is also in place to deal with any caving problems should they arise by utilising the existing Lift 1 openings for reactive propagation assistance.
Study schedule and resources The pre-feasibility study took place over six months. The work was predominantly carried out by Northparkes technical services staff with Lift 1 block caving experience and managed by the Development Division of North Limited, who had the feasibility study expertise. Consultants were used for specific areas, namely:
• ore handling system engineering; • review of geotechnical aspects, cavability and stress modelling;
• • • •
mine design review; Production simulation modelling; ventilation review; and dewatering and electrical reticulation design.
Results of pre-feasibility study At the end of this stage, four main conclusions were reached:
• development study. The Pre-Feasibility Study The objectives of the 1998 pre-feasibility study were to review alternative mining methods and ore handling strategies for Lift 2 of the E26 orebody, determine the optimum development concept and assess the project viability. The study had a sound base case to compare against in the 1997 conceptual design for Lift 2, which basically duplicated the existing layout at the then proposed Extraction Level 400 metres below Lift 1. However, for this stage of design, the team decided to use no pre-conceived ideas and started with a clean sheet of paper for many aspects of the design. This was not always possible as the potential mining methods and ore handling systems were all influenced by the existing capacities of Lift 1. An engineering audit was undertaken which documented the following information:
• existing capacities of the Lift 1 infrastructure and services; • contingencies put in place, either constructed or planned for further expansion;
• the planned expanded capacities of such contingencies; and • modifications to existing plant to achieve current productivities. The results of the audit provided the baseline from which design work could commence.
Cavability assessment The cavability of Lift 2 was investigated using empirical methods and by numerical stress modelling. The results of the assessment were that Lift 2 would cave more readily than Lift 1 due to:
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• higher in situ stresses; • favourable joint orientation; and • design changes made to improve the caving response,
• Lift 2 will add value to Northparkes; • Lift 2 will be best developed as a block caving operation; • both hoisting from a deepened shaft and conveying up to the current loading station investigating further; and
offered
advantages
worth
• significant scope existed for capital reduction and value enhancement flowing from further focussed design and engineering optimisation. The pre-feasibility study refined the mining method and ore handling options to three alternatives. Sublevel caving and open stoping were excluded on economic and risk factors. Of the mining methods, two block cave mining scenarios were evaluated:
• a single 350 m high lift; and • a double block cave of two 200 m high lifts. In parallel with the assessment of mining methods, different ore handling systems were considered which gave rise to the three main options for further evaluation. The three options were given the names Adavale, Altona and Orana for ease of identification. The options are summarised in Table 1:
FOCUSSED INVESTIGATION - EXTENSION TO PRE-FEASIBILITY STUDY Each combination of block caving scenario and ore handling system offered advantages and disadvantages. The two different ore handling systems for the single lift block cave showed no significant difference on an economic assessment at the cost estimate accuracy level. The double lift option would allow earlier production to be sourced from the higher grade reserve at
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TABLE 1 E26 Lift 2 Development options. Name of option
In summary, of the two block cave options, the single 350 metre high lift offered technical, capital and operational advantages over the double lift alternative. This is despite the better grade profile achieved by the double lift. Hence, the option known as Altona was selected as the preferred development concept for detailed study (Figure 4).
Mining method
Ore handling system
Adavale
Single lift block cave
Deepening of the Hoisting Shaft
Altona
Single lift block cave
Inclined conveyors to existing Loading Station
ALTONA DEVELOPMENT STUDY
Orana
Dual liftblock cave
Inclined conveyors and trucking
The objectives of the next stage, known as the Altona Development Study, were to:
an initially lower capital cost. Owing to these issues, it was realised that an extension to the pre-feasibility stage was required in order to meet the objective of selecting a preferred option for detailed evaluation. The pre-feasibility study was extended for five months to investigate the Altona, Adavale and Orana options.
Discussion on options For the Orana double lift option, with the decreased lead time, expenditure could be deferred until necessary for the production changeover from Lift 1. The capital commitment for the Lift 3 part of the reserve would also be deferred. However, the smaller cave height would result in coarser average fragmentation. The profitability of Lift 3 was questionable given that ore would be trucked up to the base of Lift 2 at a rate of 1.5 million tonnes per year, with a further 3.0 million tonnes per year to be sourced concurrently from E48. This would:
• bring forward the E48 construction and capital; and • lower the life-of-mine value. By mining two lifts it was also identified that there was a greater risk of waste mixing in with the ore columns. The single 350 metre high lift option offered greater scope for the beneficial impact of secondary fragmentation.
• • • •
detail the mine and engineering design; determine a cost estimate to 15 per cent accuracy; draft a project implementation plan; and seek commitment to the project.
The development study was planned to be carried out over a period of eight months. However the Lift 1 caving incident in November 1999 necessitated a diversion of study resources, resulting in a delay of about four months. The focus for each aspect of the study included:
• • • •
optimisation of the mine design and engineering; further reduction in capital requirements; further reduction in forecast operating costs; and improving the value of Lift 2.
Mine Design In the area of mine optimisation and design, the following areas were focussed on:
• • • •
review of cut-off/shut-off grade; optimisation of the extraction level and mine reserves; development ventilation requirements; stress modelling of major openings;
FIG 4 - Altona development option.
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• undercut review; • ground support requirements; and • general design optimisation.
To meet the objectives of the study it was seen as essential to use personnel experienced in block cave mine design. It was also essential that those operations personnel who had the experience of running Lift 1, and who would have to live with the designs proposed for Lift 2, were utilised throughout the stages of design.
Engineering design Alternatives to the Altona layout were proposed relatively late in the study but were thought worthy of assessment. Vertical conveying methods were investigated during the development study stage as it was initially assumed that savings could be made in reduced excavation quantities. High angle conveyors (HAC) of both the ‘sandwich’ type and ‘bucket elevator’ type were evaluated and inspected through site visits. Risk assessments of each combination of ore handling system were carried out. It was originally planned that the two jaw crushers installed on Lift 1 would last for 20 years and so process all production from Lift 1, Lift 2 and E48. The planned production for Lift 2 is 5.0 Mtpa, however each of the Lift 1 crushers are throughput limited to 3.8 million tonnes per annum. Therefore, the original designs were all based around two jaw crushers. In the pre-feasibility study stage, it was originally decided that all components of the crusher stations would be re-used with the exception of the ROM bin steelwork. However, by reviewing crusher options in the development study, it was acknowledged that significant cost-savings could be made by installing only one crusher, both in terms of crusher and associated system purchases, and excavation and support of a chamber.
LIFT 2 DESIGN METHODOLOGY It would have been easy to copy the designs from Lift 1 and apply them to Lift 2, however, the current climate of low commodity prices forced the design team to adopt a continuous improvement philosophy in order to increase the attractiveness of Lift 2. The emphasis was to minimise capital expenditure while still maintaining the excellent low operating costs achieved on Lift 1. As each section of the project was detailed, a continuous improvement process was applied where that particular section would be reviewed, by operations personnel or consultants. A critical examination was carried out questioning what each section was trying to achieve and how else might it be achieved, taking into account factors such as safety, quality, equipment and materials. The effects of any changes were also analysed for their influence on capital expenditure, operating costs and the construction schedule. On a project that is schedule driven, the design must be construction/development driven rather than the other way around. This is the case for most mining projects where the objectives are either to develop new areas to replace old, while maintaining continuous production, or fast-track a greenfields site to come on stream as commodity prices are at the top of their cycle. In the case of Northparkes, the critical date for project completion was set from the exhaustion date of the Lift 1 reserve. The construction, development and design schedules could all be back-calculated from this date. However, it was also important during the design process that consideration was given to constructability (ease of construction) issues.
Design team structure The pre-feasibility and development study teams were both small. Only two personnel were involved full-time on the project. When required, specific personnel were brought in to the team to cover areas such as mine geology, geotechnical engineering, electrical engineering, mechanical engineering, strategic planning and economic evaluation.
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Use of Experts It is easy for a designer to latch on to ‘their’ design and become blinkered to alternatives or suggested improvements. For this reason it is critical to subject the design to peer review, ideally by outside personnel with no vested interest in the project. The best design on the drawing board may also be difficult to install in the required time frame or in an operating mine environment. The designs were assessed by consultants and compared to operating mines through the use of site visits. The International Caving Study was also utilised as a source of contemporary block cave concepts.
Design review workshop As part of the development study, the mine design was reviewed in August 1999 by personnel not associated with the project. The plan was to bring together mining industry personnel with either mine construction experience or mine design expertise for a one-day design review workshop. The aim of the workshop was to review all elements of the E26 Lift 2 layout and supporting infrastructure in order to identify areas for potential capital savings and also discuss constructability issues that could save time or money. The team of ten personnel included consultants, contractors, project managers and senior mining engineers from the North group. The workshop was facilitated by a Northparkes manager who was not associated with the mining department. After setting out the objectives of the workshop, each area of the design was discussed for a set duration, typically 30 minutes per area. Lateral thinking was encouraged using a ‘what if’ approach. The issues and actions arising from each area discussion were then documented.
Design tools Mine design was executed using Datamine with detailed construction drawings later drafted in AutoCAD. Project scheduling utilised MS Project and ventilation network design was carried out on VentSim software. The results of the Altona Development Study are summarised by area in the following sections.
OUTCOMES OF LIFT 2 DESIGN PROCESS The continuous design process, always seeking a better design that would reduce capital or operating costs further has resulted in significant gains in value for the Northparkes site. Peer review by external personnel identified areas for investigation which have since been included in the final design. Based on learnings from Lift 1, several design changes were made to Lift 2 in order to improve the caving response. These include an enforced regular shaped undercut footprint and the implementation of a narrow inclined advanced undercut. The regular shaped undercut and complementary Extraction Level has resulted in a simple underground mine design, leading to both lower capital and operating costs.
ACCESS It was a recommendation from the Design Review Workshop that the Declines be steepened from a gradient of 1:7 to 1:6 to reduce the amount of development. For example, the Lift 2 Decline metres were reduced from 1652 m to 1530 m.
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Although the dual purpose access/conveyor drive concept offered some capital cost-savings due to less overall development, the effect on the project schedule was significant in that production would be pushed out by 50 weeks. There would also be an effect on construction of the ore handling system due to extended trucking of waste via the dual purpose drive. In this design, the conveyors would have to be installed in the backs to allow enough clearance underneath for trucks and crusher components. The identified savings in the dual purpose design would be further reduced once the cost of conveyor guarding and access platforms were included.
Primary crushing (similar to jaw crusher)
Secondary crushing (typical of a gyratory crusher)
ORE HANDLING SYSTEM The same basic ore handling philosophies developed for Lift 1 were kept for Lift 2. They were:
• minimise the amount of ore handling; • minimise the stages in ore block size reduction; • size the crusher to handle the largest rock size that the LHD could carry; and
• install a continuous system (the ‘Rock Factory’ concept) Engineering consultants were used to design the crushing and conveying components based on the Altona concept of using conveyors rather than deepening the Hoisting Shaft as per the Adavale option. The consultants were instructed to adopt the same philosophies but also minimise the capital cost, as Lift 2 has a relatively short-term life of around six years. The capacity requirement for all components of the system were reassessed by simulation modelling under the following guidelines:
• the capacity should be reduced where the impact on the ore handling system annual throughput is not adversely affected; and
• a reduction in required conveying capacity may enable a greater reuse of Lift 1 equipment. Of the different conveyor route options that were evaluated, the route heading north from the Lift 2 crusher then switching back south, up to the current orebins area, was selected as the preferred option. This option had the minimum number of conveyors and hence transfer points.
Crusher Engineering consultants proposed a single hybrid jaw-gyratory crusher for the required duty (Figure 5). This type of crusher was common in extractive industries in Europe but had never been used in an underground mine before. Site visits followed in 1999 and after financial evaluations were carried out, this type of crusher was selected as the preferred option. This unit accepts a larger feed size than the current jaw crushers and produces a finer, more uniform product. The single crusher design offered alternative routes for the ore handling system with potential capital savings. A total of ten options were considered each at a different compass point from the crusher. The concept of a dual purpose access/conveyor drive in a single tunnel was also investigated to compare capital costs with the base case Altona option. This option would require a second means of egress, a 300 m high vertical ladderway.
Conveyors High angle conveyors Although the excavations designed for the high angle conveyor (HAC) concept showed a saving in capital costs over the Altona inclined conveyor option, there was concern that the HAC would
340
FIG 5 - Krupp BK 160-210 jaw-gyratory crusher.
be operating at the extreme limit of current applications. The proposed vertical lift of two 175-metres high HAC installations, along with the required throughput and higher density ore, were collectively beyond the operating parameters of any existing installations worldwide. Major risks identified for application of a HAC at Northparkes were:
• suitability of E26 Lift 2 ore (density, abrasiveness, angular shape);
• operating costs; • maintenance requirements and mechanical availability; and • safety. The high angle conveyors were not selected for further evaluation due to the higher technical and operating risk.
Selected conveyor system The engineering consultants proposed that by using narrower but faster conveyor belts, the support structures and idler pulleys intensity would be reduced while still performing at the required capacity (Figure 6). The idler spacing can be increased from two to four metres, the belt speed increased from two to three metres per second and the belt width reduced from 1.2 metres to 1.0 metres. It was noted that re-use of Lift 1 conveyor components would be unlikely since during the transition to Lift 2, production will be conveyed from both Lifts concurrently. Therefore the Lift 2 conveying system must be in place, commissioned and running for about one year before any Lift 1 components can be salvaged. In the Altona option the following points are noted:
• transfer conveyors onto the main belts have been eliminated with the primary crusher discharge point aligned to coincide with the required orientation of the main conveyor;
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design and temporary support requirements, all effort must be made to ensure success at the first attempt as the consequences of failure to cave initiation and production security can be high. The narrow inclined advanced undercut proposed for Lift 2 (Figure 7) has the following advantages over the 42 m high undercut used on Lift 1:
Leg Frame Set (every 4m)
Sprinkler system
Pipe (safety barrier)
4.8 m
• the reduced ratio of ‘height to extent of the undercut excavation’ concentrates stresses over the back of what is an inherently more unstable excavation;
• • • •
1.9 m
1.5 m
5.5 m
FIG 6 - Section through conveyor drive.
it reduces abutment loading on the Extraction Level; the sawtooth shape enhances the instability; it reduces the opportunity for pillar loading; and as the narrow undercut contains less tonnes, access to the cave back is achieved quicker, giving earlier control of the cave propagation front from initial production activity.
Against this is the disadvantage of the immediate impact of coarse primary fragmentation.
• the maximum conveyor slope is set to one in 5.4 (conveyors used for Lift 2 ore handling are designed as steep as possible to minimise initial capital requirements, with the connecting conveyors from E48 designed at the required slope to intersect the transfer point); and
5m
• as most Lift 2 development waste will be mined prior to the commissioning of the ore handling system, waste handling facilities at E26 have been eliminated and conveyor capacities reduced accordingly.
18 m
Transfer points Another concept that was identified during the continuous improvement process was the use of a short transfer conveyor instead of a vertical transfer orebin. Each time a mine puts in a vertical surge bin it requires extra development at the top and bottom of the orebin, along with an associated access ramp. Although the additional transfer conveyor increases operating costs, the savings in terms of capital expenditure on excavation outweighed this increase.
12 m Drawbell
Major Apex
14 m
30 m
FIG 7 - Section through narrow inclined advanced undercut.
Overall system The selected system comprises a single Krupp BK jaw-gyratory type crusher which reduces the rock size from 3 m3 down to a -150 mm lump size. Ore is fed by a vibratory feeder directly onto a 1840 metre long inclined conveyor, known as C3, rising at a gradient of one in 6.4 to the transfer point. A 26 metre long transfer conveyor takes the ore through 90 degrees onto the second main belt which is at a gradient of one in 5.4. The second conveyor, C7, transports ore over a distance of 1140 metres up to the existing orebins above the Loading Station. Ore is automatically fed into weigh flasks then into the 18 tonne capacity skips. The skips are hoisted through 505 vertical metres to the surface by a ground mounted friction winder. The selected higher production rate of 5.0 million tonnes per annum significantly improved the base project value over the original pre-feasibility study plan of 4.5 million tonnes per annum.
UNDERCUT Undercutting is the critical component of cave initiation. As such, design work was very much driven by technical confidence and risk assessment rather than costs. In terms of sequence, blast
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Undercut footprint The regular rectangular shape is inherently unstable and maintains a minimum span dimension across the entire length of the footprint (Figure 8). Flexibility for extending the undercut was designed into the Lift 2 undercut level to allow further expansion to the east if required.
Undercut sequence Use of a narrow undercut results in a slower production ramp up than Lift 1, as only limited broken ore is generated during the undercutting sequence. This is exacerbated by the use of the advanced undercut proposed to limit the effect of abutment stress. However, the ramp-up scheduled for Lift 2 is under little pressure since production will be coming from Lift 1 and if required, from the two existing open cuts at Northparkes. In the case of an advanced undercut, the undercut must be developed, drilled and blasted to provide the stress cover for the drawpoint development below. The narrow, inclined advanced undercut designed for Lift 2 provides complete stress cover for the drawpoint development and results in full breakage of the undercut. The drawbell excavation and subsequent production of caving material then follows.
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N 162 m
206 m
Undercut design parameters The proposed undercut design consists of twin 4.2 metres by 4.5 metres drill drives placed on 12 metre centres on the corners of the future drawbell, 14.0 metres above the production level floor, as shown in Figure 8. The undercut consists of horizontal sidewall stripping holes and inclined holes overlapping adjacent rings over the major apex. This design has imposed a 50° angle on the major apex, assisting in the cleaning of undercut swell and subsequent movement of initial cave material into the drawbell. Holes were designed at 89 millimetre diameter, 8.5 metres long through the sidewall and up to 18 metres long over the apex. Rings are drilled on a 2.0 metre burden and a nominal toe spacing of 2.0 metres has been imposed by the undercut geometry. Uphole rings will be inclined forward by ten degrees to facilitate uphole charging and mucking of the horizontal face. A total of 71 drill metres per composite ring is planned at present. Rings will be charged with emulsion explosive of a relative density to ANFO of 1.0 with approximately 56 charged metres per ring. Rings will be fired initially one at a time and approximately 70 per cent of the fired tonnes will be mucked from the undercut level over a period of about eight months. Undercut initiation will be against an inclined slot fired against the western edge of the footprint.
EXTRACTION LEVEL
FIG 8 - Plan of undercut showing sequence by month.
As fewer tonnes are extracted compared to Lift 1, the narrow undercut will be retreated rapidly, allowing quicker access to bring drawbells on stream. The twin drift option allows confirmation of full breakage and ease of recovery if remnant pillars are left. The inclined section of the undercut rings are self cleaning with blasted material reporting straight to the Drill Drive as shown in Figure 9. Any narrow undercut requires extensive development and drilling which provides little return on a tonnes per developed metre or drilled metre basis. However the prime role of the undercut is to guarantee initiation of the cave, which this method will achieve. Detailed drill and blast design was carried out by the Julius Kruttschnitt Mineral Research Centre (JKMRC) utilising the 3x3Win software package, to determine the optimum parameters.
N Flat
Rill
Rill
The Lift 1 RL 9800 Extraction Level design (Figure 10) has proved to be efficient and flexible at generating high productivity with low operating costs. The design however incorporated much peripheral development in order to support the operating philosophy, which allowed two electric LHDs in the same quadrant to tip into one crusher. In order to reduce capital costs, the amount of development on the proposed RL 9450 Extraction Level has been minimised by designing alternative operating practices. The RL 9450 Extraction Level offered the most significant savings in terms of capital expenditure. As a result of the decision to originally locate the two crushers on one side of the Extraction Level, the design was radically altered from the Lift 1 layout. The change to the single crusher option simplified the layout even further. The main differences are:
• • • •
the perimeter drives were eliminated;
• • • •
six extraction drives instead of 14 extraction areas;
turning bays were eliminated; no link drives between extraction drives; drawpoint spacing increased from 14 m × 14 m to 18 m × 15 m; single LHD tip point; six rock breaking bays added; and maintenance facilities and operations support infrastructure reduced. The benefits of the layout shown in Figure 10 are:
• LHDs only interact at the ROM bin tip point, negating the need for perimeter drives;
• LHD trailing cables (assuming electric LHDs are used) are extended in a straight line resulting in less damage at the gate end panel section;
• as the airflow passes all the way through the orebody,
Drill Drives
ventilation requirements are reduced and can be easily regulated;
FIG 9 - Plan of undercut showing mucking of blasted drill drives.
• average tramming distances are reduced; and • the long straight extraction drives will allow high speed tramming and enable automation to be easily applied.
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is lost. An essential part of the design process was the use of simulation modelling to alleviate these concerns. The simulation modelling was carried out by consultants who input the operating parameters of the LHDs and secondary breaking rigs, along with the predictions of oversize rocks and high hang-ups (derived from the fragmentation predictions). The results showed that the Extraction Level layout would not restrict the planned Lift 2 productivity.
Extraction Level sequence In the proposed construction sequence, the six extraction drives will be fully developed prior to completion of undercutting, however no drawpoint take-offs will be attempted during this period. The undercut front will follow extraction drive development, with drawpoint development, drawbelling and construction activity following by no closer than 14 metres (45° shadow). The advantages of this approach are:
• abutment stresses are shouldered by 25 metre wide regular inter-drive pillars;
• the completed drives allow for through ventilation and two sided access to production level construction activity;
• drawpoint development, drawbelling and construction activity can be undertaken and accessed from one side while initial production can be accessed from the other;
• earlier production from completed drawbells will permit relief of the cave, preventing compaction of the narrow fired undercut and enable a controlled caving front to be established; and
• the strategy permits earlier production ramp-up than a complete pre-undercut approach.
Drawbell excavation FIG 10 - Comparison between Lift 1 and Lift 2 extraction levels.
Extraction Level design parameters As a result of the fragmentation predictions for Lift 2, the drawbell spacing was increased to 18 metres with 30 metres between Extraction Drives, centre-to-centre. This spacing meets Laubscher’s guidelines (Laubscher, 1995) for the predicted fragmentation and incorporates a robust major apex design. As was the case in Lift 1, the Extraction Level dimensions were designed around the tramming parameters of the Toro 450E LHD including turning radius, unit length and speed. Automation can be easily applied to this layout leading to a planned reduction in operating costs. The Extraction Drives average 266 metres in length with average tramming distances of 150 metres. Drawpoints were spaced at 18 metre centres along the Extraction Drive in an offset herringbone pattern. The angle of entry from the drive is 45 degrees with a chamfer of 22 degrees to allow for the LHD turning radius. As the LHD will load on the same line in a horizontal layout, there is no requirement to have a drawpoint much wider than the LHD unit as rocks can move behind the front tyres during loading and cause damage. The pillars between drawpoints are 14.2 metres wide. Drawpoint length has been maintained at 10.0 metres along centreline to the brow. Drawbell cross-cut length is planned at 12.5 metres. It was identified that the one way layout results in only six effective areas from which ore can be sourced and was therefore a potential restriction on productivity. If one Extraction Drive is out of service for any reason, 17 per cent of the production area
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Drawbell excavation will follow completion of the undercut overhead. ‘Skull’ shaped drawbells have been designed to similar standards as Lift 1 except that the initiation raise is planned to be excavated with a 660 millimetre diameter blind bored vertical raise instead of a conventionally mined inclined ladder raise. This raise will be bored to within 0.5 to 0.75 metres of the undercut and stripped open to provide a 1.5 to 2.0 metre relief opening for slotting across. Drawbell rings will be fired sequentially into the slot, retreating back to the drawpoint brow. This design produces drawbell shoulders ten metres above the drawpoint brow.
Roadways The drives will require good quality surfacing in order to maintain high speed tramming and hence high LHD productivity. It is a requirement that concrete roadways are installed to the same extent as Lift 1. The roadways will be graded for drainage towards the eastern access drive and to the Extraction Level sump.
DEWATERING Average water inflows to E26 are predicted at 17 litres per second. Peak water flows can be expected to be around 65 litres per second level for a one in ten year event. Once the E48 mine has broken through to surface, a total peak mine inflow of 75 to 80 litres per second could be expected. Two design options were considered for Lift 2:
• relocation of the Lift 1 pumps to a new lower pump station; and
• installation of new lower head pumps using the existing Lift 1 pump station for stage pumping to surface.
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TO SURFACE
L1 SCREEN ROOM
MPS1
RISING MAINS
BLOCK CAVE
L2 SCREEN ROOM ELPS
SUMP
MPS2
ED INCLIN R C3 EYO CONV
C3 SUMP
FIG 11 - Mine dewatering schematic.
The first alternative was not progressed due to the higher capital cost. The second alternative requires the installation of lower head pumps at Lift 2, pumping against a head of 350 metres to the current pump station. The existing Geho pumps could be used as stage pumps in their current configuration. This arrangement would also be used during E48 production as the Lift 2 pump station would remain at the lowest level of the mine. Unlike Lift 1, gravity will not be used as the prime drainage mechanism feeding the pump station. Though resulting in higher operating costs, a level secondary pumping concept has been adopted in order to reduce the vertical distance, hence capital cost, between the production level and mine bottom. The Lift 2 production level is designed as the prime water collection level. Not only will regular production water sources, such as drawpoint sprays, be collected here but also all water inflows coming through the cave (Figure 11). The Extraction Level was designed to drain to the eastern perimeter drive. Formed drains will be installed in the concrete roadways on the level directing collected water to the Extraction Level Pump Station (ELPS) located in the north east corner. Water will flow through a collection sump and into the main sump of the ELPS fitted with submersible pumps. It is assumed that some settling will occur in the main sump and provision is made for sump cleaning. Water will be pumped horizontally through one of the production drives to the screen room located in the south west corner of the level. Two 2.0 millimetre sieve
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bends located above a vertical sump will be used to remove miscellaneous trash and oversize material. The sump dimensions are 40 metres high and 3.0 metres in diameter. The Lift 2 Main Pump Station (MPS2) is located below the screen room and accessed from the Lift 2 Lower Decline. One slurry recycle pump will circulate water through the sump ensuring full mixing of solids, minimising the solids content to be pumped. The pump station will house two main pumps. The pumps have a designed capacity of 27 litres per second each. Rising mains will connect MPS2 to the existing Lift 1 screen room and vertical sump.
Emergency storage The current Lift 1 design provides emergency storage in the decline development below the pump station. For Lift 2, the RL 9450 Extraction Level represents the lowest point for any water flowing through the cave. Emergency storage is planned for the north east level sump area. Significant storage for major rain events can be provided by allowing the level to flood back up the eastern ramp without impacting on production. With the use of a back-up submersible pump in the level sump, this area can be drained over time after the event. Mine bottom infrastructure, such as conveyor tail ends and pump station, can only be flooded if the production level floods to the extent of water entering the crushers through the ROM bins. This concept limits the amount of capital to be invested in further vertical deepening of the mine.
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INFRASTRUCTURE Operations support infrastructure It was identified that the Main Workshop and Control Room can remain on Lift 1, as the facilities are located outside the subsidence region. Major servicing of electric LHDs can take place at the Lift 1 Workshop by using a generator set mounted on a vehicle travelling in tandem with the LHD. Communications and SCADA infrastructure can be extended from the RL 9800 Control Room down to Lift 2. However certain operations support activities will occur on the Lift 2 Extraction Level and excavations/facilities have been provided for:
• tyre storage and changing; • minor vehicle servicing including lubrication and daily pre-start checks;
• miscellaneous storage capacity; and • cribbing, first aid and emergency refuge. In addition, a secondary breakage magazine will be installed on the undercut access and miscellaneous parking facilities will be provided in disused level access drives and stockpile bays.
volts trailing cable which is held on an automatic reeling system at the rear of the unit. The current trailing cable capacity is 260 metres. These LHDs are currently forecast to last until 2005, two years into the production phase of Lift 2. Hence the Extraction Level design and operating methodology for Lift 2 is based around these units. However, the outer Extraction Drives (Drives 1 and 6) are too long for the current trailing cables. New longer cables will be required or diesel LHDs will be used. Simulation modelling has shown that three LHDs can supply the crusher at the target production rate. Further investigations into the use of a mixed electric/diesel fleet are in progress at time of writing.
GROUND SUPPORT For Lift 2, Northparkes Mines does not plan to have production personnel engaged in rockbolting, shotcreting or scaling. Hence, the high standards of excavation and ground support installation that were set on Lift 1 will again be employed to ensure that no revisiting of excavated ground is required over the operating life of the second lift.
LIFT 2 PROJECT SUMMARY
VENTILATION The improvements to the mine layout identified from the Design Review Workshop and through the continuous improvement process resulted in a simple ventilation network (Figure 12). The one way tramming layout of the Extraction Level simplifies the directing of air.
LOAD HAUL DUMP UNITS The current fleet of electric LHDs operating on Lift 1 consists of six Toro 450E units each with a nominal capacity of six cubic metres (about ten tonnes). The LHDs are powered by a 1000
The design for construction of the second lift currently encompasses the following:
• • • • • • •
decline and level development = 14 000 m, vertical development = 450 m, uphole raising = 900 m, rockbolts = 125 000, longhole drilling = 140 000 m, load and haul (Waste) = 770 000 t, and load and haul (Ore) = 780 000 t.
FIG 12 - Access and ventilation network.
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S DUFFIELD
From start of development to reaching the milestone of production from all 59 drawbells is scheduled to take 35 months.
• less disruption to production during the transition from Lift 1
CONCLUSIONS
to Lift 2;
The staged approach to mine design ensured that all suitable alternatives were assessed equally and that ample time was given for those assessments. This allowed the study team to refine concepts by systematically focussing on the preferred scenarios. The extension to the pre-feasibility study was justified in that the extra time taken ensured the option that offered the best value to the site was selected for detailed evaluation in the next stage. The continuous improvement approach therefore acted as a value improvement process. It is essential that the design is subjected to critical peer review. This also acts as a reality check. The one day review workshop that was held identified savings of over $A1 million. Also, the involvement of operations personnel in the design process was important as they are the ones who will have to live with the design. Block caving offers the best mining approach for E26 Lift 2 in order to minimise operating costs and maximise mine value. Based on learnings from the Lift1 operation, several design changes were made to Lift 2 in order to improve the caving response. These changes include an enforced regular shaped undercut footprint and the implementation of a narrow inclined advanced undercut. The design philosophies of low maintenance, continuous flow of ore, constructability and to minimise capital were established early and were the focus throughout the design process for Lift 2. A single lift block cave for the remaining E26 reserve has capital cost and technical benefits over a dual lift cave. In addition, the higher lift allows for greater secondary fragmentation giving better overall fragmentation.
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For the ore handling systems considered, the advantages of the inclined conveying layout over the shaft extension alternative are:
• greater hoisting capacity enabling ‘catch-up’ from poor monthly performances; and
• greater synergy with E48, reducing both development costs and lead time to bring that orebody on-line. The selected ore handling system does assist the economic viability of E48 and provides early access to the orebody to allow further drilling and caving assessment. Significant effort has been made to minimise capital costs by focussing on simple underground designs, in particular the straight-through production layout, the elimination of transfer conveyors and the use of existing infrastructure and services, either in situ or relocated.
ACKNOWLEDGEMENTS The author wishes to thank Northparkes Mines for permission to present this paper and to acknowledge all of the Northparkes personnel, the consultants and the specialist contractors who have contributed to the design of the second block cave.
REFERENCES Dawson, L R, 1995. Developing Australia’s First Block Cave Operation at Northparkes Mines – Endeavour 26 Deposit, in Proceedings Underground Operators’ Conference, pp 155-164 (The Australasian Institute of Mining and Metallurgy: Melbourne). Laubscher, D H, 1995. Cave Mining – State-of-the-Art, in Proceedings Underground Operators’ Conference, pp 165-175 (The Australasian Institute of Mining and Metallurgy: Melbourne).
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Modelling and Design of Block Caving at Bingham Canyon C J Carter1 and F M Russell2 ABSTRACT Open pit mining at Bingham Canyon commenced in 1904 and about two billion tonnes of ore have been extracted since then. There remain just under one billion tonnes of ore in the pit and this will be mined by 2015. Because the porphyry system extends below the ultimate pit limits, Kennecott Utah Copper Corporation initiated a study in 1997 to determine the viability of underground mining. The study was initiated at that time because the results were to be used to compare the economics of a further pit expansion to that of an underground mine. If a pit expansion was viable then overburden removal would have to start in the near future. This paper is a summary of the work that was carried out for the underground study. Block cave mining was determined to be the most appropriate method to mine the deposit. Two block cave areas have been planned with a total production rate of 72 500 tpd. Additional sources of ore will come from skarn orebodies that are associated with the porphyry deposit. The geology of the deposits is described together with the geotechnical techniques that were used to characterise the rock in the absence of more detailed data that are normally available. A general description of the mine plan is given along with the reasons for selecting the particular methodology used. The study was completed at the end of 1998 and demonstrated the viability of underground mining at Bingham Canyon. Funds have been approved for further drilling and technical studies, which will lead up to a feasibility study in 2004/5.
INTRODUCTION The Bingham Canyon porphyry copper deposit is located in the Oquirrh Mountains 40 km south-west of Salt Lake City, Utah, USA (Figure 1). Placer and underground mining began at Bingham Canyon in 1863 and some 5400 Mt of material have been excavated since open pit mining commenced in 1904. Bingham Canyon has been described as the largest man-made excavation in the world (the pit rim diameter is about 4 km) and one of the largest copper producers (nearly 13.5 Mt to-date). It is also one of the world’s lowest-cost copper producers. Current production from the open pit is 155 000 tpd of ore and 300 000 tpd of waste. It is a conventional drill, blast and truck haul operation utilising 314 mm diameter blasthole drills, 30 m³ and 44 m³ rope shovels and 218 t capacity trucks. Ore is hauled to an in-pit crusher that delivers to the Copperton Concentrator via a 6 km long conveyor system. About 20 per cent of the ore is loaded into rail cars and hauled to the North Concentrator. Copper concentrate is pumped to the smelter where it is filtered and dried before being smelted using flash smelting and flash converting furnaces. The Copperton Concentrator also produces a molybdenum concentrate. Electrolytic refining is used to produce finished copper. Silver and gold are recovered in a precious metal plant. Metal production in 1999 amounted to 279 300 t copper, 6300 t molybdenum, 11.4 t gold and 120 t silver. Current reserves (as at December 1999) amenable to open pit mining total 837 Mt at 0.6 per cent copper. Mineable porphyry reserves below the pit total 321 Mt at a grade of 0.7 per cent copper.
A study has been carried out to determine the viability of mining the porphyry resources below the pit using block cave mining methods and the following sections describe the results of the study.
GEOLOGY, HYDROLOGY AND RESERVES Geology The geology of Bingham Canyon has been the subject of many papers (Lanier, 1978; Titley, 1982; Babcock, Ballantyne, Phillips, 1994) and is therefore generally regarded as being well understood. The ore deposits are related to Tertiary igneous rocks which intrude Pennsylvanian sedimentary rocks, predominantly of sandstone or quartzite with interbedded limestone. The intrusives consist of stocks, dykes and sills of a variety of equi-granular to porphyritic rocks of monzonitic to latitic composition. Extrusives associated with the intrusions are generally of similar composition and include breccia, agglomerate and finer grained ash deposits. Large-scale folding and thrust faulting occurred prior to the intrusions. The present mountainous terrain at the mine is the result of these orogenic periods and subsequent extensive erosion. The orebody is of disseminated chalcopyrite with some bornite centred in and around the Bingham stock. The stock comprises an early monzonite intruded by a later quartz monzonite porphyry which is cut by latite porphyry dykes and quartz latite porphyry dykes. Early phases of rock alteration or metamorphism occurred prior to the mineralisation and were associated with the monzonite and quartz monzonite porphyry intrusions. This alteration included the formation of garnet skarn in the limestone. Later secondary biotite formed along fractures in the stock and extends locally into the sedimentary rocks. This biotite alteration is closely associated with chalcopyrite mineralisation. The mineralisation and alteration are zoned about a centre in the Bingham stock. The core zone consists of higher grade molybdenite and weak chalcopyrite mineralisation. Outward the molybdenite decreases and the chalcopyrite (and gold) increases to form the main copper ore zone. In three-dimensional presentation the deposit has the shape of a molar tooth with roots that extend to a depth of over 1000 m below the existing pit. Ore grade mineralisation extends into the sediments with especially high copper grade locally in the limestone beds. The limestone beds are locally of higher grade than their enclosing rocks and two areas, the North Ore Shoot (NOS) and Carr Fork, have been identified for potential underground mining. A geological cross-section through the open pit is shown in Figure 2 (Phillips, Smith and Harrison, 1997). The porphyry ‘roots’ of the deposit form the focus for the proposed block caving operations. Four roots have been identified, one in each quadrant around the stock, and two, the north-east and south-east, have been selected as the basis for designing an underground operation. Exploration continues in all quadrants of the deposit.
1.
Manager Mine Technical Services, Kennecott Utah Copper Corporation, USA.
Hydrology
2.
Principal Consultant, Rio Tinto Technical Services, England. E-mail: [email protected]
The open pit operation pumps about 200 L/s of water from wells in the pit established to drain the pit walls. The source of the water is from host limestones and quartzites. A hydrogeological
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C J CARTER and F M RUSSELL
FIG 1 - Location of Bingham Mine.
model, correlated with pump rates and piezometers established around the pit, was used to estimate inflows that would occur during underground mining. The model indicated that, while groundwater inflow rates would increase to 500 L/s, the drawdown would be beyond the projected caving areas and flowrates into the cave would be relatively small. Because of the large catchment area of the pit, estimates of pumping capacity also took into account seasonal rainfall as well as the possible cumulative action of spring melt of winter snowfall. Although a combination of conservative run-off factors and additional installed pumping (about three times average inflow) is used, the additional precaution of protecting the pump station with watertight doors was also taken.
Reserves Resource estimates are based on surface and underground geologic maps and information from several hundred drill holes, all but a handful of which are cored holes.
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Two methods of resource estimation are used, one for porphyry and one for skarn orebodies. In the case of porphyry the resource is modelled in 30 m x 30 m x 15 m (100 ft x 100 ft x 50 ft) blocks by kriging the assay composites. Other inputs are the rock or geologic model, grade zones to control smoothing by the kriging algorithm and limb zones to denote trends within the deposit. A three-dimensional modelling process is used for the NOS and Carr Fork skarn that takes into account stratigraphic controls for mineralisation. Much smaller blocks are used and faults are also included in this model. Categorisation of resources into reserves follows JORC practice. Table 1 lists those indicated resources that lie in the North East and South East block cave porphyry areas and the NOS skarn, below the pit, that meet the JORC requirements for probable reserves. Table 2 represents those resources that did not meet the criteria required for reserve classification. These resources include mining dilution and recovery losses. Although not a JORC category the term ‘Inferred mineable resource’ is used to represent those resources that are available for mine planning and scheduling
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MODELLING AND DESIGN OF BLOCK CAVING AT BINGHAM CANYON
FIG 2 - Geological cross-section (refer to the CD-ROM for colour explanation).
GEOTECHNICAL STUDIES TABLE 1 Probable reserves.
Introduction
Tons (Mt)
Copper (%)
Gold (opt)
Silver (opt)
MoS2 (%)
NOS Reserves
13.5
1.89
0.038
0.41
n/a
Block Cave Reserves
321.0
0.70
0.008
0.08
0.058
Total Reserves
334.5
0.75
0.009
0.09
n/a
In previous studies into mining the underground porphyry resources there were reservations about whether the weak ground conditions (from open pit experience) would permit both large-enough openings and coarse-enough flow characteristics to operate a mechanised cave (rather than gravity caving). However, the study established there are indications that:
• ground conditions improve with depth; • improved ground support techniques can support large openings economically, and
• an even draw-down of the cave will achieve good recoveries. Consequently, it was considered that mechanised caving would be operable and mining plans were developed accordingly.
Rock mass characterisation TABLE 2 Inferred mineable resources. Tons (Mt)
Copper (%)
Gold (opt)
Silver (opt)
Skarn Resources
60.3
2.59
0.036
0.44
n/a
Block Cave Resources
188.5
0.66
0.007
0.08
0.04
Total Resources
248.8
1.13
0.014
0.17
n/a
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MoS2 (%)
One of the problems facing the study was the lack of geotechnical information as most of the existing drill hole information is oriented around open pit expansions. There is limited coverage both at depth and where underground infrastructure would be developed. However Rock Quality Designation (RQD), core piece lengths and fracture frequency have been logged over a long period and compiled into a geotechnical database. As much of block cave design is related to Laubscher’s Rock Mass Rating (RMR) the study sought correlations and made assumptions that would enable RMR to be derived from the database without having to re-log the entire core. This work, summarised in Table 3, identified an increase in RMR with depth (upper and lower zones) and differences in rock type.
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TABLE 3 Rock mass characterisation. Geotech Zone
Parameter
Quartzite
Monzonite
Skarn
Upper
RMR RQD
32 18
45 35
61 65
Lower
RMR RQD
46 37
54 53
66 78
The result was considered to be within five points of RMR and sufficient for the level of study being undertaken. However, it is recognised that there are variations within each zone and rock type. Hence, when considering the rock mass on the large-scale, it is the global RMR that will control such factors as cavability, but when considering reinforcement design, the variability is more significant. This was taken account of by identifying the distributions of different RMR values within the sample population, thus taking account of subzoning without attempting to locate these zones geographically. For the purposes of subzoning, the range of rock quality encountered in the various rock types has been subdivided into four categories according to RMR (Table 4). TABLE 4 Categories as a function of RMR. Category
Description
RMR
1
Very Weak
< 25
RQD Equivalent < 10
2
Weak
25 - 40
10 - 25
3
Moderate
40 - 55
25 - 55
4
Strong
> 55
> 55
blocks were selected and analysed. The dominant five discontinuity sets identified were selected as significant where the numbers of observations were greater than ten per cent of the sample population. It was then assumed that the rock fabric orientations were consistent by rock type with depth. The data for fracture spacings relate to the upper zone of weaker and more fractured rock as exposed in the open pit. To estimate the fracture spacings in the lower zone of stronger and coarser rock, the spacings were factored according to the relative fracture frequencies between the two zones identified from the core logging.
Stress Stress measurements were carried out during earlier studies for underground mining. These indicated a relatively low vertical load due to the overburden pressures and a horizontal deviatoric stress. Studies and analyses of similar geological and mining environments confirm that the presence of a large open pit will be reflected in reduced overburden pressures and an increase in horizontal stress. At depths from surface less than 500 m, it is usual to find that horizontal stresses are greater than vertical. The presence of faulting and contrasts in rock competence will, however, tend to reduce the horizontal deviator. It would not be unusual in such conditions to encounter horizontal stresses of the order of 1.5 times vertical in a preferential direction. The in situ stresses around the planned depth of the underground production level are therefore expected to lie in the range of 10 MPa to 15 MPa, which are within the magnitudes of the stresses that have been measured at Bingham. The predictions of cavability of the orebody and the stability of the planned underground openings are not sensitive to stress magnitudes of this level. Greater stress levels would assist the caving process. Only in the unlikely event of a major horizontal deviator of greater than 2 would it be necessary to revise the conclusions concerning the undercut and the production level layout orientations.
Caving characteristics The different ranges of rock quality in each rock type and zone are reflected in the different proportions of the above categories. Table 5 summarises and necessarily approximates the rock mass characterisation of the lower zone where the block cave development and cave initiation would occur. TABLE 5 Rock mass characterisation of lower zone. Block
Rock
Global RMR
1
Percentage in Category 2
3
4
North
Monzonite
50
10
20
35
35
East
Quartzite
40
10
20
60
10
Skarn
60
0
10
15
75
South
Monzonite
55
5
10
35
50
East
Quartzite
45
5
35
40
20
Skarn
65
0
5
15
80
The principal tool for determining the cavability of a rock mass is the Laubscher stability diagram. The stability diagram relates the Mining Rock Mass Rating (MRMR) to the dimensions of the undercut required to propagate the cave, expressed as the hydraulic radius (HR), ie the ratio of the area of the undercut to its perimeter. The MRMR is the RMR factored according to other influences, such as high stresses. In the case of Bingham, a factor of unity is used because the mechanical factors influencing the cave, such as stress field and geometry, are within the normal range For all its approximations, the stability curve has been demonstrated to be a good predictor of cave initiation at a wide range of geological and mining environments. The range of MRMR for the lower zone at Bingham (cave initiation) is 40 to 55 resulting in an HR of 20 to 40. Given the footprint available, in excess of an HR of 60 m, it can be seen that cave initiation lies well within precedent for performing caves.
Fragmentation A suite of programs was used to predict the fragmentation of the orebody rocks and the achievable productivities based on the methodology developed for the Palabora project.
Rock fabric The mapping of rock fabric has been carried out in the open pit for slope stability analysis on an on-going basis. The available pit mapping data from the regions around the two target mining
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• Fragmentation Model (BCF), • Hang-up Model (HANG-UP), and • Discrete Simulation (ARENA).
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MODELLING AND DESIGN OF BLOCK CAVING AT BINGHAM CANYON
The fragmentation profiles were derived using the program BCF, which takes the rock fabric distribution and ranges and computes a sample distribution of blocks. The fragmentation was computed for each of the three main rock types: Monzonites, Quartzites and Skarns, for both the lower (Production Level) and upper (Cave Column) zones. The combined fragmentation distribution for each mining block was compiled by combining the relative proportions of each rock type in each zone. Figure 3 shows the resultant distribution. The fragmentation distribution from BCF is then input to the HANG-UP program. The program determines the proportion of hang-ups that are estimated to occur as a function of hang-up type.
• High hang-ups (>8 m above floor to top of bell); • Low hang-ups (>floor to 8 m); and • Drawpoint oversize (floor). The predicted incidence of high hang-ups is very low, becoming non-existent when the equivalent of over 100 m height of ore has been drawn down. It is therefore not anticipated that purpose built high hang-up rigs of the type at Northparkes or Palabora will be required at Bingham. When hang-ups do occur it is predicted that about 70 per cent will be low hang-ups, reducing to 60 per cent as ore is drawn down, the remainder being drawpoint oversize. The output from the HANG-UP program provides input into the production simulation model as each type of blockage event has its method of release and rules for time and resources derived from practical experience. The model then integrates all the facets of the production cycle, including:
• undercutting rate, • draw rate, and • equipment productivity. The oversize threshold at the drawpoint is taken as 2 m³. The mine plan provides for grizzlys to be located at the pass position in each production cross-cut. The simulation model therefore included rules for breaking oversize at the grizzly based on a grizzly dimension of 600 mm. For example, it was determined that for the initial, coarser material the grizzly would become blocked after 49 t of ore had been tipped and this would change to 163 t when the equivalent of 100 m height of ore had been drawn. As well as determine the number of resources, such as LHDs and secondary breaking rigs, the simulation was able to confirm that target production rates were achievable with the proposed mine infrastructure. In addition, the simulation proved very useful to determine the sensitivity of various input parameters to actual production rate.
Cave draw and dilution The range of rock mass properties in the cave columns and the contrast in the fragmentation profiles between the upper and lower zones gave concern about the draw characteristics of the column and the potential for high dilution entry and fines migration. This concern was addressed in two ways:
• a simplistic behavioural computer model of physical simulation of flow; and
• a parametric examination of the sensitivity of the economics of the project to changes in dilution parameters.
100 90 80
Percent (Mass)
70 60 50 40 30 20 10
10000
1000
100
10
1
0.1
0.01
0.001
0
Block Volume cu. m NE Lower
SE Lower
NE Upper
SE Upper
FIG 3 - Fragmentation profiles (refer to the CD-ROM for colour explanation).
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PFC modelling The behaviour of a differentiated cave column was modelled using the program Particle Flow Code (PFC) in two-dimensions. The cave column is modelled by layers of spherical balls of different sizes to mirror the fragmentation profiles and with inter-ball bonds to simulate strength and inter-ball friction to simulate differential flow. It is constrained to two-dimensions to restrict the problem size because compute time rises factorially with the number of balls in the model. A range of models was run to determine the physical parameters for the model and to examine the effects of the width of the footprint. The quantities drawn from each of the models run were accumulated and normalised according to the total material in each layer to give relationship of dilution against percentage draw of the column. One major shortcoming in the use of two-dimensional models to study dilution is that interparticle voids do not form readily to allow migration of smaller blocks through the larger ones. Within the knowledge of that restriction, the modelling exercise demonstrated the following:
variations to be mapped to any accuracy, the methodology adopted is to characterise the rock mass properties as simply as possible and to determine the proportions of each category likely to occur in each of the zones of interest. As described earlier, the variation in rock mass properties was ascribed to four categories that can be related to the range of ground conditions that will be encountered in the development openings. Table 6 is an example of a rock reinforcement schedule for infrastructure development. Other schedules were derived for undercut and extraction drives.
Instrumentation Provision is made in the planning and costing to install instrumentation to monitor the development of the cave and to give early warning of non-performance so that remedial measures can be instituted. This will include such as Time Domain Reflectometry (TDR) cables that record the intact length and a seismic monitoring system.
MINE DESIGN
• Significant inward migration occurs in the column. This models the phenomenon of wider draw zone influence in coarser material.
• The different layers in the model drew down remarkably evenly as long as the even draw was maintained over the full footprint. Within the limitations of the modelling exercise it was taken to indicate that mass draw could be achieved by even draw across the footprint and pointed to areas of further research when the project proceeds to the next phase. Dilution entry data obtained from the model confirmed the empirically derived data used in PC-BC.
Dilution sensitivity The second approach was to test the sensitivity of the schedule of tonnages and grades and hence the economics of the mine to different assumptions concerning dilution entry. It was found that the economics of the operation will be insensitive to changes in the percentage of relative migration velocities within the column. This is in part due to the fact that column outcrops in the pit floor and is substantially in ore. The current plan only calls for a single lift so there will be no mining beneath a caved out crater. ‘Cap rock’ dilution will only enter due to slope ravelling. Another source of dilution insensitivity is that the grades increase with height in the column above the planned production levels before dropping off above. Therefore, migration within the column tends to augment the earlier grades.
Reinforcement Reinforcement schedules for the different anticipated ground conditions were derived for cost estimation purposes. Because of the variable ground conditions and insufficient data to permit the
Overview The mine design provides for a conventional mechanised block cave layout utilising LHDs for loading ore from the North East and South East caves producing 54 400 tpd and 18 100 tpd respectively. Production is supplemented by ore from the high grade skarn ores using a combination of methods from drift and fill to sublevel caving. Production rates in the skarns vary from 3600 tpd to 11 800 tpd. Both porphyry and skarn mining operations will share common infrastructure, including shaft access, ventilation, pumping and ore haulage. In the block cave areas, LHDs will muck ore from drawpoints to ore passes, which will then feed ore to a heavy-duty wide conveyor belt. The conveyor will transfer the ore to a central crusher. The crusher will discharge onto a series of long inclined conveyors that will deliver ore to the existing crushed ore surface stockpile at Copperton. About 1500 m3/s of air will be circulated through the mine. Groundwater and service water will drain to settlers for pumping to surface. Maintenance workshops, stores and offices will be located underground, the only new surface facilities required are a changehouse, offices and a lamp room.
Block cave design Undercut The Bingham design has adopted the advanced narrow horizontal undercut to minimise stress effects on the extraction level. The undercut is planned to use cross-cuts at 15 m centres to mine a 4 m thick undercut. The undercut will precede development of the bells and drawpoints by approximately 45 m to 60 m.
TABLE 6 Typical reinforcement schedule. Category
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1
2
3
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Circular Arch Sets @ 1.5 m centres 300 mm cast concrete 250 mm cast concrete invert
Circular Arch 3 m grouted bolts @ 900 mm centres, 50 mm shotcrete, expanded. metal sheets 50 mm shotcrete
Parabolic Arch 2.4 m grouted bolts @ 1.2 m centres, 100 mm shotcrete, expanded metal straps @ 1.8 m
Parabolic Arch 2.4 m grouted bolts @ 1.8 m centres expanded metal straps @ 3.6 m
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MODELLING AND DESIGN OF BLOCK CAVING AT BINGHAM CANYON
Production level Production cross-cuts are spaced at 30 m and drawpoints are developed in the offset herringbone pattern at 15 m intervals. A perimeter tunnel provides for access and ventilation intake to each cross-cut. The choice between different extraction drawpoint layouts is seen to be largely dictated by the LHD usage and, in the case of Bingham, the selection of electric LHDs to be used to load ore into central orepass. Cross-cuts are planned at 4.2 m x 4.2 m sections to allow for sufficient pillar strength while maintaining suitable drawpoint spacing. The resultant drawpoint spacing (15 m x 15 m) reflects the draw control parameters for the granulometry of the initial fragmentation size range. Conventional draw cone theory would normally call for closer spacing of drawpoints to handle the finer caved rock of the upper column. For example, fine rock from secondary ores at Codelco is handled by gravity caving at 9 m spacings. However, there is substantial evidence from both experimental work and practice at other operations to indicate that mass draw of relatively fine material can be achieved evenly over a wide footprint with broadly spaced drawpoints.
The production cross-cuts are laid out parallel to the minor axis of the cave area. There is no evidence from either the stress field or the rock fabric which would dictate the orientation of tunnels in the South East area. The layout in the North East area has been rotated to minimise the impact of a major fault. Because the blocks are not elongated, however, the operational and economic impact of varying the layout orientation is not great. In production, electric LHDs with 4.5 m³ capacity buckets will muck from drawpoints to 3 m diameter orepasses located centrally in each extraction cross-cut. The orepasses are fitted with a 600 mm grizzly to prevent large rocks entering the pass. Adjacent to each orepass will be a 2.4 m diameter return ventilation raise that will downcast air to a return airway about 20 m below the extraction level. This airway, in turn, connects to the main return system for the mine. Figure 4 shows the plan for the North East extraction level.
Optimisation of extraction rate and level The method used to determine the optimum production rate and extraction level was that developed during the Palabora block caving study and utilises a technique termed ‘Economic Surface
FIG 4 - Plan of the North East Extraction Level. (Refer to the Cd-Rom for colour explanation.)
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Evaluation’ (Kear, 2000). Input data to the model consist of ore reserve data, capital costs and operating costs. Output is a ‘surface’ showing the relative value, in NPV terms, for the options under consideration. No constraint was placed on the planned column height. The resultant column heights average about 350 m and are within precedent.
Production scheduling Production schedules for use in the financial evaluation were generated using the Gemcom PCBC program (Diering, 2000). The input to the program is blocks from the ore reserve and the drawpoint positions. Parameters for dilution entry and density are generated and the program calculates the mixing between the columns and a dilution capping. The output reports total tons, grade and tons of dilution per block.
Ore haulage The selection of ore handling systems was based primarily on expected fragmentation and Bingham’s favourable experience with in-pit crushing and conveying. Orepasses, located at the centre of each production cross-cut, are used to transfer ore from the extraction level to the gathering level. There will be grizzlys at the top of each orepass to ensure that material oversized for the gathering belt is not passed through. Mobile rockbreakers will be used to fragment the oversized material. Below the North East block, the gathering level consists of two 2.4 m wide heavy duty belt conveyors converging on a centrally located transfer. The belt conveyors are loaded by apron feeders located at the bottom of each orepass and discharge into a common transfer, which feeds the gyratory crusher. The gathering level below the South East block consists of a single 2.4 m wide heavy duty belt conveyor that passes below each orepass and discharges into a 90 m high bin. The bin feeds the conveyor in the south-east transfer drift that also discharges into the gyratory crusher.
The 60 - 89 primary crusher is rated at 4200 tph for a 200 mm open end setting and is equipped with a 750 kW motor. The crusher feeds via a 500 t storage bin onto a 1.8 m wide inclined belt. The first belt is 1900 m long and connects with the second inclined belt, 3400 m long, in the vicinity of the main access shaft. Surface belts then convey the ore to the stockpile at Copperton Concentrator. The proposed ore flow schematic for the block cave operation is shown in Figure 5.
Personnel and material access The main personnel and material access will be via an 8.5 m diameter shaft located in Bingham Canyon and beyond the effect of any caving action. A service cage 3 m wide by 7.6 m long in plan will be used to convey personnel (150 capacity) or material (22 t capacity). An auxiliary cage will also be fitted in the shaft and emergency egress will be provided by emergency hoist systems in the Pit Intake shaft. Shafts and other permanent excavations are planned beyond the influence of any caving action and, for the purpose of the study, the area defined by an angle of 55° from the horizontal drawn from the base of the mining area to surface, was assumed.
Ventilation It is planned to circulate about 1500 m3/s of air through the mine to serve both the skarn and porphyry mining areas. Air will be downcast via the NOS shaft, Canyon Service shaft and the Pit Intake shaft. In the winter intake air in the service shaft will be heated to above 0 C. Air will be exhausted by the Canyon Exhaust shaft and a similar shaft located in the pit.
Other resources The ‘base case’ established during the study does not include the resources contained in the North West and South West roots of the porphyry. These areas are relatively poorly drilled and have the potential for development of additional reserves. They are targeted in a proposed drilling program to be carried out prior to the full feasibility study.
FIG 5 - Ore flow schematic.
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MODELLING AND DESIGN OF BLOCK CAVING AT BINGHAM CANYON
SUMMARY
BIBLIOGRAPHY
The study has demonstrated the viability of underground mining at Bingham. Funds have been approved for further drilling and technical studies that will lead up to a feasibility study in 2004/5. Areas for further study include the possibility of using an inclined or ‘false footwall’ cave of the North East block, rock flow analysis to determine optimum drawpoint spacing and a review of ore haulage systems. The study benefited from evaluation tools and techniques that have become available in recent years due to the increased interest in block cave mining as a safe low-cost method. It is anticipated that further interest, possibly as a result of conferences such as MassMin 2000, will prompt the development of more tools and techniques.
Babcock, R C, Ballantyne, G H and Phillips, C H, 1994. Summary of the Geology of the Bingham Canyon District, Utah, Bootprints Along the Cordillera Symposium, Tucson Arizona, October. Diering, T, 2000. PC-BC A Block Cave Design and Draw Control System, in Proceedings MassMin 2000, pp 469-484 (The Australasian Institute of Mining and Metallurgy: Melbourne). Kear, R M, 2000. The Use of Evaluation Surfaces to assist in the Determination of Mine Design Criteria, in Proceedings MassMin 2000, pp 57-62 (The Australasian Institute of Mining and Metallurgy: Melbourne). Lanier, G et al, 1978. General Geology of the Bingham Mine, Bingham Canyon, Utah, History and Pre-Excavation Geology, Economic Geology, 73:1228-1241. Phillips, C H, Smith, T W and Harrison, E D, 1997. Alteration, metal zoning and ore controls in the Bingham Canyon porphyry copper deposit, Utah, Guidebook prepared for Society of Economic Geologists Field Trip, October 23 to 25, 1997, Series No 29. Titley, S R, 1982. The Bingham Mining District, Advances in the Geology of Porphyry Copper Deposits, pp 159-164.
ACKNOWLEDGEMENTS The authors were privileged to be part of the project team that produced the study. The paper presented here is the work of all those who participated in the study and the role of the authors has simply been to present this work to a wider audience. The permission of Kennecott Utah Copper Corporation and Rio Tinto Technical Services to present this paper is gratefully acknowledged.
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Considerations for Design of Production Level Drawpoint Layouts for a Deep Block Cave A R Leach1, K Naidoo1 and P Bartlett2 ABSTRACT Premier Diamond Mine is planning a new, mechanised, block cave, with a production level approximately 1000 m below surface. A high degree of tunnel damage is expected, due to stress and the relatively low in situ strength of kimberlite. An evaluation of the timing of tunnel excavation, tunnel size and spacing has been carried out to define critical parameters for planning layouts and development schedules. Extensive use has been made of three-dimensional FLAC3D numerical models, calibrated against known rock mass behaviour at 650 m depth, to examine the influence of undercut advance on production level excavations. These were able to represent the effect of stress rotation on tunnels during undercut advance and the effect of an inclined undercut. The analysis has provided guidelines that are possibly applicable to any deep block cave layout. The choice of drawpoint layout is of considerably less significance than the proportion of ground extracted on the production level elevation prior to advancing the undercut. The analysis shows that damage to production development will be potentially severe if it is all excavated prior to advancing the undercut, or is mined directly beneath the undercut face. The development of the main production tunnels alone ahead of the undercut may be feasible, but no breakaways should be put in as they appear to trigger a rapid increase in the severity of damage. The optimal sequence is when all production tunnels are developed at least 10 m in plan behind the undercut, where the induced peak field stress is less than 20 MPa and a maximum support pressure of 100 to 500 2 kN/m is require to contain movement. The assessment showed that the conditions are not significantly improved by increasing drawpoint spacing from 15 m to 18 m. The vertical spacing between the undercut and production level beneath is investigated. Provided production level development is carried out beneath the undercut, there would appear to be little benefit in increasing the vertical spacing between the undercut and production level beyond the current 15 m to 18 m.
INTRODUCTION A new block cave operation is planned at Premier Diamond Mine, which is working a kimberlite pipe hosted in a sequence of quartzites and norites at depth. At approximately 1000 m depth the new block cave will be some 300 m deeper than current operations at the mine. Production level development for the existing BA5 and BB1 blocks has experienced a range of stability problems arising from induced stress changes and the comparative weakness of kimberlite. Given this, it is anticipated that stress damage will be considerably more severe on the new C-Cut elevation, and that design criteria for production level excavations and development sequencing require thorough evaluation. Recommendations for a range of key criteria have been developed, providing information upon which layout selection and design decisions for the new undercut and production level can be based. Specific areas of concern have included: 1.
effect of tunnel size;
2.
effect of variation in kimberlite strength;
3.
merits of different production drawpoint tunnel layouts (El Teniente and Herringbone type layouts were compared);
4.
effect of minimising production excavated prior to undercutting;
level
development
1.
Itasca Africa (Pty) Ltd, Box 38425, Booysens 2016, South Africa.
2.
Premier Diamond Mine
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5.
effect of increasing drawpoint spacing from 15 m to 18 m; and
6.
effect of increasing the vertical spacing between undercut and production level from 15 m to 18 m.
Note that excavation spacing alterations have been considered only in terms of the likely benefits in rock mass conditions around tunnels. The dimension changes are assumed to remain within tolerable limits from a practical draw zone spacing, or draw angle, point of view, although no detailed studies have been carried out to check this at this point in time. Additional work is required if significant changes in spacings are used. However, as results below indicate, substantial changes in spacings are probably not warranted. Analyses have made use of numerical models constructed using the three-dimensional finite difference program FLAC3D (Itasca Consulting Group, 1997), calibrated against known behaviour on 650 level in the BA5 block, then projected to the future 1020 level. The typical geometry of the models used is shown in Figure 1. Each model represents a number of drawpoints and includes the undercut geometry, providing a reasonable representation of changes in stress magnitude and orientation as the undercut is advanced over pre-existing production level excavations. The FLAC3D models include various levels of detail to examine arched production tunnels beneath a flat undercut, and both El Teniente and Herringbone drawpoint layouts beneath a saw-tooth undercut. In all cases the models make use of similar mining geometry, mining sequence, in situ stress regime and rock material parameters and represent a section of the kimberlite pipe, covering three main production tunnels (approximately 90 m), and extending a distance of 70 to 90 m into the pipe from the contact, approximately to the pipe centreline. Fixed boundary conditions were applied to model sides and base, hence each side behaves as a plane of symmetry. In all models an undercut was advanced sequentially across the model. This comprises undercut tunnels on a 15 m spacing, giving 15 m wide faces, each face leading the next by 15 m to give an overall undercut advance angle of 45 degrees to the production drives. In most model analyses, three main production level tunnels were mined prior to advancing the undercut.
GEOTECHNICAL CHARACTERISTICS OF THE PREMIER KIMBERLITE Although not exceptionally deep by other Southern African mining standards, squeezing ground problems can be anticipated due to the relatively weak nature of kimberlite. The estimated in situ stress fields for the 650 and 1020 levels, based on limited in situ stress measurements and mine-scale numerical modelling, are listed in Table 1. It is assumed that the Maximum Principal Stress is vertical and horizontal stresses are approximately equal in all directions. A range in kimberlite quality occurs across the pipe, and rock mass behaviour under the possible range in strengths has been examined. Rock materials include weak TKB (Tuffisitic Kimberlite Breccia), stronger TKB, and typical Hyperbyssal kimberlite in the core of the pipe.
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TABLE 1 Estimated in situ stress levels. Stress component
650 level
1020 level
Vertical
16 MPa
27 MPa
Horizontal
14 MPa
20 MPa
TABLE 2 Rock mass properties. Weak TKB
Strong Hyperbyssal Country TKB kimberlite rock
Base properties UCS (MPa)
50
100
150
200
RMRLauscher
50
50
55
70
mI
16
16
7
16
ν
0.25
0.25
0.28
0.2
Derived in situ properties Young’s modulus (GPa)
10
10
10
40
Cohesion (MPa)
3.4
4.4
5.2
10
Friction Angle (degrees)
31
37
34
47
Dilation Angle (degrees)
25
25
15
15
GENERAL MODEL BEHAVIOUR
FIG 1 - Typical external (top) and internal model geometries for the herringbone (centre) and El Teniente (bottom) layouts.
Rock mass properties used in numerical models are listed in Table 2. FLAC3D assumes the rock mass to be a continuum. It was assumed that, due to the presence of joints, all rock types have no tensile strength on a large-scale. As will be seen below, a small tensile strength may have been preferable. Rock mass properties are based on typical laboratory test values, with strengths downgraded on the basis of rock mass ratings using a methodology proposed by Hoek (1988), and more recently described by Hoek, Kaiser and Bawden (1995). Rock moduli as well as strengths are estimated from rock mass ratings. Note that the dilation angle values listed in Table 2 are assumed, based on experience.
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The overall changes in stress due to undercut advance, and behaviour of the rock mass around the excavations is similar in all models. Each point on the production level elevation undergoes a stress change during the modelling sequence as the undercut is advanced. An example graph, showing the stress change for a point 4 m into the sidewall mid-way between draw point breakaways along the central modelled production level tunnel is shown in Figure 2. Stress starts at the pre-mining value of 27 MPa and increases to typically 40 - 45 MPa when production level tunnels are mined out and sit well ahead of the undercut. A stress increase occurs as the undercut is advanced, reaching 60 MPa at the point selected, and finally a large stress decrease occurs after the undercut has passed above the point. The distribution of stress modelled around the advancing undercut is shown in Figure 3. This shows a vertical section across the model. The zone of raised stress ahead of the undercut is clearly visible and curves back over the undercut excavation, and presumably would tend to drive the caving process. Damage in the sidewalls of production level tunnels is initiated well ahead of the advancing undercut (when tunnels are pre-developed), this increases in severity as the undercut passes over, with some tendency for damage to migrate upwards towards the undercut. Stress rotation beneath the undercut appears to induce increased damage in the tunnel corners. Highest stress levels occur around the undercut drilling tunnels, and may tend to cause ring drilling and charging problems. A large volume is disturbed around the undercut, extending to roughly mid-way to the production level. Severe damage is indicated on the undercut level, ahead of the undercut, between drilling tunnels and rings excavated in the adjacent tunnels.
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CONSIDERATIONS FOR DESIGN OF PRODUCTION LEVEL DRAWPOINT LAYOUTS FOR A DEEP BLOCK CAVE
FIG 2 - Typical stress history for a point in the sidewall of a production level tunnel.
COMPARISON OF C-CUT BEHAVIOUR TO CURRENT BA5 EXPERIENCE Premier mine has considerable experience with production level tunnel behaviour at 630 m and 730 m depth in the BA5 and BB1 areas. To provide a means of interpreting results of C-Cut models at 1000 m depth in terms of realistic underground behaviour, the standard Herringbone model was run with in situ stress values corresponding to the BA5 elevation. The modelled damage at this 630 m elevation can be compared to observed behaviour to provide a means of roughly calibrating the models. BA5 model results can be compared to similar C-Cut model results to give a measure of the percentage difference in conditions on the C-Cut elevation when compared to observed conditions in the BA5 area. BA5 practice has been to excavate most production level development, excluding drawbells/drawcones, prior to undercutting, where the undercut is flat and tunnels run perpendicular to the direction of undercut face advance. The model is adequately similar to the actual situation to provide a reasonable measure of calibration. Comparison of BA5 and C-Cut model results show a considerable increase in the severity and extent of damage as mining depth is increased by approximately 400 m. Table 3 gives a summary of the main changes. These are probably a reasonably accurate indication of the differences in conditions, but may not accurately represent actual deformation levels. Note that the in situ stress in the model, as described in section
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2.2 on page 3 above, changes from 16 MPa in the BA5 case to 27 MPa in the C-Cut model, an increase of 69 per cent. In general the increase in damage and movement is significantly greater than this, and, relative to current BA5 experience, a quantum leap in rock mass mobility and support requirements would be expected. This may be countered if careful sequencing is done to minimise the exposure of production level excavations to full stress changes around the undercut.
INFLUENCE OF PRODUCTION LEVEL TUNNEL SIZE First step of the production level analysis has been to settle on a best compromise tunnel size, for the range in rock mass conditions anticipated, from weak TKB to strong Hyperbyssal kimberlite. From a production standpoint larger tunnels are preferable, to fit bigger equipment. The rock engineering view is that smaller tunnels tend to be more stable and are less costly to support. Three tunnel sizes were considered: 3.5 m, 4 m and 4.5 m, where both height and width are identical, and the hangingwall has a circular arched profile. Models of the three tunnel sizes were run in a stronger TKB rock mass. The main concern is the comparative difference in damage. Peak damage is induced around the tunnels when lying immediately below the undercut face position. The difference in damage can be summarised in terms of the maximum value of strain, plus the maximum depth of damage in sidewalls or hangingwall using five millistrains (5e-3) as a cut-off for easily
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FIG 3 - Section through typical model showing stress concentrations ahead of the undercut, and changes around production level tunnels. Arrows indicate stress orientation.
TABLE 3 Comparison of behaviour in models representing BA5 and C-Cut mining. BA5 model Width of sidewall failure zone around tunnels ahead of undercut.
Approx 1 m Approx 2 m
Extent of damage within 56 per cent of pillar area pillars between tunnels failed once undercut has passed. Maximum strain in tunnel walls (severity of damage) Modelled maximum tunnel sidewall movement
C-Cut model
1.2 x 10-3
2 cm
Percentage increase 100 per cent
Pillar area entirely failed
79 per cent
3.5 x 10-3
192 per cent
6 cm
200 per cent
TABLE 4 Effect of tunnel size. Tunnel Size
3.5 m
Maximum strain
Depth of s/wall fracture
Required support pressure
0.013
1.7 m
46 kN/m2
4m
0.014
2m
54
4.5 m
0.017
2.2 m
59
The main conclusion would appear to be that it is not essential to limit tunnel design, and that a 4 m tunnel is feasible. Rate of strain change tends to indicate a potentially large jump in damage from a 4 m to 4.5 m tunnel.
INFLUENCE OF ROCK MASS STRENGTH discernible damage. These figures are summarised in Table 4. Included in the table is an estimate of the dead-weight support capacity requirement using the conservative assumption that the full thickness of the failed zone would have to be supported. Note that these levels of support appear tolerable, as, for reference, 50 kN/m2 equates to a typical rock bolt of around ten ton capacity for every 2 m2 of exposed rock wall, while a 1 m space pattern provides approximately 100 kN/m2.
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The effect of varying rock mass strength was considered for a 4 m tunnel size. The extent of damage to tunnel sidewall zones ahead of the undercut increases as the rock mass strength is decreased, and the zone of influence around the undercut also increases and considerable shear and tensile relaxational damage appears to occur beneath the undercut. There is little difference between stronger TKB and Hyperbyssal kimberlite, however in weak TKB there is a large increase in the magnitude and extent
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CONSIDERATIONS FOR DESIGN OF PRODUCTION LEVEL DRAWPOINT LAYOUTS FOR A DEEP BLOCK CAVE
FIG 4 - Plan of production level development showing stress distribution due to advancing undercut. Top: herringbone layout, below: El Teniente layout.
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of damage. Results are compared in Table 5. Again, however, maximum deadweight support pressure does not appear excessive, remaining below 100 kN/m2.
TABLE 5 Effect of rock conditions. Rock type
Max strain
Depth of s/wall fracture
Required support pressure
Weak TKB
0.03
2.4 m
65 kN/m2
Stronger TKB
0.014
2m
54
Hyperbyssal
0.014
2m
54
The main overall conclusion to be drawn from this tunnel size, and rock mass strength assessment is that 4 m is a reasonable compromise width and height for production tunnels. Required support pressure to carry dead-weight loading of potentially loose rock, appears to be a maximum of 65 kN/m2, even in weaker ground, requiring a bolt pattern using possibly 1.2 m spacings.
COMPARISON OF HERRINGBONE AND EL TENIENTE DRAWPOINT LAYOUTS Premier mine currently uses a Herringbone drawpoint layout for the BA5 and BB1 block caves. Some problems relating to drawpoint stability have been encountered, hence a comparison to an alternative layout, favoured at El Teniente mine in South America has been carried out. The objective was to identify any geotechnical features of the El Teniente layout which would make it significantly better than Premier’s preferred Herringbone. A worst case situation was compared, where main tunnels, drawpoints and troughs have all been extracted on the production level prior to undercutting. The percentage of ground mined on the undercut elevation has been made identical in both models, to ensure a reliable comparison. Hence in both cases the main production tunnels are spaced 30 m apart, centre to centre, and drawpoints are broken off these every 17 m along them. This latter figure is greater than the 15 m distance normally used in the Premier herringbone layout. Areas where significant differences might be expected between the two layouts include the distribution of stress (Figure 4), and hence damage, in the pillars between production level tunnels, and the potential for hangingwall movement over breakaways.
With drawpoint breakaways spaced on 17 m centres along the main production tunnels in both cases, the Herringbone layout leads to a pillar with a wider, and hence stronger core. Herringbone pillar width is 13 m while the El Teniente pillar is narrower at 11 m. Note that in the normal layout used, the Herringbone drawpoint spacing is 15 m, giving an 11 m pillar width, identical to the El Teniente pillars. In general, however, peak stress concentrations and stress values centrally in pillars are similar in both layouts, Figure 4. In both cases peaks occur close to the ends of the pillars, adjacent to the main production tunnels, and reach maximum values in the acute bullnoses. Depth of failure around pillars is similar in both layouts. Failure extends approximately 2 m into the sidewalls of tunnels, with an increased depth of failure on bullnoses and corners. The maximum tunnel sidewall movement is 13 cm in the El Teniente layout compared to 12 cm with the Herringbone layout. Behind the advancing undercut, damage in pillar cores is more extensive, with proportionally more zones damaged in the El Teniente layout than in the Herringbone layout, although the difference is small. The reason for the increase in damage within these pillars behind the undercut is partly the increase in loading as the undercut passes over, the rotation of stress and loss of confinement, the removal of large volumes corresponding to drawbells/drawcones, and possibly the zero tensile strength used. One of the main differences between the two layouts, are the breakaway intersections. In the El Teniente layout, both left and right breakaways from the main production tunnels are at the same point, giving rise to large intersection spans over which more damage is induced than over Herringbone breakaways. Tunnel hangingwall deformations are greater over the breakaways, and higher strain areas tending to connect to similarly damaged regions around the undercut, above. If the full development layout results in identical percentage extractions in both Herringbone and El Teniente drawpoint layouts, then there is a marginal preference for the Herringbone layout, in terms of geotechnical criteria.
TIMING OF PRODUCTION DEVELOPMENT RELATIVE TO UNDERCUT ADVANCE Considerable damage is done to production level excavations when they are developed ahead of the undercut and then subjected to changes in stress as the undercut is advanced over them. To reduce the extent of damage, consideration can be given to excavating a proportion of the development after undercutting. Four variations in excavation timing have been examined: 1.
All production development, excluding draw-bells, is developed ahead of undercut (34 per cent of production level area).
TABLE 6 Effect of percentage extraction of production development prior to undercutting on rock mass behaviour. Develop-ment ahead of u/cut
Mined pre-u/cut
Peak Stress
Sidewall movement
Shear strain in tunnel skin
Damage to pillars
Production tunnels, drawpoints and troughs
34 %
87 MPa
6 cm
3.5 x 10-2
98% failed
Production tunnels, and draw points
22 %
75 MPa(13 % impr)
4.8 cm(20% impr)
2.8 x 10-2(14% impr)
87% failed(11% impr)
Production tunnels, only
13 %
65 Mpa(25 % impr)
3.5 cm(42% impr)
1.4 x 10-2(60% impr)
83% failed(15% impr)
No develop-ment ahead of u/cut
None
47 MPa
3 cm(50% impr)
1.0 x 10-2(71% impr)
69% failed(29% impr)
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TABLE 7 Effect of drawpoint spacing and vertical height from production level to undercut. Mining pre u/cut
Peak stress
Sidewall movement
Shear strain
Damage to pillars
3.5x10-2
98% failed
Base case: 15 m spaced draw points, 15 m below undercut to production level Prod. tunnels, d/points and troughs pre-u/cut
34 %
87 MPa
6 cm
No development ahead of u/cut
None
47 MPa
3 cm (50% impr)
1.0 x 10-2 (71% impr) 69% failed (29% impr)
Increased draw point spacing: 18 m spaced draw points 3.3 x 10-2 (6% impr)
Prod. tunnels, d/points and troughs pre-u/cut
31%
87 Mpa (nil impr.)
6 cm minimal impr.
No development ahead of u/cut
None
46.5 MPa
3 cm(50% impr)
0.9 x 10-2 (74% impr) 73% failed (27% impr)
3 cm (50% impr)
1.4 x 10-2 (60% impr)
90% failed (8% impr)
Distance from u/cut to production level increased from 15 m to 18 m No development ahead of u/cut
None
46 MPa
2.
The main production tunnels and draw point breakaways are developed ahead of undercut (22 per cent of production level area).
3.
Only the main production tunnels are developed ahead of undercut (13 per cent of production level area mined out).
4.
No development ahead of undercut. The main production tunnels are excavated immediately behind the undercut, with draw point breakaways, troughs and draw-bells completed later.
A range of criteria have been examined to compare and evaluate the four sequences. Result summaries are presented in Table 6, which lists peak values of stress, sidewall movement, rock mass strain and proportion of pillar area damaged for each of the four comparative cases. Note that movements are from specific sidewall points along the central production level tunnel in each model. Reducing the proportion of development ahead of the undercut progressively reduces the peak level of stress encountered in tunnel sidewalls from 87 to 47 MPa. Note that if tunnels were developed further behind the undercut, in the case where no development is done ahead, then even lower levels of peak stress could be achieved. The pillars between tunnels, in the area beneath the undercut, appear to be less damaged when development is done after undercutting. The models show three diagonal bands of different types of damage, paralleling the general undercut direction. These comprise: 1.
active shear failure ahead of the undercut face;
2.
past shear failure in a band just behind the current undercut face; and
3.
tensile failure in a band far behind the undercut.
These three areas are representative of the stress regimes, at current and past mining steps, in each of these areas. A large increase in severity of damage around tunnels occurs once breakaways are cut from the main tunnels ahead of the undercut. Rock mass deformation on the production level elevation is dominated by an overall downwards movement due to stress loading ahead of the undercut, followed by uplift resulting from relaxation behind the undercut. However when the amount of development is extensive ahead of the undercut, significantly increased deformation on breakaway corners is indicated by modelling. Strain and movement provide useful indicators of severity of damage and are shown as a function of the area of the production level mined prior to undercutting in Figure 5. Severity of damage appears only marginally improved if all tunnels are developed behind the undercut. This is probably an incorrect
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76%failed (22% impr)
conclusion: in the models the main tunnels were excavated at a point beneath, and not well behind, the undercut face, and consequently were still sited in a moderate stress field. If the point of development had been moved further back by 10 to 15 m the level of damage would have been reduced. The development of the main production tunnels alone ahead of the undercut may be feasible (with up to 2 m of sidewall damage). However, no breakaways should be put in and it would be preferable to put in all production development behind the undercut.
EFFECT OF DRAWPOINT SPACING The basic Herringbone model geometry was compared with drawpoints on 15 m and 18 m by 18 m spacings at C-Cut elevation. Main production tunnels were moved from 30 m to 36 m apart. Two situations were compared: 1.
all production level development was carried out prior to undercut advance; and
2.
all development was post-undercut.
Key comparative parameters are listed in Table 7. Percentage differences are shown relative to the 15 m base case with all development pre-undercut. Increasing spacing shows a minimal improvement over the corresponding closer spacing case. The benefit of increasing draw point spacing is small compared to the benefit achieved by delaying development until after the undercut is advanced.
EFFECT OF INCREASING THE VERTICAL DISTANCE BETWEEN UNDERCUT AND PRODUCTION LEVEL Models examined the effect of increasing the vertical spacing between the undercut and production level by 3 m from 15 m (as used in all models described above) to 18 m at C-Cut elevation. Draw point spacing was kept at 15 m by 17 m spacing to permit direct comparison to the range of models examined The objective was to assess whether additional stability benefits could be obtained by opening up the vertical spacing. Key parameters from the model are listed in Table 7, from which there appears little difference between a 15 m and 18 m vertical spacing. Other conclusions include, first, while excavating development behind the undercut results in avoidance of high stresses, increasing the distance from undercut to production level results in an increase in stress applied to production tunnels, ie the de-stressing effect of the undercut is reduced. Although the table shows that the peak stress, at the point where tunnels are developed, is reduced
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FIG 5 - Comparison of typical level of damage around production level main tunnels versus proportion of the production level which is developed prior to undercut advance.
because of the increased distance from the peak stress concentration ahead of the undercut, the average stress across the production level is raised. Second, at 15 m vertical spacing between production level and undercut there is a very small potential for interaction of rock mass damage around production tunnels with the undercut. Note that the disturbed zone beneath the undercut is generally limited to some 8 m and at 15 m this potential for interaction appears to be fully removed. In general, 15 m to possibly 18 m would appear to be an optimum distance for separating undercut and production level. Outside this range damage is liable to increase.
SUPPORT PRESSURE REQUIREMENTS The effect that support pressure has on rock mass stability for C-Cut production level development was examined, when excavated both ahead of, and behind the undercut. The objective was to provide a further indication of optimal locations for excavating production tunnels relative to the undercut face and to examine possible support requirements. A ground-reaction curve methodology is used, where excavations are created and a support pressure is applied to the internal excavation walls and incrementally reduced to zero, with rock wall movements monitored. A graph of support pressure versus rock movement (or closure across the excavation) is referred to as a ground reaction curve. Critical levels of support pressure to prevent rock failure, or even uncontrollable collapse,
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can be identified. Differences between curves for various locations around an excavation, or areas subjected to different stress fields, are indicative of relative difference in excavation stability. A single FLAC3D model was been used, based on the Herringbone model. It was run as a single mining step with C-cut production development completely extracted, apart from drawbells, and the undercut advanced to a point midway across the model. A range of monitoring points were selected along the main production level tunnels, at various positions relative to the undercut face. Because all mining was done in one step it was possible to simultaneously examine the effect of developing production excavations at these various positions, creating a ground reaction curve for each point. A range in ground reaction curves is shown in Figure 6. Of note is any level of support pressure below which a tendency towards rapidly increasing deformation is apparent. Because FLAC3D is a continuum program, the rock mass does not fall apart and very large closures are unlikely to be seen, as would occur underground as the rock mass unravels on the excavation skin. However, the graphs provide a good indication of the level of pressure at which severe damage is likely to be initiated. This can then be used as a target support pressure. To aid in interpreting these graphs the field stress, due to the undercut geometry, in which the tunnels are excavated should be considered. Trends in rock mass behaviour in tunnels can be identified as a function of the stress level under which the tunnel is excavated and are indicated in Figure 6. The three types of
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FIG 6 - Ground reaction curves for various points along main production tunnels showing a change in behaviour from a position ahead of the undercut (pt 1), beneath the undercut (pt 2), part de-stressed (pts 4 and 5), to positions that are completely de-stressed (pts 6 and 7).
behaviour are identifiable, as listed below, with field stress levels indicated, and are shown diagrammatically in Figure 7: 1.
High stress conditions (> 30 MPa) ahead of or directly below the undercut face. Squeezing ground, requiring yieldable support. Deformation increases when support pressure falls below 6 MPa. This level of support pressure is impractical to apply, and any tunnel developed here will require substantial support.
2.
Moderate stress (20 to 30 MPa), where the production tunnels are effectively less than 10 m in plan from the edge of the undercut. Raised stress here is due to the 15 m lead-lag situation between undercut rings in adjacent undercut tunnels. Deformation increases when support pressure falls below 2 to 3 MPa. This level of support pressure is probably impractical to apply. Conventional support will probably serve temporarily with concrete erected once undercut is advanced, possibly feasible with massive concrete.
3.
Lower stress (< 20 MPa), where production tunnels are outside the effect of the undercut lead-lag, and are more than 10 m in plan behind any point on the undercut face (effectively 20 m along the line of the production tunnel). Deformation rate increases at pressures below 0.1 to 0.5 MPa.
To provide a rough indication of the level of support required, consider that, in theory, a standard grouted rockbolt, yields typically at approximately 10 tons, or 100 kN. If installed on a 1 m spaced pattern the rockbolts would provide a support pressure of 100 kN/m2 (100 kPa, or 0.1 MPa). Addition of extra cables, a
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thin layer of shotcrete, etc is unlikely to provide much more than an additional 0.2 MPa support pressure. On the basis of the ground reaction curves, the conclusion would tend to be that development should ideally be carried out in those areas where stresses due to the undercut are reduced to below 20 MPa. If support for the higher stress areas is based on the dead-weight of loose ground, results in section 3, above, showed that typically a 2 m thick envelope of damaged rock around tunnels might have to be supported, requiring a support pressure of 50 to 100 kN/m2. However, if a system with this capacity, which is considerably less than the limiting 2 to 6 MPa indicated in the curves, was used it would have to accommodate a potentially large irresistible deformation due to squeezing conditions under the raised levels of stress.
GENERAL CONCLUSIONS AND LAYOUT GUIDELINES The following general conclusions can be drawn from the numerical modelling exercise. Damage increases roughly in proportion to tunnel size. In general, a 4 m tunnel appears to be a reasonable compromise, requiring up to 65 kN/m2 support pressure to bear the dead-weight of potentially loose ground, even in weaker ground. This is a practical, achievable, support pressure. Provided the percentage of ground extracted in the form of production level tunnels is identical in both cases, a Herringbone layout is slightly preferred to an El Teniente draw point layout on geotechnical grounds. Damage is slightly less within the Herringbone pillars between production level draw point tunnels,
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FIG 7 - Schematic diagram showing regions of rock mass behaviour relative to the undercut position, defined on the basis of ground reaction curves. These regions provide an indication of optimal positions for excavating C-Cut development relative to the undercut face.
and pillars should be marginally stronger, in part because the effective width of pillars is greater, Tunnel wall movement is approximately eight per cent less in the case of the Herringbone layout. There is a greater potential for instability over El Teniente layout breakaways. Damage to production development will be potentially severe if excavated prior to advancing the undercut. The best sequence is when all tunnels are developed behind the undercut. If however, the production tunnels are excavated at a point directly beneath, and not well behind, the undercut face, they are still sited in a moderate stress field. The development of the main production tunnels alone ahead of the undercut may be feasible (with up to 2 m of sidewall damage). However, no breakaways should be put in as they appear to trigger a rapid increase in damage severity. The optimal excavation sequence is when all production tunnels are developed at least 10 m in plan behind the undercut, where the induced peak field stress is less than 20 MPa and a maximum support pressure of 100 to 500 kN/m2 is require to contain movement. Beneath or ahead of the undercut, the induced field stress exceeds 20 MPa, and up to 2 to 6 MPa support pressure could be required to limit movement in kimberlite under potentially squeezing conditions. The effect of increasing the spacing between drawpoints in the Herringbone layout, from a modelled 15 by 17 m spacing to an 18 m spacing, appears to result in minimal improvement in rock mass conditions around production tunnels. At best an approximate eight per cent improvement in conditions is anticipated, and only if tunnels are developed ahead of the
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undercut. This level of improvement is probably not sufficient to warrant an increase in spacing on geotechnical grounds, and may compromise satisfactory draw of caved ground. However reduced development metres and cost, for example, could make it worthwhile. The models generally indicate that the C-Cut undercut will cause damage in the 8 m of rock mass directly under it. From 15 m to 18 m would appear to be an optimum distance for separating undercut and production level. At a distance less than 15 m there appears to be some tendency for damage zones to coalesce between undercut and production level. There is a risk of an increased level of rock damage due to the undercut. At distances more than 18 m there will be a progressive reduction in the de-stressing effect of the undercut, and general raising of the level of stress field. Note that this conclusion only holds if all production level development is carried out behind the advancing undercut.
REFERENCES Hoek, E, 1990. Estimating Mohr-Coulomb Friction and Cohesion values from the Hoek-Brown Failure Criterion, Int J Rock Mech Min Sci & Geomech Abstr, 27(3):227-229. Hoek, E, Kaiser, P K and Bawden, W F, 1995. Support of underground excavations in hard rock, (A A Balkema: Rotterdam). Itasca Consulting Group, 1997. FLAC3D Fast Lagrangian Analysis of Continua in 3 Dimensions. Version 2.0.
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Noranda’s Approach to Evaluating a Competent Deposit for Caving S Nickson1, A Coulson2 and J Hussey3 ABSTRACT
INTRODUCTION
Noranda Inc is a Canadian mining and metallurgical company with its activities managed through businesses in copper and recycling, zinc, aluminum, magnesium and various metals through its interest in Falconbridge Ltd. Noranda operates three base metal mines, three copper smelters, one copper refinery, one zinc refinery, one lead smelter and refinery, three copper recycling and sampling facilities seven aluminum producing facilities and one magnesium production facility in the construction stage. The mining side of Noranda has a strong history of open stope, cut and fill and shrinkage mining. The 1999 Noranda Annual Report lists a 206 million tonne resource located at Porphyry Mountain in Gaspé, Québec. Mont Porphyre is a large tonnage, low-grade porphyry style Cu-Mo deposit that was discovered in 1994. It is located at a depth between 1.0 and 1.7 km below surface and characterised by a very competent rock mass. The Mont Porphyre resource is located adjacent to the town of Murdochville, close to Noranda’s Mines Gaspé operations. Mines Gaspé closed its open pit and underground operations in 1999 after approximately 45 years of mining, but still operates a copper smelter near Murdochville. The low-grade and great depth of the Mont Porphyre resource suggested that mine design could not be based on the conventional Noranda open stoping mining practice. This paper will review Noranda’s approach to evaluating the Mont Porphyre deposit. The approach included two definition diamond drilling campaigns, a review of current caving practice, an in-house scope study, a geotechnical logging program for diamond drill core and participation in an international study on cave mining. One interesting component of Noranda’s approach was the scope study that involved a team of multi-disciplinary personnel drawn together for a short period of time. The objective of the scope study was to determine what conditions would be necessary to mine the deposit and to assess the merits of continuing with further exploration work. A review of current worldwide underground bulk mining practices, and in particular an evaluation of the applicability of caving mining methods, were key components of the scope study. The scope study team concluded that derivatives of caving could be applied to the mining of Mont Porphyre, but additional diamond drilling was recommended to further define the deposit and improve the level of geotechnical information. Following the recommendations of the scope study team, a diamond drilling campaign was completed during the summer and fall of 1997. The 1997 drilling campaign had two secondary objectives, in addition to the primary objective of defining the limits of the orebody. The first objective was to improve the level of geotechnical information available for the characterisation of primary block cave fragmentation. The second objective was to improve the delineation of a lower grade intrusive core. A geomechanical investigation was conducted and included particular emphasis on geotechnical data collection and analysis. A review of the actual data collection process, including several trials of different core orientation techniques, will be included in the discussion.
Mines Gaspé is located adjacent to the town of Murdochville in the Province of Québec, Canada. The area is within the Shock Mountains in the north central part of the Gaspé Peninsula (Figure 1), at an elevation of 575 m above sea level. Highway 198 connects Murdochville to the coastal village of Anse Pleureuse to the north (40 km) and to the town of Gaspé to the east (95 km).
1.
Noranda Technology Centre, 240 Hymus Blvd, Pointe-Claire, Quebec, Canada. E-mail: [email protected]
2.
Noranda Inc, BCE Place, 181 Bay Street, Suite 4100, Toronto Ontario M5J 2T3, Canada. E-mail: [email protected]
3.
Compañia Minera Antamina, Av La Floresta 497 440 Piso, Urb Chacarilla Lima Peru 41. E-mail: [email protected]
* This paper was compiled and edited from previous presentations at the 100th and 101st Annual General Meeting of CIM (Montreal 1998 amd Calgary 1999). Reproduced with permission by the Canadian Institute of Mining, Metallurgy and Petroleum and the Society for Mining, Metallurgy and Exploration, Inc.
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FIG 1 - Location map that also outlines the four tectonostratigraphic assemblages of the Gaspe Peninsula.
The porphyry/skarn copper orebodies at Mines Gaspé were discovered in 1938. Mines Gaspé commenced production in 1955 from the Needle Mountain open pit and produced over 141 million tonnes of copper ore at an average grade of 0.9 per cent Cu from two open pits and eight underground mining areas (A, B, C, D, E32, E29, E34, E38). Mining operations at Mines Gaspé ceased in late-1999. Open pit mining included the Copper Mountain (porphyry Cu) and Needle Mountain (A-zone: skarn/stockwork) open pit mines (Figures 2 and 3). Several underground skarn orebodies were originally mined by room and pillar, and more recently by open stope methods (Figure 3: Zones B, C and E). A total of 1.1 million tonnes of copper metal were produced from Mines Gaspé mine production and another 1.2 million tonnes from local and international concentrates (Hussey and Bernard, 1998). In 1994, a porphyry/skarn copper-molybdenum deposit was discovered at a depth between 1000 and 1700 metres below Mont Porphyre. By 1996, approximately 19 500 metres of drilling in 12 diamond drill holes had outlined 200 million tonnes grading 0.73 per cent Cu and 0.08 per cent Mo (Hussey and Bernard, 1998). The deposit was located between 1000 and 1700 metres below the top of Mont Porphyre. Figure 3 illustrates a longitudinal view through the Mines Gaspé area and the various porphyry/skarn copper orebodies. The stratigraphy is composed of bedded sediments that are gently dipping between 20° and 30° to the north. The stratigraphic sequence, from the bottom up, consists of limestones, siltstones and mudstones of the Forillon, Shiphead and Indian Cove Formations that are overlain by sandstones,
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FIG 2 - Plan view of a projection of the L2 skarn horizon, indication location of Mont Porphyre, with respect to the E - zone orebodies.
siltstones and mudstones of the York Lake and York River formations. The Forillon and Shiphead Formations (Unit #3 in Figure 3) host the Mines Gaspé orebodies and consist mainly of mudstones with two limestone units (L1 and L2). Regional scale normal (70°) and tensional (340°) faults acted as conduits for the Copper Mountain and Porphyry Mountain intrusives. Initial mining operations commenced from the Needle Mountain open pit and subsequently advanced underground into the B, C and E skarn orebodies using room and pillar and open stoping methods. The Copper Mountain open pit was used to extract the top section of the Copper Mountain intrusive. The final underground mining was located in the deeper E Zone area. The Mont Porphyre deposit is located to the north of the Mines Gaspé underground mining areas.
The alteration aureole that hosts the Mines Gaspé deposits surrounds the Copper Mountain and Porphyry Mountain quartz monzonite intrusives and is illustrated in Figure 2 at the L2 horizon. There is a skarn halo surrounding the intrusives that itself is surrounded by an outlying marble horizon and defines a total alteration halo of approximately 3 km × 3 km. The original limestones, mudstones and siltstones were metasomatised to potassic porcellanite and diopsidic porcellanite. Diopsidic porcellanite hosts approximately 70 per cent of the Mont Porphyre deposit tonnage with the remaining 30 per cent located within the intrusive unit itself. The reader is referenced to Hussey and Bernard (1998) for a more detailed discussion of the regional and mine geology. The Mont Porphyre deposit is outlined with respect to the town of Murdochville and the smelting facilities in Figure 2. The deposit outlines an area of approximately 600 by 450 metres that is composed of mineralised intrusive and surrounding porcellanite. Noranda commissioned an internal scope study in late 1996 to look at various mining options for Mont Porphyre. The scope study looked at various derivatives of block caving and sublevel caving. Cavability was challenged by the apparent competency of the Mont Porphyre orebody, but evidence suggested that current caving experience was moving to more competent rock mass conditions. The scope study concluded that additional diamond drilling would be useful for looking at subvertical fracture frequency and the possibility of associated mineralisation. It also suggested that diamond drilling would be useful to further define the extents of the lower grade intrusive. Based on the recommendations of the scope study team, a drilling campaign was completed during the summer and fall of 1997. The main objectives of the campaign were to improve on the geotechnical characterisation of subvertical jointing and the definition of the extent of mineralisation. Previous surface diamond drilling was largely subvertical in nature and therefore biased towards the definition of subhorizontal structure. One of the recommendations from the geotechnical portion of the Mont Porphyre Scope Study was to include some subhorizontal drilling and core orientation in any future diamond drilling.
FIG 3 - York Lake Fm. (1), Indian Cove Fm. (2) and Shiphead+Forillon Fm. (3). Mineable skarn/manto type mineralisation (A, B, C and E). The top of the deposit is approximately 1000 m below any surficial exploration detection limit.
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Work on the deposit has been limited since the completion of the 1997 drilling campaign, although some work has progressed in the area of bock size characterisation. Noranda has also funded the International Caving Study, an international sponsored block caving research project interested in pushing the limits of current experience. The underground operation at Mines Gaspé was closed near the end of 1999 and Mont Porphyre remains in Noranda’s list of resources. Each aspect of the history associated with Mont Porphyre will be reviewed in this paper. New developments in 2000 are refocusing activity on the research developments that would be key to making Mont Porphyre a mine.
porcellanite, but contains very little mineralisation. An important characteristic of the Forillon Formation, the formation that hosts the Mont Porphyre intrusive, is that its sediments are homogenous with very few bedding planes. This is quite different to the overlying Indian Cove and Shiphead formations that host the Copper Mountain ore deposit. The Forillon formation rocks were silicified by hydrothermal fluids, resulting in tight healing of the bedding planes. An additional T3 tuff is contained within the bulk of the orebody, just below the L2 horizon. However, this unit is only 8 m thick, and has been incorporated as acceptable geological dilution in the orebody.
Geomechanical characteristics SCOPE STUDY In late-1996, Noranda assembled a multidisciplinary team in order to perform a scope study to evaluate mining the Mont Porphyre resource through derivatives of block caving and sublevel caving mining methods. The main challenge was the very competent rock mass that exhibited a joint spacing in excess of one metre. It was recognised that experience with block caving under these competent rock conditions was very limited. The Scope Study team was drawn from several areas of Noranda and combined in some cases with external expertise. The team was lead by a project manager with significant operating mining experience. External resources were retained for the rapid development of project drawings as the scope study proceeded. On-site resources from Mines Gaspé were used for geological and mineral processing input. A project development office in Bathurst provided resources for geological reserve assessment and mine planning. The Noranda Technology Centre in Pointe-Claire, Québec provided some resources on the geotechnical and mining automation side. Several other in house experts were brought in from other Noranda operations to look at specialised topics such as hoisting and electrical power. A few external consultants were invited to review the project during the course of the Scope Study and present their recommendations. Overall, the team assignment was to complete the scope study in a three-month period and present the findings to Noranda with recommendations for future work. Elements of the Scope Study are explained in greater detail in this section.
Geology of the deposit It has been hypothesised that the formation of the Mont Porphyre deposit is not a classic porphyry deposit (Hussey and Bernard, 1998), in that during formation, most of the metalliferous fluids drained off up dip into the receptive L1 and L2 skarn horizons, preventing a large fluid/temperature convective cell. This has resulted not only in low-grade mineralisation of the porphyry (intrusive), but surrounding low-grade mineralisation of the country rock, the L1 and L2 skarn horizons and porcellanite units of the Forillon formation (Figures 2 and 3). The intrusive porphyry accounts for approximately 40 per cent of the ore deposit with the remaining 60 per cent being contained within the porcellanite P4, P5 and P6 units, the L1 and L2 skarn horizons (Figures 2 and 3). A base level of 1700 m below surface is thought to define a reasonable grade cut-off with the majority of mineralisation above this level. The top of the deposit is located at a depth of 1000 m below surface (Figure 3), and is overlain by the York River and Indian Cove formation (1, 2 and 2a in Figure 3). The porphyry intrusive is capped by the T1, tuff unit, 25 m in thickness which is very hard and void of mineralisation, but highly fractured in comparison to the surrounding rock mass. The dominant upper limit of mineralisation in the porcellanite is capped by the T2, tuff unit (50 m thick) that is similar in characteristics to the T1 unit. This is separated from the T1 unit by the W1, Wollastanite unit (50 m thick), that exhibits a level of similar competence to the
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The low-grade bulk tonnage of the deposit made it necessary to characterise the deposit in relation to parameters that are used for the design of block caving, and sublevel caving mining methods. Block caving is the lowest cost underground mining method currently used in the world. Cavability of a block cave orebody is normally determined by characterising the quality of the rock mass using the Mining Rock Mass Rating (MRMR) developed by Laubscher (1990). This rating is a percentage scale (from one to 100 per cent) that is similar to the CSIR RMR classification (Bieniawski, 1989), and incorporates factors for the intact rock strength (UCS), RQD, joint spacing or fracture frequency, joint orientation, weathering, mining induced stress and blast damage. The last three parameters are adjustments not used in the CSIR RMR system. In the collection and characterisation of rock mass quality information, the MRMR parameters were given strong consideration in conjunction with parameters normally collected within the Noranda Group for the modified (Q) system rock mass classification (Barton et al, 1974). The Q System is used regularly for open stope design with the Modified Stability Graph Method (Potvin et al, 1989).
In situ stress regime At Mines Gaspé a number of in situ stress measurement campaigns have been performed at varying depths within the existing mine; CSIR triaxial stress measurements by University of Laval (1983); NTC slotter campaign E-29, (1993), and biaxial door stopper tests performed by École Polytechnique for the E-38 (1996). From these measurements a stress gradient was derived, which when compared to stress measurements performed in the Canadian Shield (Arjang, 1991), indicate an above average gradient that corresponds to the upper limit of combined Canadian Shield measurements. Although, Mines Gaspé is not located within the Canadian shield, this was used as a reference to determine an outer limit, as no other information within the Gaspé peninsula was available. The simplified stress gradient (intercept through the origin), which was used for subsequent numerical modelling, is summarised in Table 1. The lowest mining level of the deposit was determined to be 3100L (1657 m below surface). At this depth it was expected that in situ principal stresses could approach 100 MPa with an approximate east-west orientation and a horizontal to vertical stress ratio of 2.2.
Mechanical properties A number of laboratory testing campaigns had been performed on Gaspé ore and waste over the years. It is important when performing any geomechanical evaluation of a new deposit to review all previous testing, so that the results can be put into perspective and potential similarities identified. A summary of some of the average mechanical properties for Mines Gaspé is presented in Table 2. The intact uniaxial compressive strength (UCS) of most of the different ore types present at Mines Gaspé is exceptionally strong. The testing of Mont Porphyre core indicates that the Porcellanite is marginally stronger than the
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TABLE 1 In situ stress gradient calculated based on stress measurements at Mines Gaspé. Gradient (MPa/m)
Gradient (MPa/ft)
Trend (Degrees)
Plunge (Degrees)
Stress on the 1657 m level (MPa)
Sigma 1
0.0598
0.0182
103
18
99.1
Sigma 2
0.0300
0.0091
212
45
49.7
Sigma 3
0.0167
0.0051
357
39
27.7
Sigma Vertical
0.0275
0.0084
0
90
45.6
TABLE 2 Summary of average mechanical properties from testing performed at Mines Gaspé. Density (g/cm3)
Rock type and location
UCS (MPa)
Tensile Strength (MPa)
Young’s Modulus E (GPa)
Poisson’s Ratio, υ
E-32 Waste (Porcellanite)
213
17
52
0.27
2.70
E-38 Waste (Porcellanite P4, HW)
195
-
-
-
2.87
E-38 Waste (Porcellanite P5, FW)
340
-
-
-
2.97
Mont Porphyre (Indian Cove, IC)
262
-
71
0.18
2.75
Mont Porphyre (Tuffs, T1, T2)
454
-
91
0.25
2.80
Mont Porphyre (Porcellanite, P4, P5, P6)
211
-
59
0.25
2.80
Mont Porphyre (Intrusive)
195
-
43
0.18
2.70
Intrusive, which compares well to previous Gaspé core testing. The maximum recorded UCS result was obtained from the tuffs at 619 MPa, which to the author’s record represents one of the strongest units in the world. Although this horizon was strong in nature, it was also the most fragmented of all the geological horizons investigated.
Rock mass characterisation The ore deposit in 1996 was outlined by 12 subvertical (>70°) drill diamond drill holes from surface. The objective of rock mass characterisation was to determine the degree of natural fracturing and develop a rock mass rating for each rock type, using the Modified Q’ system (Barton et al, 1974), CSIR RMR classification (Bieniawski, 1989) and MRMR Rating (Laubscher, 1990). The original geotechnical data collected from these holes included RQD, Fracture Frequency, Joint Alteration (Ja), small-scale roughness (Jr) and joint orientation with respect to the core axis. The logging campaign also had provision for 1.5 m of core to be saved for every 30 m drilled, in order to provide samples for mechanical testing before the core was split for assaying. To augment the drill core data in which it was difficult to clearly identify joint sets, some limited underground exposures which exhibited similar geological characteristics and previous Copper Mountain open pit mapping data was reviewed. Each diamond drill hole was analysed separately to determine a distribution of parameters across the ore deposit and along the holes. A global summary of averaged values for combined holes for each domain is presented in Table 3. The majority of the rock types at Mont Porphyre fall into the upper class II (Good rock) and verge on class I (very good rock) (Bieniawski, 1989; Laubscher, 1990). In the area of the anticipated block caving undercut at the 3034 level (1650 m depth), the average RMR (Laubscher, 1990) was determined to range from 75 to 91. Owing to the high stress environment and past experience of stress fracturing in deep stope backs at Mines Gaspé, a ten per cent reduction factor for stress was applied as an adjustment. This resulted in a range of MRMR values between
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67 and 82. The applied stress reduction factor at this stage is based solely on experience and will require further appraisal in the future.
Cavability of the ore deposit There are two major questions that need to be answered in order to determine the cavability of an orebody (Laubscher, 1995). First, is there sufficient area to undercut the orebody enough to cause sustained caving of the rock mass? Secondly, can the resulting fragmented rock mass be collected and handled with appropriate technology? An answer to the first question can be determined using the empirically developed Laubscher Caving Stability Graph illustrated in Figure 4. This is a ‘caving’ design graph that is similar in graphical representation to the Modified Stability Graph (Potvin et al, 1989) developed for ‘open stope’ design, but should not be confused as the design inputs and end results are different. The Laubscher Caving Stability Graph plots MRMR versus the hydraulic radius (HR). The caving zone represents a design region in which experience suggests that the complete and sustained caving of the ore column will occur, not just mass caving of stope walls as is defined in the Modified Stability Graph. From the preliminary analysis, Mont Porphyre’s MRMR rating ranges from 67 to 82 (Figure 4). Typical caving experience is related to the mining of large underground low-grade orebodies that have an MRMR rating mostly less than 50. Mont Porphyre is far more competent than any existing block caving operation. However, recent changes in mining methodology have highlighted the potential of these more competent orebodies, such as Palabora Mine (South Africa) and Northparkes Mine (Australia). Palabora has a MRMR ranging from 57 to 70 (Kear et al, 1996), and will be mined by a modified block caving method, while Northparkes has an MRMR of 35 to 55 (Dawson, 1995) and is Australia first block cave operation. The stability curve in Figure 4 illustrates the range of conditions for cavability at Mont Porphyre. Cave initiation could start somewhere between a hydraulic radius of 40 and 55 metres. This would represent a
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TABLE 3 Summary of averaged rock mass parameters Rock Type
LengthLogged
No. of Joints
FF/m
RQD %
Jn
Jr
Ja
Q'
RMR
IC
43
38
0.88
90
6
1.0
1.63
9
82
T1
45
54
1.24
67
9
1.0
1.21
6
70
W1
25.6
23
0.90
84
6
1.0
1.70
8
74
T2
125
267
2.13
61
9
1.1
1.56
4.5
70
L1
45
22
0.48
97
6
1.25
0.94
22
79
P4,P5,P6
1558
1232
0.81
98
6
1.34
0.94
21
79
IN
819
515
0.63
99
6
1.38
1.1
24
82
M R M R 100
STA BLE
90
MINING ROCK MASS RATING
80
TR AN SITION AL
DURNACOL B4 - PREMIER
IN C R EA S IN G U /C U T SIZ E
B3 - PREMIER
70 HENDERSON P1
60
URAD
50
C ONS T R U C TIO N
PALABORA
40
BA5 - PREMIER ANDINA 2nd PANEL BL6 - SHABANIE
30
KING MINE
NORTHPARKES
KING MINE
20
C AVIN G HR BIG "O"
CASSIER
10
P O R PH YR E F O O T PR INT / U /C U T EX TEN T H R =110 m
P O R PH Y RE C A V E IN IT IA TIO N
HR SKYDOME
0 0
10
20
30
40
50
60
70
80
90
H YD R A U LIC RA D IU S (m ) = A R EA /PE R IM E TER
S tab le M in e C ases T ran sitio n al C ases : M inor caving of back, potential to form stable arch C avin g C ases : P rogressive caving of cave back G aspe M R M R range
FIG 4 - Laubscher Caving Stability Graph (after Laubscher, 1990), indicationg the relative location and points of potential cave initiation for the Mont Porphyre orebody, compared to other operations around the world.
range of undercut areas from 160 m × 160 m to 220 m × 220 m. For illustrative purposes, these dimensions were compared to the size of two well known stadiums in Canada, the Olympic Stadium in Montréal and the Sky Dome in Toronto. This type of comparison was found to be a good way of illustrating the required undercut excavation sizes for management personnel. The total available hydraulic radius of the orebody footprint is 110 m (420 m × 420 m) and provides ample room for additional undercut expansion if required to initiate continuous caving of the ore column. The answer to the second question requires the determination of the size of ore blocks reporting to the draw points. This is a function of the primary fragmentation (related to the in situ discontinuity spacing) in the initial mining phase and the secondary fragmentation (related to mining induced stresses and comminution
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of caved material in the broken ore column) during maturity of the cave. It is currently anticipated that in the worst case 70 - 80 per cent secondary blasting (ore blocks >2 m3 which could not be handled with an eight yard scoop) will be required in the initial mining phase based on analysis of in situ FF distribution. The effects of stress fracturing have not been considered in much detail and is potentially an area of future research. Determination of secondary fragmentation was made with the use of an empirical fragmentation model and indicates that once the cave reaches maturity (cave height ~500 m), then secondary blasting could reduce to 30 - 40 per cent due to self comminution in the broken cave column (Gash, 1997). This is an area in which further investigation will be required to more comfortably predict secondary fragmentation of competent orebodies. The ability to
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handle this large sized material effectively on a continuous basis and maintain a smooth production cycle, is one of the technical challenges for future consideration. Experience with oversize of this nature is currently being experienced by a number of mining operations on an intermittent basis world wide.
1.
The 1300 L (1122 m depth) exploration level (Hussey and Bernard, 1998) would be used for ventilation and instrumentation, incorporating microseismic monitoring, extensometers, TDR cables and observation holes to monitor the cave advance.
Mine design
2.
The undercut level (3034 L - 1650 m depth) utilises an advanced undercut that would be developed before major development on the extraction level (Figure 6). The advanced undercut has been proposed to limit stress degradation of draw bells and development, which would be affected by high induced stress at the advancing abutment.
3.
The extraction level on the 3100 L (1670 m depth) would be instrumented with convergence stations and stress monitoring in the major apex pillars and abutments (Figure 5).
4.
The final level would be the 3200 L conveyor level at a depth of 1700 m below surface.
Although the block caving method has been the primary method presented in this paper, the scope study reviewed six mining methods in total: block caving (modified block caving), sublevel caving, modified sublevel caving, large blasthole induced caving, horizontal hole induced caving, and open stoping. Owing to the desire to maximise recovery and minimise cost, emphasis was mainly placed on caving methods, and a brief overview of the block cave option is presented here. The other methods provided for induced fragmentation of the orebody, either by conventional sublevel caving methods or more novel approaches. The Large Blasthole Induced Caving (LBIC) method provided for the application of large diameter blastholes from open pit applications, not to fully blast the ore, but rather to induce additional fracturing.
Mine layout An idealised representation of the block caving mining option is presented in Figure 5. In this option a total of four levels would be required to mine the deposit.
The extraction level haulages and the undercut level drill drives have been designed to be orientated parallel to the principal stress direction in order to reduce stress degradation. The mining sequence on the undercut level that is illustrated in Figure 6 has been proposed to start from the weakest and most highly fractured region. This was determined from analysis of drill hole information to be in the south end of the deposit. The undercut would advance perpendicular to the maximum principal stress direction to take advantage of stress degradation in the back of the cave.
Caved Ore Blasted Ore Undercut Drilling
Advanced U/C
Undercut 3034 L 50’
x pe A r ajo M
66’ ( 20 )m
Stress
Production Horizon 3100L
56’ (17 m)
112’ (34 m)
Trench Drive
Bell
Draw Point
Haulage
FIG 5 - Idealised 3D schematic of the block caving option, showing the location of the undercut and the extraction level and illustrating the advanced undercut concept.
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N
HR (m) 110 (442x442)
• hydraulic hoisting; • automation technologies for mucking, drilling and blasting,
60 (240x240) 55 (220x220)
high hang-ups and communication;
50 (200x200)
• methods of inducing rock fragmentation; • the effects of stress fracturing and healed structure in
40 (160x160) 30 (120x120) 20 (80 x 80)
reduction in capital expenditures and reduced freight charges. Although the base case return was below Noranda’s requirement for a 12 per cent return on equity, the opportunities provide a much brighter picture. Various technologies that could contribute to the success of Mont Porphyre were identified during the presentations. These included:
0
Scale
porcellanite on primary fragmentation; 150 m
FIG 6 - Plan view of the 3034L (undercut) showing the orientation of drill drifts and the expansion of the undercut. Also, indicated is the HR of the undercut at various stages.
• • • • •
block size simulation; core orientation; handling large oversize in the drawpoint; rapid round development; and cave monitoring technologies (microseismic monitoring and laser profiling).
Some additional consideration from the scope study will be presented and discussed in the next section.
Subsidence One of the major concerns related to block caving is the effect of subsidence on surface topography and infrastructure. Many case histories exist for near surface cave operations, but there are few studies for cave mining at the depths associated with Mont Porphyre. It is anticipated that the cave will go to surface based on traditionally low swell factors in the range of five per cent of the caved material. The low swell results from large mass movement of large blocks, which tend to act as a piston. As such it has been postulated that in competent rock masses cave angles will tend to be at steep angles between 75° to 90° (Nicholas, 1997; Laubscher, 1990). A subsidence model was developed in which a cave angle of 85° was used from the undercut level and projected to surface. Then it was assumed that a glory hole wall would form with a slope angle of 55° and an external relaxation angle of 45°. The depth of the glory hole was determined from calculation of the removal of the orebody using a five per cent swell factor. The projection of these expressions on the surface topography was used to identify potential areas of mass movement (tens of metres) and minor relaxation (tens of centimetres) within the glory hole and relaxation zone respectively. This type of subsidence model is supported by the subsidence failure at Henderson Mine (Stewart et al, 1981), in which after only 15 per cent draw of the column height, a steep walled glory hole formed on surface from an undercut at a depth of 1280 m.
Other aspects of the scope study The scope study results were presented at the end of the three-month period during an all day presentation and review with key Noranda personnel. The various team members made presentations on all aspects of the project work. In general it was felt that a derivative of caving was applicable to Mont Porphyre and it would likely have to include some way of inducing fragmentation in advance. Recommendations for additional diamond drilling as well as several options for accessing a potential underground exploration drive were included in the presentations. Business cases for each caving method were presented with a base case rate of return and several opportunities for increasing that return. Opportunities included a tax holiday, a compressed construction schedule, a reduction in operating cost, an increase in grade, an increase in recovery, a
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MINE DESIGN CONSIDERATIONS FROM THE SCOPE STUDY The success of caving mining methods depends highly on the primary and secondary fragmentation characteristics of the orebody. In caving terminology, primary fragmentation occurs at the cave front and is dependent on the nature of the rock fabric and associated stress induced fracturing. Secondary fragmentation refers to the size distribution that reports to the draw-bells as a result of comminution, breakage and attrition during drawing of the caved material. The characterisation and evaluation of a rock mass for caving scenarios relates to the evaluation of primary and secondary fragmentation. Production rates in a caving scenario rely heavily on the degree of secondary fragmentation, which in the case of Mont Porphyre was predicted to be 80 per cent greater than 2 m3 in the initial phases of the caving process. Secondary fragmentation prediction relies on an understanding of primary fragmentation that will result from the caving process in both the ore and the overlaying waste rock. The T2 tuffaceous unit, located above the L1 horizon, was identified as being more fractured in the scope study and therefore less competent than the ore units. The implication of this increased fracturing is a higher degree of primary fragmentation and increased probability of the T2 unit percolating down into the ore column as the draw progresses. The issue of surface subsidence and glory hole development requires a greater understanding of the Indian Cove rock unit overlaying the deposit. The timing and location of glory hole development are important considerations with any caving mining method. The 1997 scope study was based on diamond drill hole data obtained up to 1996 and included about 2700 metres of geotechnical core logging. Most of this information was extracted from within the ore horizon and very little applied to the overlying geology. The stratigraphy in the vicinity of the Mont Porphyre deposit can be determined from Figure 3, however the rock mass domains that were applied to rock mass characterisation at Mont Porphyre are given in Table 4. The order in Table 4 follows the general stratigraphic column descending from surface, with the exception of the intrusive unit that is listed last, but cross-cuts many horizons. The basic conclusion from this work was that the intrusive unit exhibited a fracture frequency of 0.6 fractures/metre and the porcellanite units were
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TABLE 4 Rock mass domains and associated geotechnical core logging data extracted for the scope study. Rock Unit
Description (Original lithology prior to alteration)
Logged Core (m)
RQD
Fracture Frequency (per metre)
York River (YR)
Sandstone/mudstone
0
n/a
n/a
Indian Cove (IC)
Calcareous mudstone with limestone nodules
43
90
0.88
Brown mudstone/tuff
44
67
1.24
First Tuff Unit (TI) First Porcellanite (P1)
Calcareous mudstone
12
87
1.35
Mudstone and siliceous limestone concretions
26
84
0.90
Calcareous mudstone
0
n/a
n/a
Second Tuff Unit (T2)
Brown mudstone/tuff + limestone nodules
125
61
2.13
Third Porcellanite (P3)
Calcareous mudstone
4
100
0.55
Wollastinite (W1) Second Porcellanite (P2)
First Limestone (L1) Fourth Porcellanite (P4)
Argillaceous limestone
45
97
0.48
Black calcareous mudstone and tuff
211
96
0.76 1.75
Second Limestone (L2)
Limestone
5
91
Fifth Porcellanite (P5)
Black calcareous mudstone
163
99
0.79
Tuffaceous mudstone
18
96
1.06
Sixth Porcellanite (P6)
Black calcareous mudstone
1153
99
0.83
Intrusive Unit (In)
Quartz monzonite intrusive
819
99
0.63
Third Tuff Unit (T3)
slightly higher at 0.8 fractures/metre. Geotechnical information outside these units was limited, however a particularly low RQD and high fracture frequency was noted in the tuff units. The tuff units, and particularly the thicker T2 horizon, were identified as exhibiting finer primary fragmentation and therefore were considered as a likely dilution source. Several observations related to geotechnical trends observed during the scope study are listed as follows: 1.
The scope study identified that limited information existed in terms of orientation of joint sets in any of the rock units, both within and overlaying the deposit. Assumptions for the scope study were drawn from unoriented diamond drill core data, underground mapping of existing development exposures in similar rock units, and previous mapping within the Copper Mountain open pit. These assumptions needed to be verified with oriented core or underground exposures.
2.
The analysis of fracture frequency information suggested that the porphyry intrusive was slightly less jointed than the porcellanite. This has a potential impact on the primary fragmentation characteristics and the viability of a caving mining method.
3.
Underground mapping at Mines Gaspé suggested that subvertical structure was more dominant than subhorizontal structure. All of the drill holes that have intersected the Mont Porphyre deposit were vertical to subvertical in nature and potentially biased towards the definition of subhorizontal structure. For this reason it was suggested that any further drilling include subhorizontal holes directed to intersect the heart of the deposit. This would assist in understanding the in situ jointing and associated primary fragmentation distribution.
4.
The fracture frequency analysis from holes within the porcellanite units suggested that there might be more pervasive jointing in the south portion of the deposit. There also appeared to be some variability with depth in the amount and intensity of jointing that suggested some lack of homogeneity within the deposit.
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5.
A limited amount geotechnical information was obtained from the drill core logging program that was completed for the scope study. Alternate assessment tools and more detailed information (such as logging in the overlaying rock units) was suggested for any future drill core logging program in order to complete a more rigorous geotechnical logging exercise.
The 1997 scope study suggested that caving or some form of induced caving was potentially applicable to the Mont Porphyre deposit. The expected block size for normal caving methods was predicted to be initially very large, and it would be necessary to efficiently handle large oversize at the drawbell. The term ‘bell blast caving’ was coined to signify that the ability to efficiently handle large oversize at the draw bell would be of paramount importance. Induced caving methods, such as sublevel caving and large diameter blasthole caving, were also included for consideration in the Mont Porphyre Scope Study. A stope and pillar type mining method was eliminated due to low ore recovery and the potential problems of such mining at depth. The challenge for any method was initially to understand the fragmentation characteristics at Mont Porphyre.
1997 DRILLING CAMPAIGN Based on the recommendations of the scope study team, a drilling campaign was completed during the summer and fall of 1997. The main objectives of the campaign were to improve on the geotechnical characterisation of subvertical jointing and the definition of the extent of mineralisation. Previous surface diamond drilling was largely subvertical in nature and therefore biased towards the definition of subhorizontal structure. One of the recommendations from the geotechnical portion of the Mont Porphyre Scope Study was to include some subhorizontal drilling and core orientation in any future diamond drilling. The 1997 drilling Mont Porphyre drilling campaign was planned to delineate ten targets at the caving undercut horizon and to obtain geotechnical information at subhorizontal inclinations. The campaign included the deepening of four existing holes, four new subvertical target holes, and five subhorizontal holes. Three diamond drills were engaged on-site
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during the campaign. Subhorizontal drilling was oriented in two major directions in order to pick up geotechnical information perpendicular to the regional 070° and 340° sets and verify mineralisation of fracture domains. Two of the subhorizontal holes were directionally drilled off an existing hole and the other three were collared at a low angle on surface. The drill program was not fixed in nature and instead was ‘results’ driven and modified as a better understanding of the directional drilling capabilities was attained. Detailed records of hole deviation and drilling conditions (such as stabilisation, bit changes, etc) were kept during the drilling campaign.
Geotechnical logging program The 1997 drill campaign incorporated a revised geotechnical logging program that required the presence of one full-time person on-site who was particularly dedicated to geotechnical logging and core orientation. One component of interest in reference to geotechnical logging is the design implication of determining natural structure during logging for caving applications. In potential open stoping situations, difficulties in determining what jointing is natural and what is induced by the drilling process can sometimes lead to the incorporation of induced drilling breaks within the logged data. In an open stoping situation, this is absorbed into a slightly more conservative design approach as a result of additional structure.
In caving applications however, additional structure can be more favorable to the mining method and result in optimistic conclusions about cavability.
Geotechnical logging sheet The geotechnical logging sheet that is illustrated in Figure 7 was developed for use during the drilling campaign. A more detailed description of the components of the geotechnical logging sheet is given in this section. Each column on the logging sheet is discussed with a single descriptive paragraph. Box No. This column refers to the appropriate core box number which can be useful if it is necessary to pull a particular drill interval at some point in the future. A note is also made if this box contains core that has been saved for rock mechanics testing with an ‘RX’ inserted after the box number. The issue of saving samples for rock mechanics testing will be discussed further in Section 4.2. Core Interval. The logged core interval is noted (From — To) here in the same units that are used by the diamond drill crews on their logging markers. A standard drill interval of 3 m (10′) is suggested as this normally coincides with the drillers markers. Lithology. The lithology is noted from the geological drill logs. Geological logging was being done at the same time as the geotechnical logging in the case of Mont Porphyre and lithological information was extracted at a later date.
FIG 7 - Geotechnical core logging sheet.
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NF - Number of Natural Fractures. The total number of natural fractures in the core interval is listed in this column. Natural fractures can generally be identified as being relatively planar oblique fractures to the core axis, and will generally have some minor alteration of the joint surface. Fractures that do not completely break the core are not usually included. It is sometimes difficult to identify the difference between natural fractures and driller induced core breaks. For this reason, the breaks that are made at the end of the core run and at the end of each box row are not normally included in the count as they are most likely to be not natural. If the core is handled well then the number of artificial fractures can be kept to a minimum. AF - Number of All Fractures (Natural and Artificial Breaks). This column represents the total number of artificial and natural fractures summed over the core interval, without artificial breaks at the end of the core run or at marker locations. The main purpose of this information is to monitor the total number of fractures whether they are considered artificial or not. Artificial breaks can generally be identified as having very uneven fracture surfaces. When the uneven parts of the core are fitted together a perfect tight fit (match) can be made. Also, artificial breaks caused by core removal from the core barrel or poor handling, will not have any alteration or coating on the fracture surface. RQD - Rock Quality Designation. RQD is a measure of the percentage of the lengths of all core pieces greater than 100 mm (4″) long and has been standardised for NQ core. If the core interval is standardised to 3 m (10′) and the recovery is 100 per cent then the length of the interval measured will be 3 m. Artificial breaks are not included in the measurement of RQD. REC - Recovery of Core. Recovery is the percentage of the core that is recovered from the drill advance. It should be possible to take this directly from the driller’s log, however, it is always useful to check. The recovery can often be reduced from 100 per cent if a fault zone or vug is encountered. JC + JO - Joint Orientation and Joint Coating. This column includes the logging of orientation with respect to the core axis and joint coating or infilling. The information is logged for each natural joint. The joint orientation is the angle of the joint with respect to the axis of the core, such that a notation of ‘9’ represents a joint cutting perpendicular to the core axis. The joint coating or joint infilling follows the same categories that were defined for geological logging purposes and a selected list for Mont Porphyre is given as follows:
Physical characteristics
C = chlorite
Cp = chalcopyrite
Q = quartz
Py = pyrite
F = fluorite
Mo = molybdenite
Ca = carbonate
V = vide (void)
T = talc
G = fault gouge
The notation for each natural joint combines a numerical value that represents the joint orientation to the nearest tenth value combined with the appropriate joint coating descriptor. Each natural joint is logged separately on a different row. The entry ‘9Mo’ would represent a natural joint that is perpendicular to the core axis and exhibits a coating of molybdenite. T - Thickness of a fault or major infill (mm). If a fault is encountered or a joint with measurable infill (>2 mm), this should be noted in the thickness column. The fault should also be identified with a ‘G’ in the Joint Coating column. Jr - Joint Roughness. A range of joint roughness should be identified for the core interval. The definition of joint roughnesscan be determined from the examples given below. It is useful to identify the a range of values and circle the most dominant range. Definition
Feels to the touch
P = polished/smooth
-
SR= slightly rough
-
Like a babies bottom Like fine sand paper
R = rough
-
Like coarse sand paper
VR = very rough
-
Like coarse sand paper with small scale undulations (mm’s)
Ja - Joint Alteration. Joint alteration (Ja) from drill core is best determined from the physical characteristic of the joint surface. A simple guide is given below that is typically used with a knife and a fingernail to identify a range of Ja over the core interval. The most dominant value is normally circled if a range of Ja is given as shown below. Distances Between Natural Fractures. The distance between each natural fracture is noted in this column in order to determine a rough estimate of block size that can be related to primary fragmentation assessment. It is important that even if there is only one fracture in the core interval, the distance should
Observable characteristics
Ja
φ
r approx residual friction angle
Very hard cannot be scratch easily with a knife Tightly healed, hard, non-softening, impermeable filling
0.75
Can be scratched with a knife
Unaltered joint walls, surface staining, or hard infilling
1.0
25° - 35°
Can be gouged (deep scratch) with a knife
Unaltered joint walls, surface staining, or hard infilling
1.5
25° - 35°
Can be deeply gouged with a knife, scratched with a finger nail and feels slippery/soapy to the touch
Slightly altered joint walls, non-softening mineral coatings, sandy particles, clay-free disintegrated rock, etc.
2.0
25° - 30°
Silty-, or sandy-clay coatings, small clay-fraction (non-softening)
3.0
20° - 25°
Softening or low friction clay mineral coatings, ie kaolinite, mica. Also chlorite, talc, gypsum and graphite etc. (Discontinuous coatings, 1 - 2 mm or less in thickness)
4.0
8° - 16°
Can be deeply gouged with a knife and a finger nail, feels slippery/soapy to the touch and has a decent thickness (ie 2 - 4 mm)
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be measured to the next natural fracture (measuring and summing the lengths of the core pieces) even if this is in the next core interval. Every natural fracture should therefore end up with a distance associated with that fracture. The first value that appears on the sheet should be the distance from the start of the core box to the first natural fracture, after which only the distances between natural fractures are noted. If the recoveries are always 100 per cent then the sum of all of these distances should be equal to the length of core logged minus that distance of the last natural fracture to the end of the last core box. General Comments. There is no ‘General Comments’ column provided, but it is useful to add any comments or observations in the last column (‘Distance between natural fractures’) using as many rows as necessary. Particular items that are worth noting here are the presence of broken ground, faults, dykes or shear zones and any observations of core disking. The presence of ‘broken zones’ or ‘rubble zones’ would also be noted here. A ‘broken zone’ would be classified as a core interval with pieces that are less than one core diameter, with 50 per cent of the pieces greater than 50 mm (1/2″) in size. A ‘rubble zone’ would be classified as a core interval with pieces that are less than one core diameter, with 50 per cent of the pieces less than 50 mm (1/2″) in size.
show up in photographs and can be referenced at a later date if the core is not split. 4.
The order of this work is up to the individual involved, but it is often best to get the Number of Natural Fractures (NF), Number of All Fractures (AF), Distance between Natural Fractures, RQD, and Recovery (REC) out of the way first as these tasks are relatively straightforward. RQD measurements are easily done by using a tape measure and summing the lengths of core that are greater than 50 mm (4″). In good quality core, such as Mont Porphyre, it may be easier to sum the lengths of core that are less than 50 mm (4″). The Joint Coating (JC), Thickness (T), Joint Alteration (Ja), and Joint Roughness (Jr) values are somewhat subjective and may vary through the interval being logged. A range of values can be given with the preferred range for the interval circled for emphasis. A preliminary assessment of these values can be made during the initial identification of natural fractures, but a second review of each natural fracture is often required in the assessment of these parameters.
Suggested logging procedure The procedure that is followed to collect the data identified on the Geotechnical Core Logging Sheet can be set up on site to meet the needs of the personnel involved. One suggestion for such a procedure is summarised in sequence as follows: 1.
Any observations of shears zones, faults, stress induced disking, or abnormal alteration zones should also be noted separately on the logging sheet, using the last column for these comments. 5.
Review geological logs to identify rock mass domains to be logged.
Lay out core boxes and note box numbers and hole depths involved. Core logging by its nature is best done with specific goals or lengths of core laid out in the core shack. In this way, geotechnical logging is aimed at a specific number of core boxes and completed before additional core is reviewed.
3.
Identify and mark all natural fractures with a highly visible (yellow) wax marker pen. This involves a review of all the core breaks and distinguishing between natural fractures and artificial fractures that are induced by the drilling and handling process. This can be somewhat subjective, but basically core breaks induced by the drilling and handling process can be identified by very rough irregular surfaces that are often not following any specific structure or horizon. In addition, the break surface can be appear somewhat fresh in nature and should exhibit minimal alteration. Normally, an ‘X’ with a circle, or some other appropriate symbol, is used to identify natural fractures. They should be marked with something highly visible and permanent so that they will
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Photograph all core and reference core boxes to a particular film and photo number. After the logging process is complete all core should be photographed as a permanent record before splitting. Photographs can include multiple core boxes, but an identifying tag should be included in the area to be photographed in order to note the hole number, core boxes and hole depths. Also it is useful to place a scale, such as a brightly colored ruler, in order to get a length reference within the photograph. Photographs are best taken from above and sometimes may require a series of overlapping photographs in order to get the area of interest. A log of the film number, picture number and core identification parameters (hole #, hole lengths, box #) should be kept. Mont Porphyre core was photographed with a digital camera and stored on CD-ROM. Try to ensure that the marks that identify natural fracture locations are facing upwards and visible in the photograph. When the films are developed, the photographs should be filed in an appropriate album.
Geotechnical core logging does not have to be done in all the diamond drill core obtained in a drilling program, however from a geotechnical viewpoint this is the ideal situation as the maximum amount of information is obtained for all potential rock mass domains. In the caving options that were under consideration for Mont Porphyre, the primary source of geotechnical information was the orebody since the issue of cavability was of paramount importance. Information within the overlaying rock units was treated as a secondary source of information, but was also important as the initial core logging campaign concentrated mainly on the ore horizon. The amount of core logged within each horizon can be modified to suit the time constraints involved as long as reasonable coverage is obtained within the different rock units. 2.
Log the information identified on the Geotechnical Core Logging Sheet.
6.
Identify samples to be saved for rock mechanics testing. In the rock units that will be split, it has been suggested to retain 1.5 m (5′) of every 30 m (100′) of core that will not be split and can be used for laboratory strength testing if necessary at a later date. These areas should be flagged in each core box so that they will be easily identified during the splitting process. Laboratory strength testing requires good quality core that does not have weak bands or structure that will affect the test results. Also individual samples that are roughly 2.5 times the diameter will be prepared from this core, so several samples can be obtained from core that is not broken up much. The location of the 1.5 m section may be moved around a bit in order to meet the above constraints.
Observations on diamond drilling It is recommended that whoever is doing geotechnical logging and orientation try to spend some time at the drill to observe
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conditions and procedures. It is also important to relate what is coming out of the core tube to what is seen in the core shack. During the Mont Porphyre drilling campaign a great deal of time was initially spent at the drill in order to observe core unloading procedures and get an idea of natural fractures and artificial breaks that result from fitting the core into the core box. In the intrusive unit, artificial breaks were identified by a clean, rough, generally fresh looking break with no alteration and were fairly easy to identify in the core shack. Certain varieties of porcellanite were found to have many healed structures that preferentially broke apart when fitting the core into the core box and often appeared natural. This was identified as a potential benefit to cavability. In terms of core removal procedures, two core springs rather than one were being used to break the core at the bottom of the hole. Core removal with two core springs was difficult and frequently accomplished by banging the end of the core against a metal trailer. This method had the potential of damaging core and a wooden block was substituted for the metal trailer in order to minimise core damage. Another method that also worked but was more time consuming, was to manually loosen the core spring tube segments and force the core back out into the core tube with a wooden stake. Occasionally a metal ring around the core tube was used to take the impact of a hammer and ease the removal of core. The use of a rubber hammer was suggested if this procedure was deemed absolutely necessary. Productivity is usually the defining parameter and drill core treatment is likely to vary significantly by crew and by day. Several visits were made to the drills just to observe and document procedure during the drilling campaign. Optimal crews and procedures were identified for purposes of critical drilling and orientation trials. Artificial breaks in the P4 rock unit, over a length of approximately 30 metres of logged core, were observed to be in excess of 60 per cent of all the fractures noted. Discussions on the importance of minimising artificial fractures were held with each drill crew and a procedure was established to have each drill helper mark natural breaks that come out of the core tube. This process was started initially with a marker pen but required special lumber crayons for use in wet conditions. The importance of minimising artificial breaks helped to ensure the correct recording of natural fractures and also to improve the ability to puzzle core together for orientation purposes.
Core orientation program A program of core orientation was included in the 1997 drilling campaign for Mont Porphyre. The initial phase of this program was designed to trial different orientation techniques and adopt a preferred procedure for the remainder of the drill campaign. Different techniques that were identified for trial are listed below. a.
Cole Directional Scribe Method
b.
Foster Technologies Acid Tube Method
c.
Noranda Technology Centre Borehole Camera
d.
Clay Imprint Core Orientor
e.
Orienting to Known Bedding
No underground access was available for Mont Porphyre and it was likely that the depth of the deposit would place heavy reliance on drill core information. There was underground access available in the porcellanite unit close to the existing Mines Gaspé operations, but none was available into the Mont Porphyre intrusive unit. The Copper Mountain open pit provides some access to a the Copper Mountain intrusive, but it is located near to surface and far from the Mont Porphyre intrusive.
Cole directional scribe method This orientation technique involves the scribing of three lines on
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the core as drilling progresses. One scribe line is referenced to an orientation lug that is attached to a Sperry-Sun Type B Multishot to reference top of the hole. The technique was tested on site, but was not successful due to induced rotation of the core tube while coring and the resulting variable spiraling and slipping indicated by the scribe line. The hardness of the rock at Mont Porphyre and the difficulty in reducing the core tube rotation were not favorable for this type of test. One drawback of this method is that the data processing phase is understood to be fairly extensive and requires expertise with Sperry Sun Multishot system.
Core Tech Canada acid tube method (Foster Test) A system was rented from Core Tech Canada (1988) that involved etching an acid tube that is keyed to a slotted core tube. The slotted tube will be referred to as the ‘Foster tube’ in this text and the overall procedure is noted as the ‘Foster test’. A diamond tip pen is used to scribe the core through the slots in the core tube after retrieval of the core tube. The acid test gives the orientation of the high side of the hole (low point on the acid tube) and can be used to offset the scribed reference line. A simple goniometer is used in conjunction with the surveyed hole location to orient structure in the hole. A series of three trials were conducted in order to examine the applicability of this method. The first trial evaluated the logistics involved and familiarised site personnel with the procedure and equipment. The trial involved setting up a core barrel assembly with the Core Tech equipment and matching the length required by the drill crew, coring a 1.2 m (4′) section of rock, leaving the acid to etch, retrieving the core barrel and then scribing a line on the retrieved core. The Foster tube will allow for approximately a 1.4 m (4.5′) cored section before the core would impact on a pin that holds the acid tube in place. The test was completed in three hours and 15 minutes at a depth of 1159 metres (3803′), but this included extra time to allow the acid to etch and organise various equipment requirements. The results of the trial suggested that a completed Foster Test would likely take about three hours in the 1200 metre (4000′) range. In order for the Foster test to be successful, the core must not rotate in the core tube before the line is scribed on the core after retrieval. The possibility of rotation was looked at in a second trial that involved the completion of two separate Foster tests that were separated by a cored length of 34 m (110′). The top of hole orientation was determined from the results of both Foster tests. A top of hole line was scribed on the core from the first Foster test to the second test by puzzling together each successive piece of core and using a special scriber prepared at NTC. The top of core scribe line matched for both Foster tests and suggested that core rotation was not occurring prior to breaking. A Sperry Sun Multishot survey by Cole Directional was incorporated into both Foster tests in order to monitor potential rotation. The Multishot results for the first test, using a film speed of 1 frame per minute, did not indicate any rotation. The second Foster test did not show any rotation until the very last frame before breaking occurred, where a 14° change was noted. The actual timing of the last frame before the break has a degree of uncertainty attached and likely occurred after the core break had occurred. The last minute uncertainty provided by the Sperry Sun results in the second trial lead to a third trial with the Foster test to increase the level of confidence with this orientation method. The third trial consisted of three Foster tests with a 4.9 metre (16′) cored interval between each test. The third Foster test was necessary since it was impossible to piece together a section of broken core in the first 4.9 m coring interval. The top of hole scribe line was again successfully matched between the second and third Foster tests. Assuming core can be successfully pieced together, which was very likely in all rock units except the tuff horizons in the case of Mont Porphyre, two separate Foster tests that line up would
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suggest a good test. Core can then be pieced together and scribed as far as possible on either side. A procedure for core orientation based on two successive tests that line up was developed and adopted for Mont Porphyre. The procedure relied on the ability to successfully piece together core between successive tests.
similar to the clay imprint technique except the core imprint is taken by a number of pins. It was felt that it would be difficult to interpret the imprint with the small number of pins usually incorporated into pin imprint orientation devices on the market. One group however had used a 12 pin orientor, which sounded a bit more promising.
The Noranda Technology Centre (NTC) borehole camera The Noranda Technology Centre (NTC) borehole camera can be used as a tool to provide information on joint orientation and rock mass condition. The camera is designed to fit in a 76 mm diameter borehole, but unfortunately can only be used to a maximum depth of 600 metres (2000′). It incorporates a head assembly that is designed to rotate 360° and provide the ability to locate and orient structure within a borehole. Open structure, stress induced spalling, fracturing, and joint orientation, location and condition can all be visually assessed with the camera. Borehole camera logging of structure within the Indian Cove unit at shallower depths was originally proposed as part of the core orientation program. This work was not completed during the drill campaign due to time constraints, however this work can still be completed at a later date if required.
Clay imprint core orientor The clay imprint core orientor (Call, Savely and Pakalnis, 1982) is weighted by a steel tube half filled with lead and is able to find gravity as it travels down the hole. It requires a drill hole that is inclined between 40° and 70° (Call 1993) so is not good for subvertical drilling. The Foster test was found to be readable up to 80°, at which point reading the acid tube became difficult. The method involves taking a clay imprint of the core left at the bottom of the hole and then matching the imprint with the top of the next core run. Several trials were conducted to evaluate the clay orientor procedure, and one test was successful at obtaining an imprint at a depth of 1768 m. The clay imprint method has a lot of potential as it reduces the time of an orientation test by approximately 60 per cent at a depth of 1200 m. Obtaining a good imprint however was found to be difficult in the deeper holes and the ‘feel’ of the driller was difficult to utilise. With deeper holes the pressure was found to build up quickly, and the imprint had to be taken quickly in order to avoid blowing a hose. Taking an imprint with new drill bits, when the diamond impregnated portion was at its thickest, seemed to be more difficult. The solution of adding more clay was tried in this situation, but became unstable and was knocked off. Although the clay imprint core orientor seems to have had reasonable success at shallower depths (Call, Savely and Pakalnis, 1982) the deep holes at Mont Porphyre presented some problems. Orientation imprints also provided an imprint of the inside of the drill bit, which was found to be useful in evaluating bit wear without pulling rods.
Orientation to known bedding The stratigraphy at Mines Gaspé can be related to over 100 tuffaceous marker horizons that have been identified. Diamond drilling over the years has identified the regional variation in dip and dip direction of these marker horizons. It was thought possible to orient core from the identification of a known marker horizon, given a rough idea of its dip and dip direction in a particular area. The key to the success of this method, is identification of the correct marker horizon and the ability to extract a dip and dip direction from the regional scale information.
Pin imprint orientor The use of a pin imprint orientor was considered and several contacts made to discuss the applicability of the method. It is
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Core orientation strategy The fracture frequencies that were logged during the scope study were used to provide an indication of the required amount of oriented core. It was assumed that at least 100 fractures per rock mass domain would have to be collected. Initial efforts were to concentrate on the porcellanite, intrusive and T2 (tuff) horizons. An orientation procedure was developed based on the Foster test trials and completed with all three crews on one drill to familiarise everyone with the procedure. The procedure incorporated two separate Foster tests that were completed one after the other. A successful test was indicated by the top of hole line for each test being within 10° of each other. Once a good top of hole scribe line was obtained, core was pieced together as far as possible and individual structures oriented to the scribe line. The success of this orientation procedure depended on the ability to piece core together between individual tests, and required that particular care was placed on core handling. Piecing together core in the T2 unit was found to be difficult and limited the amount of data collected from this horizon. Core orientation data was recorded on a separate logging sheet created for the purpose. Oriented core was initially saved from splitting and assaying until the procedure of orientation and geotechnical logging ran smoothly. Some lengths of oriented core were kept on a longer term basis in case additional information was required later. One Foster test as outlined in the orientation procedure was found to take approximately six hours at a depth of around 1200 metres. There was a significant amount of ‘wait time’ as the Foster tube was pumped down or retrieved by the wireline. Various operational problems were encountered with broken acid tubes, apparent slipped core on retrieval, and piecing together core between tests. Complete orientation tests frequently required several orientation runs in order to get subsequent tests that lined up correctly. The orientation program however did eventually provide over 400 oriented joints in the end and much of the credit for this goes to the diligence of several Mines Gaspé personnel involved over the course of the drill campaign. The orientation of natural fractures was the initial goal of the orientation program. Some of the porcellanite units were noted to have healed structures, which were considered as a potential source of weak break points during the caving process. Some orientation of these structures was completed for reference. Simple drop tests, or ‘belt buckle’ tests, were also used to evaluate the effect of healed structure. These tests involved dropping samples from belt buckle height and observing the core breakage after dropping. Drop tests in the P6 unit showed that breaks preferentially followed healed structure first, while tests in the P4 unit showed breaks crosscutting both intact rock and healed structure.
Geotechnical data processing Geotechnical core logging and orientation data was collected and entered into a spreadsheet for future manipulation. The data display and geostatistical tools available through the Gemcom software that was used at Mines Gaspé, was considered for the production of geotechnical sections associated geostatistical analysis. A compilation of the geotechnical logging data from the 1997 drill campaign is given in Table 5. The geotechnical logging program from the 1997 Drill Campaign did not improve the primary and secondary fragmentation predictions from the Scope Study. It did however
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TABLE 5 Rock mass domains and associated geotechnical core logging data from the 1997 drill campaign. Rock unit
Description (original lithology prior to alteration)
Logged Core (m)
RQD
Fracture Frequency (per metre)
York River (YR)
Sandstone/mudstone
1924
96
1.30
Indian Cove (IC)
Calcareous mudstone with limestone nodules
4391
97
1.30
Brown mudstone/tuff
131
89
1.80
First Tuff Unit (TI) First Porcellanite (P1) Wollastinite (W1)
Calcareous mudstone
85
96
1.20
Mudstone and siliceous limestone concretions
356
98
0.80
Second Porcellanite (P2) Second Tuff Unit (T2)
Calcareous mudstone
71
98
1.10
Brown mudstone/tuff + limestone nodules
456
87
2.20
Third Porcellanite (P3)
Calcareous mudstone
70
83
2.50
Argillaceous limestone
276
97
0.70
Black calcareous mudstone and tuff
1680
97
0.90
First Limestone (L1) Fourth Porcellanite (P4) Second Limestone (L2)
Limestone
98
96
0.60
Fifth Porcellanite (P5)
Black calcareous mudstone
314
98
0.70
Third Tuff Unit (T3)
Tuffaceous mudstone
30
94
1.40
Sixth Porcellanite (P6)
Black calcareous mudstone
2274
98
0.90
Intrusive Unit (In)
Quartz monzonite intrusive
1559
99
0.50
provide a much better database of structural data for future reference and identified opportunities in the areas of stress fracturing and healed structure. Block size assessment was one area that could be explored in more detail with the updated structural database and oriented core information.
Block size assessment from core In the geotechnical logging sheet discussed in the Section entitled: ‘Geotechnical Logging Sheet’, a reference was made to recording the distances between natural fractures. This information was collected in order to create distributions of block lengths, examples of which are illustrated in Figures 8 and 9. This information is collated from the ‘distances between natural fractures’ and manipulated to provide a distribution of potential block sizes assuming equal dimensions in all directions (Nicholas, 1997). It can also be used in order to extract fracture frequency and RQD information from core without having someone particularly log this information. Both fracture frequency and RQD could be determined from the this information since it is related to core depth, recovery and the notation of broken or rubble zones. The examples in Figures 8 and 9 show the difference in size distribution using this method for the intrusive and porcellanite rock units. The block length estimation for porcellanite suggested a primary fragmentation block size distribution that was 30 per cent finer than 1.2 m. The intrusive unit, which was understood to be more competent from the rock mass characterisation data, exhibited a block size distribution that was 12 per cent finer than 1.2 m. Both of these results were in close agreement with an empirical prediction using a software called Block Cave Fragmentation (BCF). BCF was developed by E. Esterhuizen at the University of Pretoria and is commonly used for caving fragmentation prediction. The 2 m3 cut off was used to relate to the maximum size that could be loaded in a 8 yd3 scooptram.
Oriented core data The oriented core data included 403 joints with the distribution illustrated in Table 6. The joints were logged as either natural
380
FIG 8 - Block length estimation from porcellanite core.
breaks (n), artificial breaks that looked natural (dn), or natural breaks that did not look natural (nn). The main reason for differentiating between natural and artificial breaks was to provide information on the susceptibility of a particular rock mass to induced breakage along healed or internal structure. It also provided some information on the frequency of artificial and natural breaks. Natural breaks that were observed during core unloading, but did not look natural, were logged separately and noted with a ‘nn’. They may be related to breaks that occur during the drilling or core retrieval process. Some of this data was collected from observations at the drill by geotechnical logging personnel, but the drill crew eventually marked natural breaks themselves as the core was unloaded. The latter procedure was not as accurate but was considered worthwhile. An exhaustive analysis of the oriented core data has not been fully completed at this stage and is an opportunity to explore further. Additional work on block size assessment is proceeding,
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campaign. The major plane data extracted from these stereonets is summarised in Table 7. This data suggests that there may be three sets in the porcellanite and intrusive units, both of which make up the orebody. The Indian Cove unit and all the data treated together exhibit only two joint sets.
CONCLUSIONS
FIG 9 - Block length estimation from intrusive core.
TABLE 6 Oriented core data. Rock Unit
Natural Break (n)
Natural Break (nn)
Artificial Break (dn)
Total
IC
34
16
4
54
T1
10
1
4
15
L1
12
1
13
26
P4
65
14
44
123
P5
19
0
0
19
P6
82
2
8
92
In Total
31
18
25
74
253 (63 %)
52 (13 %)
98 (24 %)
403
but at the moment is not first priority. The importance however was placed on collecting the information for future use as opportunities develop. Basic stereonets of the oriented core have been created by rock type in order to identify the potential number of joint sets in each rock unit. A basic premise of cavability in the scope study was that basic block formation would require three joint sets. Where less than three joint sets exist, cavability would rely heavily on stress induced structure or a mining method that incorporates some form of drilling and blasting to induce additional structure. Figure 10 illustrates the stereonet analysis performed for the most dominant oriented core data collected during the orientation
This paper has presented a discussion on the geotechnical investigation including geotechnical logging and orientation data of potential cave mining of the Mont Porphyre resource. The investigation to-date has been completed using conventional methods, but the use of geostatistical analysis for the compilation of geotechical data is also being reviewed. The advantages of using geostatistical methods are that the same tools that permit the visualisation of diamond drill hole geological and assay data could be used for geotechnical purposes. Sections can be produced that illustrate parameters such as rock mass rating or RQD and provide the ability to compile a true 3D representation. The geotechnical investigation of Mont Porphyre has revealed a very competent orebody that would normally eliminate any further consideration of a caving mining method. Current trends in practice and theory however suggest that there may indeed be a possible application of caving techniques to Mont Porphyre. The recent trend in caving experience is towards the exploitation of more competent ore deposits. Northparkes in Australia and Palabora in South Africa are two particular operations that are pushing current experience and have or will shortly initiate caving applications in competent rock. Noranda has joined the three year International Caving Study that has been jointly developed by the Julius Kruttschnitt Mineral Research Centre and Itasca Consulting Group. The main motivation for the project came from the growing worldwide interest in the application of caving mining methods to stronger massive low-grade deposits that lie outside current caving experience. The basic premise of world-wide caving expertise is that anything will cave if sufficient area is available for undercutting. The important issues then become the size of the broken ore and the ability to handle large size material at the drawpoint. It is believed that the benefits of reduced mining costs provide the incentive to further study the application of caving to competent orebodies. The identification of new technologies and methods that can be applied to traditional caving techniques has been a part of the work associated with Mont Porphyre. Methods of inducing and maintaining the cave using open pit blasting approaches and new technologies associated with the handling of large block sizes are two examples that have been considered. Subsurface diamond drill investigations at Mont Porphyre has included detailed geotechnical logging and a successful core orientation program. The associated data analysis discussed in this paper is far from exhaustive, but suggests that Mont
TABLE 7 Summary of major planes exracted from oriented core data. Data Set
No of Poles
Set #1 (Dip, Dip Dir, Strike*)
Set #2 (Dip, Dip Dir, Strike*)
Set #3 (Dip, Dip Dir, Strike*)
# of Sets
All Oriented Data
403
22/007/277
81/071/341
n/a
2
Natural Breaks Only
249
22/007/277
85/070/340
n/a
2
Indian Cove Unit
54
26/007/277
88/042/312
n/a
2
Intrusive Unit
74
16/063/333
87/061/331
46/136/046
2-3
P4, P5, P6 Units
233
10/132/42
75/089/359
52/028/298
2-3
*Strike is based on the right hand rule
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All Data
IC Unit
In Unit
P4, P5, P6 Units
FIG 10 - Stereonets for all oriented data, Indian Cove (IC) unit, intrusive (In) unit and the porcellanite (P4, P5, P6) units.
Porphyre will be a challenging application of cave mining methods. Opportunities for technology development were identified as part of the Scope Study. Efforts in core orientation and block size characterisation have since complemented the case for Mont Porphyre. The geotechnical logging program was designed to obtain as much data as possible and was particularly directed at caving type applications. Not all the collected data has been fully explored, but the information remains available for future use. Core orientation can provide valuable information, especially where underground access is not available. The experience with core orientation at Mont Porphyre has demonstrated that several orientation techniques can be incorporated into a geotechnical logging program. The geotechnical logging and orientation programs developed for Mont Porphyre can also be applied to the delineation of future deposits. This paper illustrates that there are many different core orientation techniques. One technique that could be used at Mont Porphyre for relative core orientation, is simply piecing together core in the core shack and orienting structure to a random scribe line. Although this technique does not provide true oriented core information, it can provide a reasonable representation of joint families if sufficient lengths of core can be puzzled together. In the case of Mont Porphyre, the general competence of the rock
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units allowed for long lengths of core to be assembled. Processing tools for dealing with large amounts of geotechnical information, such as the Mont Porphyre data, are not well developed. In the early-1990s, the Noranda Technology Centre had developed a psion based geotechnical logging software that was incorporated into underground geotechnical mapping work. This technology has since become dated and recent efforts are associated with utilising pen based systems and the Microsoft Access database product. Gemcom software has become widely used within Noranda for drill hole database information, reserve calculation, and geological modeling. The incorporation of geotechnical data processing is within the capabilities of a product like Gemcom, but is only now being aggressively pursued. An alternative is Gocad, which is a 3D geological modeling software that allows for sophisticated 3D visualisation of complex data sets and quantitative data integration. This tool is being explored further for the purposes of handling rock mass characterisation data. One effort that grew out of the Mont Porphyre geotechnical logging and orientation program is a project on fracture simulation. This effort is working on the development of a geostatistical method to assess the block size distribution of a deposit. Geotechnical logging of fracture frequency and core orientation data will serve as input to this technique. The work is currently in progress and has utilised the Mont Porphyre data for model formulation.
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ACKNOWLEDGEMENTS The authors greatly appreciated contributions during the scope study of the project team members and associates involved. The group was led by Bill Rogers and included Gaston Morin, Jacques Gagné, Pierre Bernard, Phil Gaultier, Claude Jacob, Robert Vaillancourt and J C Bélanger. We especially would like to thank the efforts of all the personnel who contributed to the geomechanical study including Paul Germain, J P Basque, David Gaudreau, Lionel Catalan, Luc St. Arnaud and Ken Liu from the Noranda Technology Centre; the Mine and Geology Departments of Mines Gaspé, especially Peter Marenghi and Harold Vachon; external consultants Peter Gash, Dave Nicholas, Maged Rizkalla and Laval University. During the 1997 drill campaign special thanks must be given to the Mike Cole, and in particular the field assistance from C Tremblay, E Stephenson, K Janssen, and V Vezeau. Visits to several caving operations, De Beers Premier Mine, Palabora Mining Company and the Northparkes Mine, were conducted as part of the work associated with Mont Porphyre. The assistance of various personnel at these operations and the welcome extended by the associated companies is greatly appreciated.
REFERENCES Arjang, B, 1991. Pre-mining stresses at some hard rock mines in the Canadian Shield, CIM Bulletin, January: p 80-85. Barton, N, Lien, R and Lunde, J, 1974. Engineering classification of rock masses for the design of tunnel support, Rock Mechanics, 6 May, pp 186-236. Bieniawski, Z T, 1989. Engineering rock mass classifications (Wiley: New York). Call, R, 1993. Clay imprint core orientor manual. (Call and Nicholas). Revised April 1993. Call, R D, Savely, J P and Pakalnis, R, 1982. A simple core orientation technique, in Proceedings Third International Conference on Stability in Surface Mining, 1-3 June 1981, Vancouver, (Ed: C O Brawner) pp 465-481 (Society of Mining Engineers of AIME: New York).
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Core Tech Canada, 1988. Manual of instruction for Core Tech Canada Diamond Drill Core Orientation System. Coulson, A, Nickson, S and Hussey, J, 1998. A geomechanical investigation of the application of caving mining methods to competent rock at the Gaspé Mont Porphyre deposit, Québec, Canada, presented at the 100th annual general meeting of CIM in Montréal, Québec from 3-7 May. Dawson, L R, 1995. Developing Australia’s first block caving operation at Northparkes Mines - Endeavour 26 Deposit, in Proceedings Sixth Underground Operators Conference, pp 155-164 (The Australasian Institute of Mining and Metallurgy: Melbourne). Gash, P, 1997. Review of block caving potential. Internal report to Noranda Mines and Exploration Inc, Mines Gaspé Division, Mont Porphyre Project. February 1997. Hussey, J and Bernard, P, 1998. Exploration of the Porphyry Mountain Cu-Mo deposit, Mining Engineering, 50(8):36-44. Kear, R M, Fenwick, F and Kirk, R L, 1996. The sizing of Palabora Underground Mine, SAIMM Colloquium: Massive Mining Methods, Johannesburg. Laubscher, D H, 1990. A Geomechanics classification system for the rating of rock mass in mine design, J S Afr Inst Min Metall, 90(10):257-273. Laubscher, D H, 1995. Cave Mining - The State-of-the-Art, in Proceedings Sixth Underground Operators Conference, pp 165-175 (The Australasian Institute of Mining and Metallurgy: Melbourne). Nicholas, D, 1997. Review of block caving and induced blast hole caving methods. Internal report to Noranda Mines and Exploration Inc, Mines Gaspé Division, Mont Porphyre Project. March 1997. Nicholas, D, 1997. Personal communication. Nickson, S, Hussey, J and Coulson, A, 1999. Rock mass characterization for block caving fragmentation assessment at Mont Porphyre, presented at the 101st annual general meeting of CIM in Calgary, Alberta from 2-5 May, 1999. Potvin, Y, Hudyma, M and Miller, H, 1989. Design guidelines for open stope support, CIM Bulletin, 82(926):53-62. Stewart, D, Rein, R and Firewick, D, 1981. Surface subsidence at the Henderson Mine, in Design and operation of caving and sublevel stoping mines, (Ed: D Stewart) pp 203-212 (Society of Mining Engineers of AIME: New York).
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The Past Focuses Support for the Future A D Wilson1 ABSTRACT The history and milestones in the evolution of support systems utilised in the mass-mined chrysotile mines of Shabanie and King are discussed in detail. The use of block caving, false footwall or sublevel caving mining methods brings with it some major support problems, which are exacerbated, in our low rock strength mining environments. The paper discusses the various support elements and systems that have been used successfully or otherwise in these operations. Some predictions for the future of rock reinforcing in our ground conditions are highlighted, in relation to geomechanics, levels of mechanisation and production risk analysis.
MINING ENVIRONMENT BACKGROUND The Shabanie and King operations in Zimbabwe (Figure 1) are both large, heavily mechanised chrysotile producers mining in relatively poor ground conditions. The King section is situated at Gaths Mine, 58 km NE of Shabanie Mine.
FIG 1 - Location plan of mines
FIG 2 - Generalised cross-section through Shabanie mine.
Shabanie - at Zvishavane The ore occurs as discrete, elongated and pod-like bodies in partly serpentinised dunite at the base of a differentiated ultramafic sill. The bodies are separated from each other by wide, sheared talc zones. 1.
Geological Consultant, African Associated Mines (Private) Limited, Chrysotile House, 93 Fife Street, Bulawayo, Zimbabwe. E-mail: [email protected]
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The ore-forming stress-controlled hydraulic fracturing together with residual thrusting and faulting has created an environment characterised by various degrees of rock competency. Populations of slips, joints, fibre seams as well as their spacing, orientation, continuity and condition play a major role. A carbonated dunite footwall alteration zone separates the ore zone from the massive to sheared talc footwall. The large range in rock competencies has required the adoption of the detailed geomechanics classification assessments using Laubscher’s MRMR classification system (Laubscher and Taylor, 1976; Laubscher, 1977; Laubscher, 1983). The applied geomechanics system plays a major role in determining support recommendations as well as aiding in safe siting of longer term installations. A major structural feature is the graphitic talc chlorite schist footwall shear zone, which is several metres wide and dips between 40° to 60° (Figure 2).
The more competent carbonated footwall has been used where possible to site key service development. The graphitic footwall shear zone poses a major support challenge and layouts tend to keep out of the shear. However, this is not always possible and then specific support designs are required for the area. Further detail on the characterisation of the Shabanie mining environment can be gleaned from numerous previous publications (Marano, 1983; Fomison, 1983; Wilson, 1983; Marano and Everitt,1987).
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Block caving, pre-break caving and sublevel caving at Shabanie mine The discrete ore zones have generally lent themselves to caving methods but with differing degrees of success depending on ‘local’ factors. The key impact elements in the large mining blocks here are:
• • • • •
hanging wall competencies (significant joint populations); the relative proximity to adjacent caves; the relative mining geometries; the location and shape of the undercut; and the rates and geometry of cave advance.
The problem of large ore block footprint (+/-150 m × 300 m) as well as significant ore heights (+ 100 m) means that it has been difficult to completely undercut an entire block rapidly ahead of the abutment stresses. Invariably, the mining has been restricted to phases within blocks and development in the periphery of the first phase has been exposed to significant damage. Major repair cycles are often required. Even with a major effort in terms of undercut and production level support, the second and third phases in Block 58 were subjected to severe production stoppage due to failure/closure of development. A major repair cycle was put into phase three with intermediate undercut development as well as reengineered support of the production level (Figure 3). This did enable almost complete extraction of the ore.
A significant cause in the damage in Phase 2 and further SE, was the exposure of incompletely supported development to the abutment stresses. In view of this damage, the block has been reconfigured to a sublevel cave to give more flexible control on the advance of the face shape. The time and stress-change factors at Shabanie are significant. The in situ MRMR may indicate relatively competent ground conditions during the core-drilling and initial development phase (reducing the motivation to support). However, unless there is recognition that the induced stresses will have a severe impact and need to be catered for in the face shape and MRMR support recommendations, major failures take place.
King (Gaths Mine) at Mashava The ore zone is in the footwall of an embayment of serpentinised peridotite bounded by steeply dipping metasediments and is separated from the footwall by partly sheared talc carbonate and more competent carbonated serpentine rocks. In view of the location of the host ultramafic in the greenstone belt, there has been considerable focus of tectonic stress. The resulting environment exhibits considerable shearing roughly parallel to the footwall geometry with several populations of low condition sympathetic and secondary joints. This environment has also been assessed carefully in terms of relative distribution of zones of different geomechanics classification and further detail can be obtained from previous authors (Marano, 1983; Wilson, 1983; Fomison, 1983).
FIG 3 - Plan showing Shabanie’s block 58 damage areas.
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Block caving mining methods at King Several authors have chronicled the progression from horizontal hand-worked grizzly drawpoints (first double sided then single-sided), through a reclamation horizontal herringbone LHD layout (Brumleve, 1987) to the inclined, multilevel LHD layout known as the False-Footwall (Fomison, 1983; Buchanan, Whewell, DeKock and Parshotham, 1987). Refer schematic cross-section (Figure 4). With increase in mining depth coupled with impact of the known inclined draw from the hill feature and progression of the mining faces into previously non undercut areas, the significance of abutment loading has increased. The resultant damage on the production and services development has dictated a major review of mining geometries, drawpoint spacing and requisite support. The importance of prior knowledge of in-situ rock conditions (RMR) together with the application of adjustments for the mining stress changes (MRMR) during the life of the production/services area cannot be over emphasised.
HISTORY OF SUPPORT SYSTEMS Shabanie and King have progressively transformed their mining and support methods to cope with increasing depth and deteriorating rock conditions. The mines have been in the forefront of developing innovative solutions to their often unique problems. The key support milestones in this evolution are commented on in Table 1.
FIG 4 - Schematic cross-section; King.
TABLE 1 Brief History of support systems.
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TABLE 1 (continued) Brief History of support systems.
THE FUTURE SUPPORT CHALLENGE The ground at both Shabanie and (more so) at King behaves in a typical plastic mode. Mining-induced high abutment stresses cause significant deformations, mainly in tunnel sidewalls. This is often accompanied by various degrees of floor lift. This is due to a quasi-elliptical failure zone, which spreads out horizontally from the sidewall and renders effective rockbolt anchorage difficult. This elliptical zone is effectively pressed out into the open development by the vertical abutment load, and, because there is little restraint against it, causes major damage/repair cycles. The impact increases with mining depths. There is also evidence of local inclination to the abutment stress at King due to the modifying effects of the slip-circle toe load exhibited by the caving hill feature. This has introduced an inclined draw column scenario, which has been confirmed by recovering a comprehensive series of markers from working drawpoints. However, development in this environment has the potential to degenerate after production begins such that the sidewalls bulge inwards and ultimately close off access. Depending on the MRMR of the area, floor-lift is also common. However, in most cases, the hanging wall remains as a stable arch and only cracks up/disaggregates when the sidewalls have reached an advanced stage of failure. A substantial proportion of King has very low MRMR conditions and hence support has to cater for this environment as a ‘norm’. Controlling the rate of failure to allow an uninterrupted production cycle is the challenge.
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DETAILS OF SUPPORT ELEMENTS USED AND THE RATIONALE BEHIND THEIR IMPLEMENTATION Rockbolts Pre mid-1980s, the fully cement grouted bolts had a forged head but these could not be used to tighten up additional integral support items. Now, both operations use 16 mm square twisted and threaded bolts with domed washers; this facilitates pulling mesh and/or straps tight against the sidewalls to limit initial movement. Bolt lengths are limited to three metres because of development size constraints (±3.2 m × 3.2 m). The bolts are ±17 - 19 ton UTS capacity.
Spiling bolts These are used successfully in prestabilising the immediate back of development in highly sheared ground that tends to flow (MRMR class 5) prior to blasting the next round. The bolts are installed as an overlapping fully grouted curtain which helps to interlock the small rock blocks sufficiently to allow the face and back to be shotcreted before the next round of spiling bolts and blast-holes is drilled.
Mesh Chain or diamond mesh currently 4 mm wire at 75 × 75 mm aperture used to:
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• add limited reinforcing to shotcrete (non-galvanised ); and • aid in spalling control when held in place with straps (galvanised). Although mesh does play a reinforcing role, that role is severely limited in environments where significant deformation is expected. The sourcing of heavier gauge mesh is being followed.
Shotcrete This is an essential starting element to support for King in the predominantly Class 4 to 5 MRMR rock conditions. In general, the layer is a minimum of 150 mm thick and is now ALWAYS mesh reinforced. The application of non-reinforced shotcrete in production areas within our ground conditions is purely cosmetic and can itself create dangerous conditions when it starts failing. This application has been described in detail elsewhere (Fomison, 1983; Wilson, 1991; Bell and Wilson, 1998).
FIG 5 - Typical tendon strap on sidewalls with interlocked heavy straps with cables in hanging wall.
Tensionable cablebolts The introduction of the 15.2 mm cables facilitated deeper anchorage to the in-rock reinforcing and enabled us to integrate this with the more effective surface restraint elements/straps. Key application points here are, for example, the wide-span turnouts of drawpoints off the collection drive. Neither Shabanie nor King truly tension these cables but simply pull up ancillary straps under tension on the fully grouted unit.
Tendon straps (and heavy straps) These are key units in conferring surface restraint to the rock between rockbolts and/or cablebolts. The design necessitates that the bolts act through the holes in the plates and thereby maximise the confinement potential from the steel working under tension. All that is required to achieve this co-incidence of bolts/holes is discipline in following the successful drilling and installation standards as laid out. When these straps are used in conjunction with mesh and shotcrete, any deformation of the sidewall immediately transfers load to the tendons. The support becomes active in limiting further movement. There are currently two different capacity units depending on application and matching current bolt/cable capacities. The tendon straps are best installed in a horizontal pattern (Figure 5); vertical units are not effective against shearing/ shortening of the sidewall. The tendon straps are currently being reengineered to incorporate a dome into the plates so that washers are no longer required on the rockbolts. Variations of this system have wide application and are fully described elsewhere (Wilson, 1991).
Rope trusses at bullnoses The successful use of old Koepe rope (32 mm up to 50 mm diameter) continues as a major deformation constraining element for both acute and obtuse turnouts as well as hanging wall and floor reinforcing. Early applications at bullnoses (Brumleve, 1983) used too much cable running along the surface which resulted in potential for lateral deformation (Figure 6). The main problem here is the inability to fully tension the ropes along the sidewalls prior to the grout setting. This geometry has been revised to reduce the surface acting dimension in favour of more units, each acting over a smaller area, thereby limiting deformation potential. In the case of bullnoses, more rope is grouted through the apex as opposed to around it; this restrains the apex from splitting open.
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FIG 6 - Surface ropes; too long but not damaged in this case; also note the ‘through-apex’ ropes at right.
‘Yielding’ TH arches and accessories Conventional 29 kg section TH arches used in a fully yielding role are not effective on their own because there is no rock-reinforcing content. The arches form a discontinuous surface skin, which is subject to intense point loads resulting in inward buckling and eventual inaccessibility. In order to maximise the support capacity by equitably distributing loads across several arches, we incorporate substantial channel-iron arch spacers (Figure 7 at ‘A’) together with reinforcing section panels, which are integrated into the concrete poured behind the arches. Hence, with this added stiffness, when the units come under load and deform, the enhanced system tends to give time for production/access. This is not an ideal support situation but has been adopted because:
• experience shows the system works acceptably (but limited); • the brow damage situation has improved from introducing a more coherent ‘unit’; and
• it can be improved by adding rock reinforcing elements before the arches are set.
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hanging wall), the shortening exhibits itself as buckling; usually towards the void. This, coupled with the inward ‘push’ from the deforming sidewalls can rapidly close off access.
FUTURE DIRECTION OF OUR SUPPORT EFFORT IN HIGH RISK AREAS Analysis of the past experiences and practical limitations within our environment suggests the following design requirements:
• Eliminate high abutment stresses in the mine design by innovative planning.
• Increase the depth of anchorage by using fully grouted cable or rope.
• Higher capacity anchorage should be intimately connected to enhanced surface restraint elements.
• Consider routine double TH arch legs slide in certain areas as an initial load take-up before major deformation occurs (Figure 7 at ‘B’).
• Increase in steel reinforcing potential for shotcrete to enable more effective restraint acting between the surface restraint elements.
• Focus of enhanced support effectiveness in the sidewall area which will in turn limit the progression of damage to the roof.
• Limit the use of TH arches to brows; favour the integrated new elements designed to give higher capacity and greater depth to rock reinforcing together with significant increase in surface restraint. This system will also be required to stabilise the floors.
FIG 7 - TH arches application showing arch spacer channels and clamps at ‘A’; double arch legs at ‘B’ to take up initial deformation.
NEW SUPPORT ELEMENTS IN EXPERIMENTAL INSTALLATIONS WHERE MAXIMUM SUPPORT IS REQUIRED SUCH AS POOR GROUND IN ABUTMENT AREAS Heavy mesh panels
KEY SHORTCOMINGS TO THE CURRENT SUPPORT SYSTEM(S) In low competency areas that require maximum support, the current support elements have the following limitations:
• Debonding of 3 m grouted rockbolts has, in many cases, allowed the entire sidewall to deform and move into the drift.
• Exposed sections of mesh reinforcing tear preferentially along the cracks when mesh reinforced shotcrete comes under load. Without reinforcing, the underlying shotcrete disaggregates further and the rockmass starts to unravel.
• Where cablebolts have been installed in sidewalls with barrels and wedges but not properly tensioned, the system fails with the cables pulling through the wedges and releasing any surface restraint. Properly tensioned cablebolts work effectively provided corrosion is not significant in the area.
• There is limited restraint capacity in the existing interlocking heavy duty strap system under severe stress conditions where the straps fail at ±19 - 20 tonnes. However, from experience underground, these units are the last to fail and offer a good solution alone in moderate to high load conditions or as enhancements in extreme cases. Density is the key.
• TH arches have definite limitations in that as the abutment loads increase (transferred from the downward closure of the
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These are made from 10 to 12 mm diameter +500 MPa steel wires woven at 120 to 150 mm square centres into panels that will replace the current wire mesh. The panels will be overlapped and be held in place by overlapping rock staples.
Overlapping rock staples These are custom designed support elements made from either cable or rope that will give the required depth of anchorage as well as act along the surface. The cable or rope is bent into the required angle and held by short lengths of steel piping, which will aid in stress distribution along the unit. The units are designed to overlap each other on the surface at a standard spacing. This is achieved by using drill director pipes attached to a lightweight stencil which ensures accurate hole spacing. These staples, together with the heavy mesh panels will confer much more restraint capacity to the sidewalls than has been available with our other support elements. Figure 8 shows an experimental area ready for the final shotcrete layer. In the relevant King environment where maximum support is required (WC/387, etc), staples made from redundant koepe rope are essential to give the correct restraint.
• Final shotcrete layer this is essential so that any deformation is transferred directly to load the steel in tension and will also protect the support from LHD damage.
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STAPLE MANUFACTURE
RESTRAINT CAPACITIES QUANTIFIED
A custom-made hydraulic rope + pipe bending unit (Figure 9), has been acquired for mass-production of the various configuration staples needed in high-risk support areas. This unit allows staples to be made at different bending angles depending on the support case. The most common will be both legs at 90° with some cases designed for one leg at 45° and one at 90°.
In order to counter the failures that have been seen, back analysis of the restraint capacities of the installed support has been done. The analysis has been based on assuming the installed steel elements will act in tension and confer a load-bearing restraint capacity acting in the tension direction. This has been quantified as a tonnes capacity per square metre of development surface. The capacity has been split into three active directions.
FIG 8 - Experimental cable staples with heavy mesh; awaiting final shotcrete.
FIG 9 - Staple-bending machine; The spacing between ‘bends’ as well as the bend angle can be varied.
FIG 10 - Sidewall view and section of original koepe staple support system; a proportion of ‘vertical’ units have since been replaced by more. Horizontal ones.
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In both ‘horizontal’ and ‘vertical’ above, the restraint capacity calculation is an average over a representative area; obviously, these must be intimately connected to the radial potential for success. Table 2 lists the restraint capacities of our most frequently used support systems and is the basis for the comparative discussions. What is evident from Table 2 is that minimal restraint capacities are offered by support codes 1 to 6 with only 5 and 6 offering substantial surface restraints from the arches and concrete-reinforcing units. The ubiquitous radial restraint from rockbolts is also considered inadequate and hence the figures for the staple option (code 7) is a quantum leap in restraint capacity in all directions (also refer Figure 12).
Radial That load-bearing potential of the aggregated steel elements acting against inward deformation, ie the highest effective UTS conveyed by the grouted bolts/cables/ropes within the square metre.
Horizontal That load-bearing potential conveyed by steel elements acting along and within the surface support layer in a ‘horizontal’ direction (parallel the development axis) which will limit the spreading apart of dislocated rock blocks under stressed conditions. For example, in a normal tendon strap system, the straps only exhibit tensile strength in the horizontal direction between their securing rockbolts (Figure 5). There is no vertical restraint capacity potential unless mesh is being used; in which case, the mesh figure will be accepted.
COMMENTS ON COSTS (Figure 11) VERSUS RESTRAINT CAPACITY (Figure 12) In assessing the data in Table 2 and the associated graphs Figure 11 and Figure 12, very high costs for the (imported) TH arch systems are not proportionately matched by increases in their restraint capacities. However, the significant advantage in restraint capacities (in vertical, horizontal and particularly radial directions), for the experimental Koepe hoist rope staple system comes with a cost that is less than half the TH systems. The anticipated benefits from the deep-anchored, higher capacity and well distributed restraint are that the system will not only be cost-effective but still leave room for refinement. This will come from enabling mechanisation and different materials within the constraints of our mining environment. The staple system is however, less stiff than the TH arch + concrete reinforcing systems and hence some deformation should be expected which will be confined by the enhanced quantity of steel working in tension.
Vertical That load-bearing potential conveyed by steel elements acting along and within the surface support layer in a ‘vertical’ direction (perpendicular to development axis) which will limit the spreading apart of dislocated rock blocks under stressed conditions. For best results, all steel elements conferring the restraint capacities should all be intimately linked so that the resultant constraint can act in at least two and preferably three dimensions.
TABLE 2 Data on support system relative costs and restraint capacities. Restraint capacity ‘##’ from steel in tonnes/m2 of face In-rock Prog $US/lin m
Radial
Horizontal
Vertical
3 metres long; spaced @ 0.95 m × 0 .6 m
91
91
87
nil
nil
4 mm thick; 375 MPa 75 × 75 mm apperture
37
128
nil
5
5 nil
Specifications
Rockbolts Mesh Shotcrete
Support code
&&
Acting along surface
Actual^^ $US/lin m
Elements
150 mm thick
1
92
220
nil
nil
Tendonstraps
1.2 m vertical spacing
2
29
157
87
15
5
Double tendon straps
0.6 m vertical spacing
3
59
187
87
27
5
Double tendon straps + shotcrete
0.6 m vertical spacing
4
59
279
87
27
5
0.5 m + spacers + reinforcing cages + backing concrete
5
1434
1621
87@@
35
45
0.5 m + spacers + reinforcing cages + backing concrete
6
2075
2354
87@@
35
45
95
95
nil
40
40
418
513
414
112
80
TH arches; 3 m × 3 m brows shear zones; TH arches; 3.5 m × 3.5 m in WC Heavy mesh panel
10 mm; 125 × 125 mm min 500 MPa
Koepe rope staples
per standard; Figure 10
7
Notes: @@ Radial capacity comes from initial bolting otherwise it would be ±20 t ## Only the highest number used ^^ Cost in $US/lin metre for the stated element on its own && Progressive cost of all elements that make up the support code
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FIG 11 - Graph showing relative costs of support systems.
FIG 12 - Graph showing relative restraint capacities of support systems.
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REFERENCES
TABLE 3 Explanation of support codes. Support Codes
Contribution Support Elements
1
Rockbolts + mesh + shotcrete
2
Rockbolts + mesh + t strap @ 1.2 m spacing
3
Rockbolts + mesh + t strap @ 0.6 m spacing
4
Rockbolts + mesh + t strap @ 0.6 m spacing + shotcrete
5
Rockbolts + mesh + shotcrete + reinforcing sections + 3.0 m × 3.0 m TH arches + shotcrete cover AT BROWS
6
Rockbolts + mesh + shotcrete + straps + reinforcing section + 3.5 m × 3.5 m TH arches + shotcrete cover in WC/387 shear zones
7
Rockbolts + heavy mesh + deep anchored koepe staples + shotcrete cover
Referred to in Figures 11 and 12.
SUPPORT TIMING It is widely acknowledged and confirmed by our experience that support must be installed before deformation takes place that will reduce the overall effectiveness of the system being installed. The experimental staple system also needs this consideration particularly for practical drilling of the deep holes.
ACKNOWLEDGEMENTS The author thanks the Directors of African Associated Mines (Pvt) Ltd for permission to publish this paper and recognises the contribution and assistance from the Mining Services staff at both mines for some of the data and figures used.
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Bell, N J W and Wilson, A D, 1998. Application of shotcrete at Shabanie and Gaths Mines, in School; Shotcrete and its application, (South African Institute of Mining and Metallurgy, Randberg). Brumleve, C B, 1987. Rock reinforcement of a caving block in variable ground conditions, King mine, Zimbabwe, in Proceedings Conference Africa Mining ‘87, pp 31-46 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Buchanan,G F, Whewell,B W, De Kock, H C G and Parshotam, C B, 1987. The ‘false –footwall’ mechanized caving method. King section, Gaths mine, Zimbabwe, in Proceedings Conference Africa Mining, pp 53-62 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Fomison, D W, 1983. Mining Operations at Shabanie and Mashaba Mines, in Proceedings Conference Mining and Metallurgical operations in Zimbabwe, Vol II pp 236-264 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Laubscher, D H, 1977. Geomechanics classification of jointed rock masses - mining applications, Trans IMM, Vol 86. Laubscher, D H, 1983. The design and effectiveness of support systems in different mining environments, in Proceedings Conference Mining and Metallurgical Operations in Zimbabwe, Vol II, pp 265-289 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Laubscher, D H and Taylor, H W, 1976. The importance of a geomechanics classification of jointed rock masses in mining operations, in Proceedings Symposium on Exploration for Rock Engineering, Johannesburg, pp 119-135. Marano, G, 1983. Rock mechanics requirements to define the Mining Environment, in Proceedings Conference Mining and Metallurgical Operations in Zimbabwe, Vol II pp 198-235 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Marano, G and Everitt, A P, 1987. Selection of mining method and equipment for block 58, Shabanie mine, Zimbabwe, in Proceedings Conference Africa Mining, ‘87, pp 229-238 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Wilson, A D, 1983. Geological and geomechanics requirements to define the Mining Environment, in Proceedings Conference Mining and Metallurgical Operations in Zimbabwe, Vol II pp 173-197 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare). Wilson, A D, 1991. Advances in cost-effective support technology for rock reinforcing in some deforming ground conditions, in Proceedings Conference Africa Mining ‘91, pp 31-46 (Institution of Mining and Metallurgy, Zimbabwe Section, Harare).
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Rock Mechanics as Applied in Philex Block Cave Operations R S Dolipas1 ABSTRACT The Sto Tomas II orebody of Philex Mining Corporation is a porphyry copper deposit located at Padcal, Tuba, Benguet, Philippines. It was mined through open pit method for the first two years starting 1958 and subsequently through underground method using the Block Cave System. The slusher type of draw was employed at the upper levels of the orebody from mid-year of 1959 up to 1994 and then shifted to LHD as the ore became more competent at the lower levels. Application of rock mechanics played a vital role during the conceptualisation of the LHD method of draw, the first of its kind in the Philippines. Rock mechanics is a combined responsibility of the geology and the mine engineering staffs, with the geology providing the vital information to the mine engineering to include in the mine planning the following: a.
the general geology of the orebody;
b.
the ground characterisation through rockmass classification;
c.
determination of the cavability of the ore column (this includes our experience on airblasts which occurred during the early stage of our last slusher blocks);
d.
undercutting sequence;
e.
fragmentation of the rockmass in the ore column;
f.
drawzone spacing;
g.
proper draw control procedure;
h.
lay-out of openings; and the
i.
design of ground supports based on the rockmass condition. For the proper interpretation of the rockmass that would cater to the diverse mining situations, Philex adopted the modified geomechanics rockmass classification (Mining Rockmass Rating, MRMR) as this system is recognised and used by most of the block cave mine worldwide. This paper aims to discuss and present the role of rock mechanics in the analysis and determination of geomechanical issues encountered in Philex mechanised block cave operation.
TABLE 1 Estimation of intact rock strength (after Laubscher 1990). Equivalent σc
Descriptive rock strength
MPa
psi
Easily molded in fingers; shows distinct heel marks
0.05
5
Molds in fingers with strong pressure; faint heel marks
0.07
10
Very difficult to mold in fingers; difficult to cut with hand spade
0.15
20
Cannot be molded in fingers; cannot be cut with hand spade and requires hand-picking to dig out
0.5
70
Very tough and difficult to move with hand pick; requires pneumatic spade for digging
0.7
100
Crumbles under firm blows with sharp end of geological pick and can be peeled off with a knife; too hard to cut out a test specimen by hand
3
450
Can just be scraped an peeled with a knife; indentations up to 3 mm show in the specimen with a firm blow of the geological pick point
7
1000
Cannot be scraped or peeled with a knife; hand-held specimen can be broken with one firm blow of hammer end of geological pick
20
3000
Hand-held specimen breaks under more than one blow with hammer end of geological pick
70
10 000
Many blows with geological pick required to break through intact specimen
200
30 000
INTRODUCTION Philex adopted the modified geomechanics rockmass rating classification of Dr D H Laubscher in 1994, the time when Philex was planning to go mechanised in their block cave operations. This mining rockmass rating (MRMR) classification system as introduced by D H Laubscher in 1974 was a development of the geomechanics rockmass classification system of Z T Beniawski to cater for diverse mining situations. The fundamental difference was the recognition that in situ rockmass ratings (RMR) had to be adjusted according to the mining environment so that the final ratings (MRMR) could be used for mine design. Basic parameters of the RMR are the following: a.
b.
1.
TABLE 2 Intact rock strength rating (after Laubscher 1990). Intact Rock Strength (IRS) Mpa
% Rating
185
20
165 - 185
18
145 - 164
16
125 - 144
14
105 - 124
12
Intact Rock Strength (IRS) which is the unconfined uniaxial compressive strength of the rock between fractures and joints. The intact rock strength of a rock can be estimated using Table 1 and the ratings for the different IRS values can be found in Table 2.
85 - 104
10
65 - 84
8
45 - 64
6
35 - 44
5
Fracture Frequency per Metre (FF/m) is the number of naturally occurring discontinuities in a specified length of core or sidewall (refer to Table 3).
25 - 34
4
12 - 24
3
Rock Mechanics Superintendent, Philex Mining Corporation, Padcal, Tuba, Benguet, Philippines.
MassMin 2000
5 - 11
2
1-4
1
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TABLE 3 Fracture frequency per metre rating (after Laubscher 1990). Fracture Frequency per metre (FF/m) Average per metre
Rating 1 Set
2 Sets
3 Sets
0.10
40
40
40
0.15
40
40
40
0.20
40
40
38
0.25
40
38
36
0.30
38
36
34
0.50
36
34
31
0.80
34
31
28
1.00
31
28
26
1.50
29
26
24
2.00
26
24
21
3.00
24
21
18
5.00
21
18
15
7.00
18
15
12
10.00
15
12
10
15.00
12
10
7
20.00
10
7
5
30.00
7
5
2
40.00
5
2
0
c.
Mining induced stresses and whether clamping stresses across blocks are increased or decreased (refer to Table 9).
d.
Blasting effects/excavation technique and depth of damage - Blasting creates new fractures and loosens the rock mass, causing movement on joints, so that the following adjustments on Table 10 should be applied. The 100 per cent adjustment for boring is based on no damage to the walls. It should be noted that poor blasting has its greatest effect on narrow pillars and closely spaced drifts owing to the limited amount of unaffected rock.
Laubscher (1983) showed that it is also possible to use the ratings (RMR) in the determination of empirical rockmass strength (RMS) and then in the application of the adjustments to arrive at the design rockmass strength (DRMS). This classification system is versatile and the rockmass rating (RMR), the mining rockmass rating (MRMR) and the design rockmass strength (DRMS) provides good guidelines for the purposes of mine design. However, in some cases a more detailed investigation may be required, in which case greater attention is paid to specific parameters of the system. Narrow and weak geological features that are continuous within and beyond the stope or pillar must be identified and rated separately. *MRMR = RMR × {(% Weathering)(% Joint orientation) (% Mining induced stresses)(% Blasting Effects)} *RMS =
(RMR − IRS Rating) 80 × Intact Rock Strength × 80 100
*DRMS = RMS × % Adjustments (blasting, weathering, and orientation) c.
Joint Condition which is the assessment of the frictional properties of the joints (not fractures) and is based on expression, surface properties, alteration zones, filling and water. The procedure for the determination of joint condition is shown in Table 4, which divides the joint assessment section into subsections A, B, C, D, and E. Subsection A caters for the large-scale expression of the feature, such as across a drift or in a pit face, B assesses the small-scale expression, C is applied only when there is a distinct difference between the hardness of the host rock and that of the joint wall, D covers the variations in joint filling, and E caters the cement hardness of cemented joints. The ratings of these three parameters are summed up to arrive at the RMR (refer to Table 5).
The adjustments to arrive at MRMR are the following as shown in Table 6: a.
Potential for weathering - certain types of rock weather readily, and this must be taken into consideration in decisions on the size of the opening and the support design. Weathering is time dependent, and influences the timing of support installation and the rate of mining (refer to Table 7).
b.
Joint orientation - The size, shape, and orientation of an excavation affects the behaviour of the rock mass. The attitudes of the joints, and whether or not the bases of blocks are exposed, have a significant bearing on the stability of the excavation, and the ratings must be adjusted accordingly. The magnitude of the adjustment depends on the attitude of the joints with respect to the vertical axis of the block. As gravity is the most significant force to be considered, the instability of the block depends on the number of joints that dip away from the vertical axis. This also dictates the direction of undercutting (refer to Table 8),
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The Design Rockmass Strength is used in the design of supports and can also be used in the calculation of the strength of a certain pillar as shown by the formula below, Stacey and Page (1986): Pillar Strength, MPa = DRMS × (W0.5/H0.7) where W = 4 × SI H = height of the pillar The evaluation of the following geomechanical issues shows the importance and usage of rock mechanics in the design of Philex mechanised block cave operations.
CAVABILITY Aside from the often asked question of, ‘Will it or will it not cave?’ the real question is, ‘Can we afford to make it cave, carry the rock away and extract the mineral?’ Laubscher (1994) stated that the cavability of an ore deposit is based on many aspects, but clearly, if a large enough area is undermined (hydraulic radius), any rockmass will cave. The manner of their caving and the resulting fragmentation size distribution need to be predicted if cave mining is to be successfully implemented. Controlling the draw as the cave can only propagate if there is space into which the rock can move can slow the rate of caving. Advancing the undercut more rapidly can increase the rate of caving but problems can arise if this allows an air gap to structures, heavy blasting and the influx of water can result in damaging air blasts. Rapid, uncontrolled caving can result in an early influx of waste dilution. The rate of undercutting should be controlled so that the deterioration of the undercut and extraction openings is avoided. Stagnant cave front should also be avoided since this will cause column or point loading on the apexes.
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TABLE 4 Joint condition rating (after Laubscher 1990). Parameter
Accumulative % adjustment of a possible rating of 40 Description
Dry
Adjustment (%) Moist
Moderate Pressure 25 - 125 li/min
High Pressure >125 li/min
Wavy/Multi-directional
100
100
95
90
Uni-directional
95
92
90
87
Curved
90
87
85
82
Slight Undulation
85
82
80
77
Straight
80
74
70
72
Very rough
95
90
85
80
Striated / rough
90
85
80
75
Smooth
85
80
75
70
Polished
80
75
70
65
C Joint Wall Alteration
Weaker than wall rock
75
70
65
60
D Joint Filling
Non softening and sheared material
Coarse
95
90
70
70
Medium
90
80
70
60
Fine
80
70
60
50
A Large Scale Joint Expression
B Small Scale Joint Expression
Soft sheared material (eg talc)
E Cement Hardness
Coarse
70
60
50
40
Medium
60
50
40
30
Fine
50
40
30
20
Gough Thickness < Amplitude of irregularities
40
30
20
10
Gough Thickness > Amplitude of irregularities
20
10
Flowing material
Anhydrite (5+)
95
Calcite, Fluorite (3-4)
90
Gypsum (1-2)
85
Example: A straight joint with a smooth surface and medium sheared talc under dry conditions gives A = 70 %, B = 65 %, D = 60 %, E = 85 %; total adjustments = 70 × 65 × 60 × 85 = 23 %, and the rating is 40 × 23 % = 9. The rock mass rating (RMR) is the sum of the individual ratings; RMR = IRS Rating + FF/m Rating + JC Rating
TABLE 5 Meaning of the ratings (after Laubscher 1990). 1 Class Rating Description
A
2 B
A
3 B
A
TABLE 6 Parameters/adjustments to arrive at MRMR (after Laubscher 1990).
4 B
A
5 B
A
B
Possible adjustment, %
Weathering
30 - 100
100 - 81
80 - 61
60 - 41
40 - 21
20 - 0
Joint Orientation
63 - 100
Very Good
Good
Fair
Poor
Very Poor
Induced Stresses
60 - 120
Blasting
80 - 100
An understanding of the characteristics of the orebody and its geological setting allows the subdivision of the total rockmass into groups or ‘Structural Domains’ each of which tends to behave similarly in response to engineering activities. Several lithologic units may be lumped together, while a single lithologic unit can be divided into multiple domains. Major fault structures often form their own domain, and the direction of engineering activity may alter rockmass behavior, resulting in different domains. Delineation of these structural domains then facilitates determination of the cavability of the ore deposit.
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Parameter
The most favourable rock structural condition or fracture geometry for caving is represented by a rockmass containing at least two prominent sub-vertical joint sets, and a subhorizontal set. At our present mining level (908 ML), two nearly orthogonal steeply dipping joint sets (N48°W - 72°SE and N51°E - 88°SW) and a subhorizontal set (N37°E - 18°SE) have been identified. Hydraulic radius is the plan area divided by the perimeter. For the same area, the hydraulic radius or stability index (SI) will vary depending on the relationship between the maximum and
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TABLE 7 Adjustments for weathering (after Laubscher 1990). Degree of weathering
Potential weathering and adjustments, % ½ year
1 year
2 years
3 years
≥ 4 years
Fresh
100
100
100
100
100
Slight
88
90
92
94
96
Moderate
82
84
86
88
90
High
70
72
74
76
78
Complete
54
56
58
60
62
Residual Soil
30
32
34
36
38
TABLE 8 Adjustments for joint orientation (after Laubscher 1990). No of joints defining the block
TABLE 10 Adjustments of blasting effect (after Laubscher 1990).
No of faces inclined away from the verticle
Technique
70%
Boring
75%
80%
85%
90%
Smooth-wall blasting
3
3
2
4
4
3
5
5
4
3
2
1
6
6
5
4
3
2.1
2
Adjustment, % 100 97
Good conventional blasting
94
Poor blasting
80
Note: An additional adjustment is applied when dealing with shear zones. The adjustments for the orientation of shear zones with respect to development are as follows: 00 - 15° 76 % 15 - 45° 84 % 45 - 75° 92 %
TABLE 9 Factors affecting mining-induces stress (after Laubscher 1990). • drift-induced stresses • interaction of closely spaced drifts • location of drifts or tunnels close to large stopes • abutment stresses, particularly with respect to the direction of advance and orientation of the field stresses • uplift • point loads from caved ground caused by poor fragmentation • removal of restraint to sidewalls and apices • increases in size of mining area causing changes in the geometry • massive wedge failures • influence of major structures not exposed in the excavation but creating the probability of high toe stresses or failures in the back of the stope • presence of intrusives that may retain high stress or shed stress into surrounding, more competent rock.
minimum spans. For example, a block having a dimension of 50 m × 50 m has the same area as 500 m × 5 m, but the SI of the first is 12.5 m whereas the SI of the second is only 2.5 m. As shown in Figure 1, the large 50 m × 50 m block is less stable than a 500 m × 5 m tunnel and this is well illustrated by the difference in the SI. Brady and Brown (1985) found that the state of stress in the block exerts a major influence on the initiation and development of caving. Mahtab and Dixon (1976) showed that even if rockmass structure and geomechanical properties are well
398
FIG 1 - Relationship between maximum drawzone spacing and fragmentation for different LHD operations (after Laubscher1994).
conditioned for slip, the existence of high confining stress (horizontal stress) can inhibit rockmass failure. Therefore, the magnitude of the horizontal stress should be determined as this dictates whether there is a need to delineate the block by weakening its boundaries through isolation raises, or any other means of boundary weakening. In our case at Padcal, our resulting MRMR for virgin blocks would range from 60 to 65, therefore, we need to have a block having a hydraulic radius of 30 m to 35 m or a block with dimensions of 100 m × 200 m. Caving of adjacent blocks is much easier considering that we are only increasing the
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ROCK MECHANICS AS APPLIED IN PHILEX BLOCK CAVE OPERATIONS
hydraulic radius of the previous cave. The important factor is to ascertain that a connection between the new block is made with the previously caved block. This will ensure that no pillars are left in between the old and new block at the undercut level. At the last slusher blocks at 908ML where air blasts were experienced, the hydraulic radius during the first air blast was only 26 m. Two more air blasts were experienced when all the blocks were undercut. Final hydraulic radius of the four blocks was 34 m but due to increased MRMR due to high horizontal stress, caving rate was slowed down due to the clamping of joints. Actual caving rate through back calculations is eight centimetres per day. After these experiences, it is now a standard operating procedure for virgin blocks to be driven with isolation raises at the block boundaries to cut the high horizontal stress.
FRAGMENTATION Primary fragmentation Caving results in primary fragmentation and as defined by Esterhuizen (1994) it is the particle distribution that separates from the cave back and enters the draw column. The primary fragmentation from stress caving is generally finer than from subsidence caving owing to the rapid propagation of caving in the latter case with disintegration of the rockmass, primarily along favourably oriented joint sets and little shearing of intact rock. The orientation of the cave front/back with respect to the joint sets and direction of principal stress can have a significant effect on primary fragmentation.
Table 12 shows the resulting primary and secondary fragmentation at our southwest LHD Blocks at 908 ML using the Block Cave Fragmentation Program of Dr D H Laubscher. From this summary, we can conclude that the resulting fragmentation in our ore column will be medium to coarse throughout the life of the block. Though finer materials are expected during the later stage of extraction, medium to coarse blocks are significantly present in the ore column.
DRAWZONE SPACING Drawzone spacing should always be determined using the results of the secondary fragmentation study and interaction of adjacent drawpoints. As Brady and Brown (1985) stated, the spacing should also be carefully chosen so that the three-dimensional flow ellipsoids from adjacent drawpoints overlap slightly. This produces almost complete extraction and minimises dilution. Based on the resulting secondary fragmentation using numerical modelling of our SW LHD Blocks, blocks greater than 2 m3 comprises ±25 per cent to 33 per cent of the ore column. Using these data on Figure 1, we can have a drawzone spacing of 21 metres to 24 metres across the major apex (Distance A in Figure 2). Increasing the draw zone spacing will result to lower capital cost due to lesser development drives and supports but then proper draw control should be practiced to ensure interaction of the draw points.
Secondary fragmentation Secondary fragmentation is the reduction in size of the original particle that enters the draw column, as it moves through the draw column. The processes to which particles are subjected determine the fragmentation size distribution that reports to the drawpoints, ie a strong, well-jointed material can result in a stable particle shape at a low draw height. A range in rockmass ratings will result in wide range in fragmentation size distribution as compared to the fragmentation size distribution produced by rock with a single rating as the fine material produced by the former tends to cushion the larger blocks and prevents further attrition of these blocks. Table 11 shows the relationship between the adjusted rating to the hydraulic radius, cavability, expected fragmentation, and frequency of blasting. The degree of fragmentation is controlled by the interplay of: • Original rock block size • Comminution effect • Strength of the rock
During caving, block size is continually reduced by the comminuting effects of rock mass movement. Additional breakage is influenced by the strengths and hardness of the rocks present, the distance of travel, the stress condition, or depth of burial.
• Tendency of large blocks to float down to lower levels
When rock is only slightly fractured or cemented together, large blocks can ‘float’ down virtually intact if surrounded by softer rock. The smaller, softer pieces do not have the ability to break the block, and cannot effectively transmit the shearing and differential forces that would ordinarily have broken the rock in the undercutting operation.
• Thickness and nature of fracture fillings • Fines likely to be formed by grinding or travelling in from elsewhere
The fracture intensity and block size, together with the strength of the rock and the nature and thickness of fracture filling materials, dictate the percentage of fines.
MassMin 2000
FIG 2 - The three critical distances in Philex block cave layout.
DRAW CONTROL To improve fragmentation as stated by Brady and Brown (1985), the height of cave should be maximised and a slow initial rate of draw should be used. The rate of draw should be such that the volume of ore removed during caving is equal to the volume increase or bulking of the caving rockmass. The sequence and rate of drawing ore from several drawpoints in a block should be selected so as to produce controlled development of the cave line. It is the common practice not to attempt to draw down a complete block uniformly but to retreat the cave line across the block at a controlled rate. The drawpoints removed from the developing cave line should be drawn uniformly to promote the even lowering of the caved ore. Funneling can occur if ore is drawn from a given drawpoint at an excessive rate, and if the cap rock is weaker than the ore. Likewise, column loading or point loading can occur if drawpoints are kept in isolation or not being drawn.
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TABLE 11 Relationship of MRMR with hydraulic radius, cavability, fragmentation and frequency of secondary blasting (after Laubscher 1990). Adjusted Class Area Undercut as Hydraulic Radius
1
2
3
4
5
Not Practical
30 m
30 - 20 m
20 - 8 m
8m
Cavability
Nil
Poor
Fair
Good
Very Good
Fragmentation
-
Large
Medium
Small
Very Small
Secondary Blasting
-
High
Variable
Low
Very Low
TABLE 12 Resulting fragmentation at our Southwest LHD Blocks at 908ML using the Block Cave Fragmentation Program (BCF). Average Block Size, m3 Primary Fragmentation Secondary Fragmentation range from start to exhaustion of the mining block
Maximum Block Size, m3
0.37
13.21
39.8
56.5
6.30 - 12.80
52.70 - 62.70
65.80 - 76.40
Mean = 1375 Standard Deviation = 682/100 = 7 Draw Control Factor = 0.75 (168 − 120 ) × 100 Dilution Entry = 168 = 28 %
LAYOUTS
DILUTION ENTRY The dilution entry percentage is the percentage of the ore column that has been drawn before dilution appears in the drawpoint and is a function of the amount of mixing that occurs in the draw columns. The mixing is a function of the following: a) ore draw height, b) range in fragmentation of both ore and waste, c) drawzone spacing, d) range in tonnage drawn from drawpoints. The range in fragmentation size distribution and the minimum drawzone spacing across the major apex will give the height of the interaction zone (HIZ) as shown in Figure 4. There is a volume increase as the cave propagates so that a certain amount of material is drawn before the cave reaches the dilution zone. The volume increase or swell factor are based on the fragmentation and applied to column height. Typical swell factors are fine fragmentation 1.16, medium 1.12, and coarse 1.08. A draw control factor is based on the variation in tonnes from working drawpoints as shown in Figure 3. The formula shown below was based on the above factors to determine the dilution entry percentage. A−B × C × 100 % A
where A = Draw Column Height × Swell Factor B = Actual Height of Interaction Zone (Calculated HIZ/Draw Control Factor) C = Draw Control Factor Example: DPs
W1
E1
W2
E2
W3
E3
W4
E4
Monthly tonnage
2000
800
1000
2500
600
1500
800
1800
400
Blocks less than 3 2m ,%
0.18 - 0.34
At the last slusher blocks at 908 ML where air blasts were experienced, it was found out through back calculations that the rate of draw was much greater than 17 cm than the actual caving rate of only 8 cm. The process emptied out the available ore at the undercut thereby creating a big void resulting to the series of air blasts. It can be concluded that during the early stages of a block, the draw rate should be minimised to ensure that the gap between the broken ore and the cave back is at a minimum but enough to ensure propagation of caving.
Dilution Entry =
Blocks less than 1 m 3, %
The objective of the lay-out design is to achieve optimum extraction while minimising problems that may interrupt production and require expensive remedial measures. A mining area will be divided into blocks, which are mined sequentially in a systematic manner. The size of each block will be determined by the area of the undercut required to promote satisfactory caving (hydraulic radius). The LHD system of draw offers high productivity of coarse ore but this system requires wide openings, large radius turnouts, and regular maintenance of the haulage invert. Owing to larger excavation compared to the slusher and gravity type of draw, stability of the openings requires special attention during production. Wherever possible, the permanent mine installations and important production openings such as drawpoints, dumping points, breaker stations, crusher stations, etc should be located in the better quality sections of the rockmass. Openings or excavations should be oriented as closely perpendicular to the strikes of the dominant, low shear strength structural features. As shown in Figure 5, Philex adopted the El Teniente layout due to its versatility, efficiency and particularly its stability as compared to the other layouts, Henderson and Herringbone layouts. Generally, the production lines are oriented in an east-west direction making it parallel to the major principal stress direction and perpendicular to the minor principal stress direction. With this layout, the major apexes are in a stable condition. The undercut lines are also oriented in an east-west direction to give a shielding effect to the production lines during the undercutting stage. However, scanline mapping results at our Southwest LHD blocks showed that the major joint sets direction in the host rock (meta-andesite) are almost parallel to the openings. This condition causes spalling of rock blocks and slipping of fan holes during the blasting of adjacent undercut lines. To remedy this, the drilling drifts at the undercut level were now oriented north-south or perpendicular to the trend of the joints.
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ROCK MECHANICS AS APPLIED IN PHILEX BLOCK CAVE OPERATIONS
FIG 3 - Determination of the draw control factor (DCF).
FIG 4 - Determination of the height of interaction zone based on RMR.
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R S DOLIPAS
FIG 5 - Philex block cave operation showing production level layout.
UNDERCUTTING Undercutting is one of the most important aspects of cave mining as not only is a complete undercut necessary to induce caving, but also the undercut can reduce the damaging effect of induced stresses. The normal undercutting sequence as being practiced in Philex is to develop the drawbell and then to break the undercut into the drawbell. In high stress environments, the pillars and brows are damaged by the advancing abutment stresses. The damage caused to pillars around drifts and drawbells by abutment stresses is significant and the major factor in brow wear and excavation collapse. Rockbursts are also located in these areas. The solution is to complete the undercut before the development of the drawpoints and drawbells. In the normal undercutting sequence, breaking of the undercut should be slower than the failure of the back, but it should be faster than the failure of the extraction horizon caused by high abutment stresses.
402
Where possible, it is preferable to mine from a weaker to a stronger rockmass, or from a failed to an intact rockmass rather than vice versa. The mining of a new block should retreat from an adjacent caved block rather than advance towards it. This procedure will avoid the creation of a potentially highly stressed pillar between the two mining blocks. It would also be preferable to configure the cave line perpendicular to dominant structures. At Padcal, the general mining retreat is from west to east making the caveline almost perpendicular to dominant fault structures. Succeeding blocks are mined with the caveline slightly slanted towards southeast to avoid stress concentration at the corner it makes with the boundary of the existing block. What happened to the slusher blocks at 908 ML was the inverse of what should be the proper retreat. Undercutting was started at the east side of the block and retreated towards the west. The caveline was advancing against the dip of the dominant fault structures as well as the major joint sets. The process resulted to the slow slipping of blocks.
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ROCK MECHANICS AS APPLIED IN PHILEX BLOCK CAVE OPERATIONS
between the components of the initial and final stages. The initial support will be installed concurrently with advance to control deformation and will preserve the integrity of the rockmass. The final support will cater the induced stresses and stress changes that will result from mining operation. An integrated support system consists of components that are interactive and the success of the system depends therefore on correct installation and on the use of materials of the right quality. Furthermore, the effectiveness of a support system can be related directly to an appreciation of the rockmass and the stress environment, as well as to the correct selection, timing and installation of the support. At Philex, where the mining rockmass rating at the production level falls under class 3A and 3B, supports at the production drifts and draw cross-cuts consists of rockbolts, wiremesh, cable bolts, and shotcrete. At the pillar walls of the draw cross-cuts, four rows of cable straps at 0.80 metres vertical distance are installed. The straps are grouted at both ends. Rockbolts, wiremesh, tendon straps, cable bolts and shotcrete in place of steel sets and reinforced concrete now reinforce most of the draw points. This support system is installed at areas where the rockmass are blocky. Total savings from this support system amount to at least P 150 000 per draw point. Draw points where lithologic contact are present are supported by reinforced massive concrete with rigid steel sets. Tables 13 shows the relationship of the rockmass rating (RMR) as against the adjusted rockmass rating (MRMR). Table 14 also shows the calculation of the length of rock bolt and cable bolt in relation to MRMR and the state of stress.
SUPPORT REQUIREMENTS It is ironic in a block caving operation that we want the extraction openings stable throughout the life of the block and just 16 to 20 metres above it, we want the ore column to cave as easy as possible. Accessibility of the production drifts, ventilation and access drifts and cross-cuts and dumping points should be maintained to ensure continuous and efficient extraction operation. The rock mass surrounding the extraction openings is subjected to four stress cycles in all caving situations and a fifth cycle if the retreat is towards an unfavourable geological structure. a.
adjustment of the rock mass to the openings;
b.
abutment stresses ahead of the undercut, and re-distribution of stresses around the caved area;
c.
uplift after the undercut is complete, with the removal of the vertical stresses;
d.
vertical loading on the apexes from point loads and an increasing column of caved material; and
e.
high toe stresses if wedge failure occurs against a structural fault.
With these conditions, it is a must that all openings in the production level should be reinforced and supported to last until the life of the block. The support system should be designed and agreed prior to the development stage so that there is interaction
TABLE 13 Support table using MRMR (after Laubscher 1994). MRMR
RMR 1A
1B
2A
2B
3A*
3B*
4A
4B
5A
5B
←Rock reinforcement – plastic deformation→ 1A 1B 2A 2B
a
a
3A
b
b
a
a
3B
b
b
b
b
b
c
4A
r
r
c
c
c
d
d
e
f
f
c+l
f/p
h + f/p
h + f/l
h+ f/l
h + f/p
f/p
t
4B* 5A 5B
d
t
*Philex Support technique: a
- Local bolting at joint intersections
k
- Rigid steel sets
b
- Bolts at 1 m spacing
l
- Massive concrete
c
- ‘b’ and straps and mesh if rock is finely jointed
m - ‘k’ and concrete
d
- ‘b’ and mesh/steel-fibre reinforced shotcrete bolts as lateral restraint
n
- Structurally reinforced concrete
e
- ‘d’ and straps in contact with or shotcreted in
o
- Yielding steel arches
f
- ‘e’ and cable bolts as reinforcing and lateral restraint
p
- Yielding steel arches set in concrete or shotcrete
g
- ‘f’ and pinning
q
- Fill
h
- Spiling
r
- Bolts and rope-laced mesh
i
- Grouting
s
- Rock replaced by stronger material
j
- Timber
t
- Development avoided if possible
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FIG 6 - Conditions without the aid of rock mechanics in planning.
TABLE 14 Calculation of rockbolt and cable bolt length (after Laubscher 1994). Support Element
Low Stress
High Stress
Bolts - Length - Spacing - Type
= 1 m + (0.33 W × F) =1m = Rigid Rebar
= 1 m + (0.5 W × F) = 1m = Yielding, eg Cones
Mesh
0.5 mm × 100 mm aperture
0.5 mm × 75 mm or 50 mm aperture
Deep Seated
Cables = 1 m + 1.5 W
Steel ropes = 1 m + 1.5 W Long Cone Bolts
Shotcrete linings
Mesh reinforced shotcrete
Mesh reinforced shotcrete
Arches
Rigid Steel arches Massive Concrete
Yielding Steel arches Reinforced Concrete
Surface Restraint
Large Washers (triangles) Tendon Straps
Large Plate Washers Yielding Tendon Straps
Corners
25 mm rope cable slings 25 mm rope cable slings
Brows
Birdcage Cables From Undercut Level Inclined pipes
Birdcage Cables From Undercut Level Inclined pipes
Repairs
Grouting Extra bolts and cables Plate straps and arches
Grouting Extra bolts and cables Plate straps and arches
Where W is the span of the tunnel and F is based on:
CONCLUSION Geomechanical issues such as cavability, fragmentation, drawzone spacing, dilution entry, draw control, lay-out design, undercutting sequence and support requirements can be assessed provided adequate and accurate geotechnical data are gathered and that geological variations are recognised. It is essential that classification data are made available at an early stage, so that the correct decisions are made on mining method, layout and support requirements. Although, rock mechanics in Padcal’s mechanised block cave operation is still in its infancy stage still, it is in line with the philosophy that one must crawl before he can walk. The key to a sound understanding of the rockmass behaviour in block caving is by first knowing and documenting cause-effect relationship with a full knowledge of ground conditions, draw history and ground movement. With this in hand, much can be learned to use rock mechanics as a design and predictive tool. Without rock mechanics, mine planning in block caving is carried out in the dark as shown in Figure 6.
ACKNOWLEDGEMENT The author is very much indebted to PHILEX Management, most especially to Messrs Gerard H Brimo and Leonard P Josef, Chairman of the Board and Senior Vice President for Operations respectively for giving me the opportunity to prepare this paper for the MassMin 2000 Conference. Heartfelt gratitude and appreciation is also extended to Rosauro Fernando G Agustin, Vice President for Operations for his unwavering encouragement and assistance. Special acknowledgement is also due to the geology and exploration department personnel and also to the author’s family for their support and guidance.
MRMR = 0 - 20;
F = 1.4
MRMR = 21 - 30;
F = 1.3
MRMR = 31 - 40;
F = 1.2
REFERENCES
MRMR = 41 - 50;
F = 1.1
MRMR = 51 - 60;
F = 1.05
MRMR = >60;
F = 1.0
Brady, B H G and Brown, E T, 1985. Rock mechanics for underground mining, Chapters 12.4.8 and 15.5, pp 312-313 and 392-402. Laubscher, D H, 1977. Geomechanics classification of jointed rock masses-mining applications. Laubscher, D H, 1984. Design aspects and effectiveness of support systems in different mining conditions. Laubscher, D H, 1995. Cave Mining – State of the Art.
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Block Cave Undercutting — Aims, Strategies, Methods and Management R J Butcher1 UNDERCUTTING - AIMS AND STRATEGIES
ABSTRACT Many block caving operations experience difficulties during undercutting, resulting in draw horizon damage, production loss and extensive drift repairs for the life of the block cave. This paper deals specifically with block cave undercut aims strategies, methodologies and management to avoid the problems indicated above.
INTRODUCTION In block cave mining, undercutting (Figure 1) is probably one of the most important of all the activities involved. Therefore, the correct design of the undercut is critical to the success of the block cave. Poor undercut strategies and designs can result in severe draw horizon damage, extensive drift repairs and production loss. Today many different undercut strategies and designs exist, each having its own particular advantages, disadvantages and applications. This paper details the main undercut strategies and designs and describes undercut management to avoid block cave damage and production loss. 1.
Senior Mining Engineer, SRK Consulting, 265 Oxford Road, Illovo, Johannesburg, South Africa.
Aims of undercutting The primary aim of the blockcave undercut is to extract a void of sufficient dimensions to allow caving to occur. However, in developing an undercut of the required area, an increase in stresses acting on the periphery of the undercut area normally results. This in turn causes damage to the rock mass surrounding the undercut. It therefore follows that the second aim of undercutting is to achieve the required undercut dimension to initiate caving, with minimum damage to the surrounding rock mass. The size of undercut necessary to initiate caving is a function of the relationship between the quality of the rock mass and its cavability. This is commonly quantified by the relationship between the Mining Rock Mass Rating, (MRMR), derived from rock mass classification of the orebody (Laubscher, 1994) and the extent of the undercut area at which caving will propagate. An example of this is the relationship between the MRMR and the hydraulic radius (the area of the undercut divided by its perimeter) shown in Figure 2 (Laubscher, 1994). As a general rule, the weaker the orebody, the smaller the hydraulic radius
FIG 1 - Trackless block cave layout (Brumleve and Maier, 1981).
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A common misconception is that stress levels will automatically reduce once the undercut has reached the caving hydraulic radius. However, in practice, stress reduction only occurs with cave propagation, which in turn relates to the cave draw. Therefore, the third aim of block cave undercutting is to advance (in time as rapidly as possible) to caving hydraulic radius, initiate caving, propagate the cave and consequently reduce undercut abutment stresses. In practice this requires that the undercut preferably advances from the weakest ground to the strongest, and away from the direction of action of the maximum principal stress. Observations of undercut abutment damage from an Asbestos Mine in Zimbabwe during the late-1990s have shown that even though a significant stress reduction occurs once the cave has propagated, stress levels only appear to return to the pre-undercut magnitudes once 25 per cent to 30 per cent of the block’s tonnage has been drawn. This means that peripheral excavations on the draw horizon may be subject to abutment stresses for longer than might have been expected. It is clear that the second and third aims of undercutting are focused on the control and reduction of stress levels. It therefore follows that undercutting strategies must directed at stress abatement to satisfy these aims.
Undercut strategies Three strategies are used for the undercutting of orebodies: FIG 2 - Stability diagram (after Laubscher, 1994).
required for cave initiation. In practice, orebodies normally require an undercut with a hydraulic radius of between 8 to 26 for caving to occur. A hydraulic radius of eight corresponds with an undercut area of 30 m × 30 m, and of 26 with an area of 105 m × 105 m. The magnitude of undercut stresses (before cave initiation) can be related to the extent of the undercut as a percentage of the actual hydraulic radius required for cave propagation. Table 1 (Butcher, 2000) shows undercut stress conditions based on simple numerical modelling of a blockcave with a required hydraulic radius of 17 and an MRMR of 30 to 35. TABLE 1 Undercut stress conditions related to required caving hydraulic radius. Undercut extent as a % of required caving hydraulic radius
Undercut stress condition
Remarks
25
Low
Limited rock mass damage
50
Stress levels start to become of concern
Onset of critical rock mass damage
75 to 110
Stress levels at maximum
Severe rock mass damage
From the data in the table, undercut rock mass damage starts to become critical once 50 per cent of the required caving hydraulic radius has been undercut. Stress levels are at their maximum when the undercut has reached the block’s hydraulic radius. The practical significance of this is that undercut level support requirements and blasthole losses increase significantly once 50 per cent to 110 per cent of the required caving hydraulic radius has been undercut.
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Pre-undercutting In a pre-undercut situation, the undercut is fully developed to the hydraulic radius position. The drawhorizon is then developed below in a de-stressed conditions. The main advantage of such a method is that draw horizon damage is reduced. The disadvantages are:
• a time delay between undercutting and draw horizon development;
• a sequencing problem between draw horizon development and undercutting;
• production delays resulting from the above and the subsequent interruption to project cash flows;
• possible high stress remnants due to compaction of blasted undercut ground making draw horizon development difficult; and
• slower initial production due to drawpoint hang-ups resulting from ground compaction. As a result of the above, pre-undercutting of deep level block caves is not recommended unless a high level of project planning and mining discipline is achievable.
Post-undercutting In this method the draw horizon is fully developed, and the undercut is then advanced over this extraction level horizon. The main advantage of a post-undercutting strategy is that blocks can be brought into to production quicker. This is due to the fact that the draw horizon is fully developed and that production from the complete block can take place as soon as undercutting has been completed. The main disadvantage is the stress induced draw horizon damage caused by the advancing undercut abutment, and the subsequent drift repairs which are required. South African experience (Bartlett, 1992) has shown that during the advance of the undercut over the draw horizon, tunnels are subjected to stress magnitudes about three times higher than the pre-undercut levels. In this environment inter-drawpoint/drawbell pillars are sometimes subjected to severe stress induced damage. This damage is related to competency of the rock mass and the size of
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BLOCK CAVE UNDERCUTTING — AIMS, STRATEGIES, METHODS AND MANAGEMENT
the inter-drawpoint/drawbell pillar. This relates directly to the percentage extraction on the draw horizon. During the early-1990s, it was found that draw horizons suffered severe damage from undercut abutment stresses when more than 80 per cent of the draw horizon was extracted. It was further found that extraction drifts suffered slight to moderate damage when less than 60 per cent of the draw horizon was extracted (Butcher, 1999). A review of undercutting history at Kimberley mines (Gallagher and Loftus, 1961) indicated that post-undercutting became unacceptable in terms of draw horizon stress induced damage once block operational depths exceeded 500 m. It can be concluded that, despite the apparent advantage of quicker block production build-up, the possibility of severe stress induced drift damage must be seriously considered before employing a post-undercutting strategy. As a general guideline, the use of post-undercutting strategies should be critically assessed when the operational depth of the cave is greater than 500 m, when the caving area has a hydraulic radius of greater than 17, and when draw horizon extraction percentage exceeds 50 per cent.
Advance undercutting In this strategy, only a limited amount of development takes places on the draw horizon before undercutting. The undercut is then advanced over the draw horizon, and the remaining extraction level development is conducted in de-stressed conditions. This is in essence a compromise between post and pre-undercutting strategies, in that:
• draw horizon damage is reduced, because, compared with the post undercut strategy, the extraction ratio on this horizon is decreased;
• the cave is brought into production quicker than with a pre-undercutting strategy, reducing the problems associated with increased development times;
• the probability of the formation of stress-inducing remnants, owing to muck pile compaction, is reduced;
• a separate level is still required for undercutting; and • this strategy is slower than post undercutting, due to the remaining horizon level development that has to take place after the undercut has advanced (this excludes the time that is required for draw horizon drift repairs that will almost certainly be required with the post undercutting strategy). In reality, drifts are fully developed before undercutting, with cross-cuts and drawbells being developed once the undercutting is complete. South African mining experiences (Butcher, 1999) have shown that advanced undercutting can reduce draw horizon damage to tolerable levels if the extent of development on the draw horizon is limited to less than 60 per cent of the area prior to undercutting. A problem may exist if the cutting of drawbells is carried out too close to the undercut face, since these excavations can then sustain abutment damage. It would therefore be essential that drawbells are developed behind the undercut face, by a distance at least equal to the middling between the draw horizon and the undercut level. The current trend in block cave design is to use the advance undercutting strategy, with the discussed disadvantages being reduced to tolerable levels by good planning, appropriate support, and equipment to deal rapidly with hang-ups in drawpoints.
UNDERCUT EXTRACTION METHODS There are five main considerations in the selection of the undercut extraction method. These are:
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Operational depth/stress regime In general, to minimise stress levels, the undercut extraction system should remove as little tonnage as possible. Narrow undercuts remove limited tonnage compared with fan blasted undercuts and are therefore preferable in deep level situations.
Skill levels of undercutting crews The success of undercutting relies heavily upon the skills and mining discipline of undercut miners and production officials. These skills are important to ensure that the undercut can be correctly cleaned without the occurrence of remnant pillars due to choke blasting. In essence, satisfactory undercutting revolves around the quality of drilling, blasting and undercut cleaning, rather than design aspects. The selection, training and motivation of undercut crews must take the highest priority.
Block cave tonnage build up profile This is important in a panel caving situation, in which production comes from both cave and undercut tonnage sources. In this situation, undercutting is considered to be a means of production. Fan undercuts have a greater tonnage capacity than narrow undercut systems.
Technology This should be taken into account in the design stages, as many narrow and fan undercutting methods require the use of sophisticated long hole drilling equipment, for the accurate drilling or undercut blast holes. Owing to the narrow height associated with crinkle cut method, accurate drilling using sophisticated equipment is required to prevent the formation of pillars. This may not be available, nor be capable of being maintained in many developing countries. In such situations, flat or fan undercuts of limited height are considered to be preferable.
Undercut heights Experience has shown that fan and narrow undercutting tend to have average undercut heights in the region of 15 m to 4 m respectively. From a stress damage reduction point of view, the height of deeper level undercuts should be as low as possible. Large undercut heights do not affect the cavability potential of the block (Laubscher, 1994). However, production tonnage requirements may necessitate that heights have to be increased to meet project cash flows. Taking the above five considerations into account, a description of the main undercutting extraction methods is given in the following section.
Undercut extraction methods Three undercut methods are considered – fan, flat and narrow crinkle-cut.
Fan This type of undercut has probably been the most widely used. It has been used for undercutting of grizzly, LHD and slusher-drift caves. Fan undercutting of caves can be done with or without the development of a separate undercut level (Figures 3 and 4). Fan heights can be increased to produce additional undercut tonnage. Fan undercuts can be drilled with a variety of equipment from hand held stopers to sophisticated trackless equipment. The fan system can easily be adapted if a mine changes from open stoping or sublevel caving methods to block caving.
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R J BUTCHER
FIG 3 - Block cave with separate undercurt level (after Rood and Bartlett, 1994).
FIG 4 - Block cave undercut without separate level.
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The main problem with this undercut method is pillar formation due to blasthole loss, and choke blasting. Poor design is also a major factor affecting pillar formation (Grobler et al, 1994). These problems can be alleviated by correct design, reducing the number of fan pre-drilled blastholes, and by ensuring that choke blast conditions do not exist before blasting.
Most narrow undercuts have been implemented in mines that have had many years of undercutting experience (Stewart, 1981), and therefore the required skill levels were available for implementation of the flat undercut method. Flat undercuts have limited tonnage capacity and therefore cannot be considered as a method of production.
Flat
Narrow crinkle cut
These undercuts have been used successfully at Bell and El Teniente mines (Bartlett, 1997). They can be implemented using both trackless and conventional technology. An example of a narrow undercutting layout is given in Figure 5. Narrow methods are normally used in deep level undercut situations for the following reasons:
This type of undercut is depicted in Figure 6. Crinkle cut undercuts are being introduced in an effort to overcome the problem of choking and stacking, by inclining the undercut over the major apices. In theory, this inclined section is self cleaning. However, in order to facilitate self-cleaning, the inclined section should be developed at an angle of 55° or steeper. Experience from narrow vein mining has shown that stopes dipping at less than 55° may not be totally self cleaning. Therefore, if the inclined section is shallower than this, the possibility of choking may still be present. The problem is further compounded by the fact that these inclined sections cannot be manually cleaned. In addition, in the flat sections of the undercut, the original cleaning problems will still exist. At the intersection of the inclined and flat undercut portions, brow damage may occur in deep level block caves. The sequencing of blasting and cleaning both the flat and inclined portions may also prove difficult due to a lack of access (Figure 6). This type of undercut should have greater advance rates due to its self cleaning ability and therefore should be suitable for high stress situations. Owing to the inclination of the blast holes, high quality drilling procedures will have implemented to prevent pillar formation.
• greater advance rates, due to the fact that a lesser amount of drilling and charging is required;
• undercut blast hole loss is less due to the fact that fewer holes are required;
• lower undercut heights reduce the magnitude of undercut stresses. However, flat undercuts require additional management to prevent:
• occurrence of remnant pillars between drilling drifts (Figure 6); and
• undercut choking and stacking, due to cleaning difficulties (Figure 6); The cleaning of this type of undercut is especially important. Owing to the narrow height, muckpile stacking can result, causing remnant pillars due to choke blasting. Recent experiences in South America (Bartlett, 1997) have shown that the above problem can easily occur. In this regard it is essential that undercut draw discipline is maintained.
UNDERCUT MANAGEMENT Management is probably one of the most neglected aspects of block cave implementation, with undercutting problems more often being due to poor management than to design faults. Poor
FIG 5 - Conceptual plan of flat undercut showing possible problem areas.
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FIG 6 - Crinkle cut undercut potential problem areas.
management is mainly due to the lack of appreciation of the required level of control needed for successful undercutting. In this respect the following are the main problems associated with poor undercut management.
once undercut tunnels suffer stress induced damage, blasthole loss increases, rates of advance reduce and ground control becomes difficult and costly. Table 3 shows the required maximum tolerable lag distances and intervals at which the face configuration should be reviewed.
Poor undercut configuration control Large irregularities in the horizontal geometry of the undercut front (cave line) cause an increase in stress concentration, resulting in serious damage on the undercut front. It is therefore imperative that all large irregularities are kept to a minimum. The irregularities can be thought of in terms of panel lag distances (as in tabular longwall stoping). Based on current knowledge, suggested guidelines for lag distances are given in Table 2.
TABLE 2 Undercut face lag distances and corresponding stress conditions and damage (Butcher, 1999). Undercut front lag distance (cave line)
Undercut front stress condition (cave face)
Damage
60
metres
FIG 3 - 9920 Level geotechnical plan.
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P L WOOD, P A JENKINS and I W O JONES
draw, with the concurrent extraction of equal ore volumes from a minimum of three draw points (Wood, 1999). This strategy produced a marked improvement in both the amount and grade of the extracted tonnage and was continued, with a grade shut-off, down to the 10 055 level (470 m below surface). At the 10 055 level, extremely poor ground conditions were encountered in the south of the disseminated orebody, resulting in the loss of one of the transverse cross-cuts. Further cross-cut collapses occurred soon after on the 10 077 and 10 000 levels. The poor conditions correlate with widening of the hangingwall shear as mining approached the inflection zone. Above the inflection zone, SLC development was mined well in advance of production and initially supported by split sets and mesh with cable bolts installed at intersections and in areas of poor ground on completion of cross-cuts. In response to the deterioration of production brows on the first two production levels, 50 - 70 mm of shotcrete was sprayed over the mesh on all existing orebody development. The frequency of brow failures was further reduced by the installation of 5 m cable bolts in the backs between production rings. Despite these measures, floor heave occurred near the hangingwall shear and there was significant cross-cut closure due to shear failure and dilation of sidewalls in the more competent ground, that formed a broken zone of 0.5 - 2.0 m thickness behind the shotcrete. The inability to keep drives open through this shear zone led to the additional development of footwall drives to enable footwall retreat of transverse cross-cuts on some levels. However, the presence of swelling ground on the footwall side of the orebody prevented this strategy from being a complete success. The spraying of fibrecrete as part of the development cycle was first introduced on the 10 055 level and grouted 3.0 m long split sets were used as spiling bars to help get through the hangingwall shear zone. Fibrecrete, with 50 kg/m3 steel fibre, was applied in two layers to 125 - 150 mm thickness with mesh reinforcement between the layers. The initial fibrecrete layer and mesh, installed with split sets, was placed at the face with the second layer following after the next cut. Cable bolt installation followed 10 - 20 m behind the face to complete this integrated development and support cycle. Through the hangingwall shear zone, the rings of 13 cable bolts were installed at a 1.25 m spacing and advance rates reduced to 1.5 m per cut. The support strategy above the inflection zone is further described in the paper by Struthers and Keogh (1996), which also details other remedial support measures used in localised areas. Mining of the SLC was abandoned on the 10 030 level due to major ground problems. On the 10 000 Level access from the hangingwall was prevented by cross-cut collapses although some mining was achieved in the northern section by reverting to footwall access. However, drive closure and poor back conditions severely hampered these activities.
Caveability of the disseminated orebody To assess the two preferred methods the caveability of the disseminated orebody through and below the inflection zone was evaluated. Existing geotechnical data from core logging was combined with additional detailed mapping of rock type, mineralogical zonation and point load testing of core from the 10 030, 9920 and 9760 Levels for this work. The results of the caveability assessment, in which RMR values were converted to MRMR (Mining Rock Mass Rating, after Laubscher, 1977) with adjustments for alteration and stress effects, are summarised in Table 3 and illustrated on the stability diagram, Figure 4 (Laubscher, 1994). TABLE 3 Summary of caveability assessment.
RMR
Lift I (9875 - 10 000 Level)
Lift II (9760 - 9875 Level)
40 - 60
50 - 80
Alteration Adjustment
82 %
100 %
Stress Adjustment
70 %
85 %
23 - 34
43 - 68
19 % >2 m3
70 - 75 % >2 m3
MRMR Drawpoint Fragmentation
RE-EVALUATION OF MINING BELOW THE 10 000 LEVEL Geological conditions were known to be more disturbed through the inflection zone, with a greater frequency of fracturing within the disseminated orebody as well as on the hangingwall contact. Below the inflection, from the 9920 m Level down, the orebody steepens and the degree of geological disturbance lessens with the hangingwall shear zone tending to decrease in width, though still prominent along the southern section of the hangingwall contact. Owing to the expected difficulty of mining in the inflection zone, an investigation of alternative mining methods and strategies for the disseminated orebody below the 10 000 Level was conducted (AMC, 1996). A number of methods were considered but two preferred methods were identified, these being to continue the SLC through the inflection or Block Cave the column of ore from below the inflection.
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FIG 4 - Stability diagram, proposed Perseverance block cave (after AMC, 1996).
These findings indicate that the orebody between 10 000 m Level and 9870 m Level would indeed cave and the draw point fragmentation would result in 81 per cent of material 50 m inside the footwall) was reported in 1989 for the first time at levels 275 m, 320 m and 370 m. Extraction took place around level 610 m at that time. Cracks and movements appeared in the floor and in some infrastructures.
Since then, the instability has essentially extended northward and southward, but has not deepened westward. It is now mapped between production level and surface from mine co-ordinates Y14 to Y26 (the first north quarter of the mine, ie about 1 km). Figure 3 presents a synthetic view over the footwall’s slope at co-ordinates Y23 where major moving structures are represented. Figure 4 shows the failure’s lateral and longitudinal extension on level 320 m. Failure is obvious through shear and open cracks in the walls and floors of the drifts (see Figure 5), concrete posts shearing and bending failures (see Figure 6), railroad movements relative to the wall, shotcrete scaling and fallen wedge. In the worst case, the roof collapsed. Between mine co-ordinates Y22 and Y26, the actual instability’s outer limit is made of a major structure or series of intermediate structures dipping steeply between the surface and level 230 m (about 70 degrees), and flattening between level 230 m and 540 m (50 to 60 degrees). This structure (or group of structures) is undulating and filled with 1 mm to several centimeters chlorite and clayey minerals and has been reopened everywhere. Normal faulting mismatched the structure and created several centimetres-large apertures (see Figures 5 and 6).
FIG 3 - Mapped failure on section Y23 until year 2000.
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FIG 4 - Mapped failure on level 320 m.
FIG 5 - Borehole sheared by a gliding structure on level 320 m.
A joint set that dips at 45 degrees toward S-E also controls the failure at the northern closure and at the slope’s toe (systematic shotcrete scaling between level 540 m and 740 m). Very near the ore contact, the roof has often collapsed, but this could be the effect of superficial perturbations or blasting at the sublevels beneath. While structures that moved in the upper part of the footwall are separated by zones where rock has nearly no signs of deterioration, under level 540 m, joint displacements are visible everywhere.
FIELD INSTRUMENTATION Compared with an open-pit and except at the top, slope failure cannot be observed from ‘outside’ at the Kiruna mine because of the covering caved rock-fill coming from the hangingwall. An arsenal of field instrumentation has been applied to locate and eventually quantify displacements in the footwall. Most of it has been placed underground.
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FIG 6 - Sheared concrete post on level 540 m.
Total station A network of point lines has been installed at surface all around the mine, on the hangingwall and the footwall. These points are surveyed annually with total station equipment. Although this equipment has a precision of about two millimetres under perfect conditions, some points are installed in soil and not in rock and due to the half-automated measurement mode, movements less than one or two centimetres cannot be detected. This resolution is not sufficient for the footwall since rockmass differential displacements in the order of a centimeter can cause visible roof collapses.
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The same kind of equipment has been used underground for years in order to catch displacements of the order of a millimetre. Points where set along galleries and, from a starting reference point whose absolute co-ordinates were supposed constant over the time, co-ordinates of the other points where measured in sequence. This one-way measurement method appeared very ineffective because small measurement errors on a point were reported and cumulated on the others. As one measurement series can consist of a 4 km long gallery with up to 40 points installed, errors on the last points’ deduced X, Y, or Z co-ordinates could reach 50 cm. On short distance, results where quite acceptable. A new methodology with double check is now being tested.
Distometer Distometers are robust equipment allowing measuring deformations of the order of the millimeter. When lines of five to ten metres are installed along drifts orthogonal to the mine’s strike, failure progression is monitored. Figures 7 and 8 present an example on level 230 m where line 1 was installed closest from the extracted orebody and line 13 about 100 m further away. These lines have been in function since 1987 and last measurement is dated to 1999. Instability cut line 1 early in the late-1980s and then reached lines 12 and 13 that expanded by about 10 and 14 cm about seven years later. During that time, rock mass relaxation affected all the other lines in between.
Time Domain Reflectometry (TDR) The TDR technology has been used since the 1980s at LKAB and gave very satisfactory results. A lot of cables have been installed in the mine and especially in the old levels. Most of them are now inside the failure zone and do not give any more information. Compared with distometer, TDR cables allow for deep rock mass monitoring, far away from the drifts influence zone.
FIG 8 - Distometer lines localisation on level 230 m.
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FIG 7 - Distometer measurements on level 230 m between 1987 and 1999.
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New cables would be needed now but their installation in disaffected levels would require huge restoration costs (scaling, electricity, pressured water, …). A solution had to be found to replace TDR cables and the macro-seismic solution arose.
Macro-seismic A macro-seismic system developed by the CANMET Mining and Mineral Sciences Laboratories has very newly been installed with purpose to locate seismic events occurring in the inner footwall. The system registers the whole waveform, which gives more information on the failure mechanism: tension or shear failure in intact rock, slips along existing joints, with possibility for orientating faulting planes. Most of the galleries are located in the vicinity of the orebody and there is only one access deeper in the footwall (located at Y24). If failure deepens westward, the macro-seismic system would then be the only means to monitor it. This system should also help distinguishing between biased gallery observations that are attributed to superficial alterations from the more global footwall failure. For example small rock-bursts and wall spalling are observed around level 740 m. It is however very difficult to say if they result from superficial and local overstresses around the openings due to mining process or if the phenomenon affects the whole rock mass. Seismic events’ localisation should solve that problem. Efficiency of the macro-seismic system will greatly depend on the frequency and the energy of the events that may occur. Little information is available in the literature about hard-rock slope seismic monitoring. The Kiruna macro-seismic system will be installed primarily for a trial period of one year. After this period, an evaluation will be done and a decision on how to go on will be made.
FAILURE MECHANISM, A LITERATURE REVIEW Since the 1980s, several engineering reports and doctoral theses have been written on hanging- and footwall stability at the Kiruna mine.
Planar shear failure along major structures Among them Hansagi (1967 and 1972) and Paganus (1973) advised delineating footwall failure by lines at 55 degrees starting at the production level and going up to the surface. Structures were poorly known at that time and this angle was chosen arbitrarily. However it now makes sense when considering planar failure along structures sub-parallel to the ore contact. This model has been successfully used for placing new infrastructures.
Circular shear failure From that model, Danhér-Lindqvist (1992) began to look at the possibility for a circular shear failure using Hoek and Bray’s charts (Hoek and Bray, 1981). This model has been calibrated and is still in use for mine planning. This theory implicitly assumes that no external pressure exerts onto the slope. Yet the slope is covered by the caved rock-fill which is moving downwards as the mine is being under production. Sjöberg (1999) studied large slope circular shear failure. He defined a methodology for estimating rock mass properties and used FLAC simulations in open-pit slope analysis. Problems arose to model the pit’s caved rock-fill behaviour. He showed that considering the caved rock-fill at rest would result in a stable slope. Representing drawing chimneys by very low stiffness zones gave a failure geometry that was not in agreement with the actual failure geometry. Sjöberg (1999) also looked at replacing caved rock-fill by inclined traction forces, deduced from the Silo Theory, in limit equilibrium routines and obtained quite satisfactory results.
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Limit Equilibrium Theory Lupo (1996) studied in details the possibility for a wedge-like failure using Hoek’s (1974) Limit Equilibrium Theory. A tensile failure seed is explicitly placed (5 m deep in Lupo, 1996) and grows vertically downward. The tensile failure transforms into a shear failure across the rock mass from a certain depth. A factor of safety is calculated for each seed position and tensile fracture’s depth. The configuration that gives lowest factor of safety is retained. Prognostics for the hangingwall and footwall at the Kiruna mine versus the extraction depth have been edited (Lupo, 1996). Looking at the hangingwall, good agreement with surface deformations was obtained. Lupo (1999) also confirmed the validity of this model for the hangingwall with numerical simulations with FLAC (Itasca, 1998). Rock mass cohesion and friction angle can be much higher in the Limit Equilibrium Theory than in the Circular Shear Failure Theory. Cohesion used in Lupo (1996), 1.2 MPa, was twice as big as in Danhér-Lindqvist (1992), 0.6 MPa.
CAVED ROCK-FILL’S EFFECT ON THE KIRUNA MINE’S FOOTWALL, DISCUSSION AND NUMERICAL SIMULATIONS WITH FLAC Discussion Several observations are to be made on these various theories. First, it is difficult to get representative rock mass and caved rock-fill parameters and to estimate the pressure conditions at the interface between the footwall’s intact rock and the caved rock-fill. The difficulty in obtaining caved rock-fill parameters is mainly a question of accessibility. Consequently, the discrimination between the three models (planar shear failure, circular shear failure or wedge-like failure) can’t be absolutely reliable. Depending on the assumptions made on the parameters, these models can fit the observations with a certain degree of agreement. Using the silo theory and calculating forces inside the funnels and against the footwall should be possible provided that the funnel’s width, the caved rock-fill parameters and the friction angle broken rock mass – footwall could be estimated. These parameters are still hardly known and research ought to be focused on these three points.
Numerical simulation with FLAC The FLAC-model simulates the Kiruna orebody with a 60 degrees inclination. A Mohr-Coulomb and a Ubiquitous Joint constitutive model are used. From the original topography, 15 m high levels of ore are removed at a time. Traction forces are applied onto the footwall and the model is stepped until equilibrium is reached. Methodology proposed by Sjöberg (1999) was applied to estimate the footwall’s cohesion, friction angle, tensile strength and deformation modulus from the RMR (Rock Mass Rating, Beniawski, 1989). The RMR varies horizontally and vertically in the footwall but 50 to 60 would be a fair estimate. Cohesion was then set to 0.6 MPa and friction angle set to 30 degrees. A mean level of 250 m was considered for the caved rock-fill. Lupo (1999) suggested that the caved rock-fill could be replaced by traction forces calculated with formulae from Soil Mechanics and Silo Theory. The vertical stress on the footwall results from the column of caved rock-fill placed above, while the horizontal stress would be derived from the Coulomb active earth pressure relationship:
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σ h = K a ⋅ σv
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Where σh is the horizontal stress, σv the vertical stress and Ka the Coulomb active earth coefficient. Sensitivity analysis was performed on Ka, supposed constant over the slope’s height as a gross estimation. Footwall failure did not occur for Ka values higher than 0.3 with the Mohr-Coulomb constitutive model. In the Ubiquitous Joint model, using joints dipping at 50 degrees toward the slope and having a friction angle equal to 30 degrees, instability was reached for Ka as high as 0.35 (see Figure 9). This is considerably less than the values needed in the Limit Equilibrium Theory as presented in Lupo (1996). The outer footwall failure’s limits were clearly circular for the Mohr-Coulomb constitutive model. The tendency to a circular failure shape could also be observed in the Ubiquitous Joint constitutive model (see Figure 9).
CONCLUSION Footwall failure should always be envisaged when planning a new sublevel or block caving operation. This hazard is indeed an observed fact at the Kiruna mine. The monitoring program should insure a correct failure localisation, quantification and interpretation. This last task is maybe the most difficult because it means determining the exact leading mechanism. Concurrent theories are still seriously discussed in Kiruna and each one has its defenders: planar shear failure along existing faults, circular shear failure and wedge-like failure across the rock mass.
A lot of key-parameters are still hardly known. Among them, caved rock-fill action against the footwall is significant. It has been showed that even for a rather poor rock mass quality, lateral support must not necessarily be huge for stabilising the footwall against circular shear failure. Coulomb active earth pressure coefficient up to 0.3 could be envisaged in Kiruna. Intuitively, re-mobilisation of existing structures is easier than breaking the rock mass. Seldon, Pakalnis and Board (1999) reported a complex wedge failure in the Kidd Creek’s hangingwall. Faults delineated a huge wedge with diffuse toppling at its toe. This kind of instability, indeed a parent of the Hoek’s (1974) one, could also apply to the Kiruna mine.
BIBLIOGRAPHY Beniawski, Z T, 1989. Engineering rock mass classification (John Wiley and Sons: New York). Dahnér-Lindqvist, C, 1992, Liggväggsstabiliteten i Kiirunavaara, in Bergmekanikdagen 1992, pp 37-52 (BeFo-Swedish Rock Mechanics Research Foundation: Stockholm), in Swedish. Hansagi, I, 1967. LKAB internal memorandum. Hansagi, I, 1972. LKAB internal memorandum. Herdocia, A, 1991. Hanging wall stability of sublevel caving mines in Sweden, Doctoral Thesis 1991:096D, Luleå University of Technology, Luleå. Hoek, E, 1974. Progressive caving induced by mining an inclined orebody, Trans Inst Min and Metall, Section A, London, 83:A133-A140.
FIG 9 - Shear failure in the footwall- Cohesion, 0.6 MPa, Friction angle, 30 degrees, ubiquitous joint dipping at 50 degrees and having a friction angle of 30 degrees.
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Hoek, E and Bray, J W, 1981. Rock slope engineering, (Institution of Mining and Metallurgy: London). Hustrulid, W, 1991. Hängväggstabilitet, koncept, LKAB internal report, in Swedish. Itasca Consulting Group Inc, 1998. FLAC V3.4, User’s Manual, Minneapolis. Itasca Consulting Group Inc, 1999. PFC2D V2.0, User’s Manual, Minneapolis. Lupo, J F, 1996. Evaluation of deformations resulting from mass mining of an inclined orebody, PhD dissertation, Colorado School of Mines, Golden. Lupo, J F, 1999. Numerical simulation of progressive failure from underground bulk mining operations, in Rock Mechanics for Industry (Eds: Amadei et al), pp 1085-1090, (Balkema: Rotterdam).
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Paganus, T, 1973. LKAB internal mine maps. Paganus, T, 1992. Kiirunavaara norra sprickor och sättningar i markytan, LKAB internal report, in Swedish. Seldon, C S, Pakalnis, R T and Board, M P (in prep, 1999). Complex failure of a large block affecting surface and underground operations at Kidd Mine. Kidd Operation Mines. Timmins. Sjöberg, J, 1999. Analysis of large scale rock slopes, PhD thesis 1999:01, Luleå University of Technology, Luleå. Stephansson, O, Borg, T and Bäckblom, G, 1978. Sprickbildning i Norra Kiirunavaaras hängvägg, Technical report 1978:51 T, Luleå University of Technology, Luleå, in Swedish.
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Sublevel Caving Design Sublevel Caving — Today’s Dependable Low-Cost ‘Ore Factory’
G Bull and C H Page
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Sublevel Caving — Today’s Dependable Low-Cost ‘Ore Factory’ G Bull1 and C H Page2 ABSTRACT Much of the theory upon which the Sublevel Caving (SLC) method is based was developed in Scandinavia many years ago. It was based mainly on ‘bin’ theory and ‘ellipsoid of draw’ configurations derived from sandbox models. In many cases, classical SLC layouts influenced by these early theories obtained relatively poor results. Early dilution entry, low tonnage factors and low recoveries were commonly experienced. SLC soon gained the reputation of being a high dilution, low recovery and development intensive method. This led to SLC falling into disfavour generally, other than as a pillar reclamation method used at the end of the life of mines. More recently a number of mines, notably those in Australia, have adopted SLC as a primary extraction method and have achieved good recovery results. The reason for the better results is believed to be an outcome of modifications in some of the basic design parameters and operational procedures. This has led to questions being raised regarding the original ‘bin theory’ assumptions and classical SLC layout designs. 1.
MAusIMM, Technical Director, SRK Consulting, 25 Richardson Street, WestPerth WA 6005.
2.
Corporate Consultant, SRK Consulting, Suite 800 – 580 Hornby Street, Vancouver, BC V6C 3B6, Canada.
It is clear that the fragmentation and flow characteristics occurring in a choke blast situation (compaction of waste, variable size and shape of fragmentation throughout the ring) conflicts significantly with the assumptions of regular loose flow of uniform, fine grained, material in a bin. Interactive versus independent draw conditions can also materially effect the outcome of a SLC operation. The good results obtained more recently can only be explained by using ‘new’ models of behaviour, which have to a large extent been borrowed from block caving. If these models are correct, they point to a number of design and operational changes that can be applied to improve the effectiveness of SLC. This paper discusses the factors that may not have been fully considered in early SLC design and modelling. It emphasises the importance of delayed dilution entry as a primary objective of sound SLC design and provides a basis for efficient SLC design and operation.
INTRODUCTION The apparent contradiction of SLC is a slice of broken material being drawn relatively clean while being surrounded by broken waste (Figure 1). It seems to defy logic, but experience shows it is possible. When a ring is blasted (typically 1000 – 2000 tonnes), the first part is drawn clean. Waste from above and
FIG 1 - Schematic view of sublevel caving.
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behind the ring then starts to come into the draw point and a mixture of ore and waste is drawn. The proportion of waste increases until shut-off is reached. When draw from each ring is stopped at shut-off point, some ore will be left behind. This remnant ore mixes with the previous ore/waste in the cave and this ‘dilution’ material increases in grade as the cave matures. ‘Dilution’ is any material (barren or mineralised waste and remnants of ore from previously drawn rings) that is not part of the current fired ring. The objective is to keep the waste out of the draw point for as long as possible, but to try and make the most out of zones of higher grade ‘dilution’. For this reason it is very important that a practical, common sense model of how, or why, SLC works is agreed. This model will then be the basis for determining critical design and operating aspects of SLC. The basis for any SLC model is as follows:
• The overlying rockmass must be able to cave freely and/or a large volume of broken material or introduced fill material must occur immediately above the ore to be mined. This broken material provides confinement to keep each consecutive blasted ring of ore ‘in place’ at the draw point.
• When ore is drawn from the blasted ring it is replaced by the broken waste rock from above or behind the ring (this ‘dilution’ material may carry grade).
• Because the waste and the ore are moving downward towards the draw point mouth, some mixing occurs. This mixture determines the grade of the material being drawn at the draw point.
• The layout design and operational effort should be directed at extracting as much of the ore as ‘clean’ as possible, delaying the appearance of waste/dilution and thereafter keeping the ratio of ‘dilution’ material to ore as low as possible for as long as possible. As dilution entry point, quantity and value of the diluting material affect the economics of the operation, it is important to be able to predict, with confidence, these parameters in order to calculate the ore reserves and planned tonnage and grade factors. This calls for an understanding of the material flow mechanisms in a SLC situation, which can be obtained from SLC design theory, current SLC mining practice and ideas ‘borrowed’ from block caving experience. Besides the large iron ore mines in Sweden that have used SLC as a primary mining method with apparent success for more than three decades, more recently a number of mines in Australia have adopted SLC as a primary extraction method and have achieved good results through improved layouts and better draw management. This has led to questions being raised regarding the original ‘bin theory’ assumptions and ‘classical’ SLC layout designs. ‘Classical’ theory and practice have dilution entry typically at between 20 – 40 per cent of ring tonnage draw and high cumulative dilution (20 – 40 per cent), with grade factors typically around 60 per cent, giving resultant low recoveries. More recent operations using ‘improved’ layouts claim to have achieved dilution entries in the 40 per cent – 70 per cent range, lower cumulative dilution (15 – 30 per cent), higher tonnage factors, with grade factors in excess of 80 per cent and recoveries around the 100 per cent level. The focus of this paper is, therefore, to examine and discuss the changes in SLC design and draw management that have contributed towards the ‘improved’ operational results. Emphasis is placed on the importance of delayed dilution entry and interactive draw in producing superior results.
BRIEF DESCRIPTION OF THE SLC METHOD The SLC method functions on the principle that the ore is fragmented by blasting, while the overlying host rock fractures and caves in under the action of mining induced stresses and
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gravity. It is a ‘top down’ method, with ore being extracted level by level working downwards through the orebody. The caved waste rock from the overlying rock mass fills the void created by ore extraction. Figure 1 is a schematic view through a SLC mine. The orebody is divided into sublevels at regular vertical spacings. A network of production drifts are developed across the full width of the orebody footprint at pre determined horizontal spacings. These drifts connect on one end to a slot drift and on the opposite end to a footwall (perimeter) drift which in turn connects via an access cross-cut to the main decline. The perimeter drift has a series of cross-cuts leading off it to ore passes. The ore passes end on a gathering level where the ore is transferred to the crusher and thereafter, conveyed or hoisted out of the mine. Where ore is trucked directly to surface from production levels, ore passes and crosscuts are exchanged for stockpile bays. The volume of ore immediately above each sublevel production drift is drilled with long holes in a fan or ring pattern. The drilling is undertaken as a separate operation, and completed well before blasting and loading commences (usually several months ahead). Blasting commences on the hangingwall or far end of the orebody and retreats towards the footwall or near side. A slot is blasted first and thereafter the rings or fans of production holes are fired and loaded out in succession. Ore extraction normally retreats on an approximately straight front, or series of panels each with a straight front, so that adjacent cross-cuts can be operated simultaneously. Production blasting and loading can take place on a number of sublevels simultaneously provided that the each successive level lags behind the level above by a distance not less than one to two times the vertical difference between the sublevels. Breaking the ore by blasting removes the dependency on ‘natural’ fragmentation as the mechanism for ore breakage shared by most other methods of caving. As mining advances downwards, level by level, through the orebody, the interface between the caved and solid rock above will move up toward the surface as the cave back constantly cracks up and fails providing rock to fill the void made by extraction of the ore. Passive support for the walls of the void created by removal of the ore in SLC operations is provided by the caved overlying material which functions as backfill. Provided that the orebody is not too deep seated, after sufficient ore is extracted the cave will propagate through to the surface. After this has occurred, with further ore extraction, the broken material is drawn down, reducing confinement on the upper parts of the cave walls, causing further failure of the cave walls. This process goes on until the orebody is depleted and the cave walls stabilise.
Advantages of SLC The following is a list of a few of the important practical advantages SLC has over other bulk mining methods:
• Top down and low capital: It is a top down approach to mining which means that the mine can get into production earlier than most other mining methods and at substantially less initial capital outlay.
• Flexible: It can deal with changes in the outline of the orebody identified by information gained during mine development. This feature enables sublevel caving to access ore in the deposit where other bulk mining methods may have more limited opportunities.
• Selective: If low-grade or barren zones are encountered the bulk of this unpay ground can be left behind, loading out only the ‘swell’ material. The ring pattern in such waste zones can be altered to create coarse fragmentation, to reduce dilution effects on subsequent levels.
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• Rock ‘factory’: In large deposits where conditions permit, development, production drilling, blasting and loading are carried out on separate levels or independent areas and are operations that are conducted independently of each other. Each different type machine ideally has several work areas available to operate in at any point in time. These factors allow SLC to be a continuous repetitive operation that can maintain an efficient mining process. It may be likened to an industrial ‘factory’ process where optimum utilisation of personnel and machinery is possible.
• Low-risk: Very little ore is at risk at any one time (a few hundred to a few thousand tonnes in an individual ring). ‘Lost’ ore can be recovered by ‘overdraw’ on the next level down.
• Safe: All work is carried out in well-supported drifts. CLASSICAL AND IMPROVED SLC THEORY The knowledge base for SLC is comparatively small as relatively few mines use the method. Most of the available SLC ‘theory’, generally found in mining handbooks and journals, comes out of Scandinavia and is many years old and mostly based on sand box model studies and classical ‘bin’ theory. Application of the ‘old’ or ‘classical’ theory in the design and operation of SLC mines over the years gained SLC a reputation for being a method that is development intensive and high in dilution. Janelid (1968) and Cox (1967) quote the following facts and figures:
• ore from development = 15 – 20 per cent of total production; • ore loss = 15 – 20 per cent; • waste dilution = 15 – 30 per cent (for iron ore 15 – 30 per cent by weight and 35 – 50 per cent by volume); and
• height between sublevels is 7-15 m and rings drilled vertically. However, good results from some recent Australian operations, and the changes the Scandinavians are making (high-level intervals) show that much improved results are possible with this method. These more recent results show the following:
• ore from development reduced to as little as seven per cent of total production (level intervals typically 20 – 30 m);
• cumulative dilution between 15 – 24 per cent; • metal recoveries between 90 per cent –100 per cent; and • dilution entry points in excess of 50 per cent. It must be stated, though, that data on draw behaviour is difficult to come by and in situations where ‘dilution’ carries relatively high grades it is difficult to assess ‘true dilution’. Here, ‘dilution’ refers to the material that is drawn which is not part of the current blasted ring. So the determination of true dilution from quoted results is difficult. Some of the Australian SLC mines quote very high recoveries and grade factors but these figures are not always adjusted for the grade in the ‘diluting’ material. A related problem is that many mines judge themselves against a reserve grade which already includes an estimate for mining dilution. However, the metal being recovered with respect to the tonnages being milled indicate that the ‘improved’ SLC operations are a lot more efficient than their earlier ‘classical’ predecessors. The main differences between the ‘classical’ and ‘improved’ SLC layouts may be seen by comparing Figure 1 (the ‘classical’ SLC layout) with Figure 2 (the ‘improved’ SLC layout). These differences are also shown in Figure 3. Essentially, in the improved layout, drill holes are shorter and side holes flatter. Most notable is that interactive draw is not possible in classical layouts as the draw cones on any particular level do not intersect. In the improved model interactive draw is possible.
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The ‘classical’ SLC model The ‘classical’ model of SLC (Kvapil, 1982,1992) was based on ellipsoids of motion and isolated draw, taking place at individual draw points. Even where straight lines of draw points were worked simultaneously, the layout did not permit interaction of draw columns to take place. Further problems associated with this model are that it ignores differences in fragmentation, material types, the significant weight of the rockmass within the cave that compacts the cave material immediately above and adjacent to production draw points, and the compressing action of the blast which also further compacts the waste behind the blasted ring. The conventional draw curves that most designers have used in the past are shown in Figure 4 (Kvapil, 1982). Although this figure shows dilution entry at around 50 per cent, conventional design typically has dilution entering at low extraction percentages (20 – 40 per cent) giving rise to the reputation of SLC being a high dilution method. It is agreed that in certain circumstances (ie drawing coarsely fragmented ore beneath a blanket of very fine waste material) dilution can enter very quickly (ten per cent of draw) and soon flood the draw point. However, with improved designs in many, more recent, operations the results show that dilution may enter at much higher draw extractions (well over 50 per cent) giving improved grade factors. The information used to justify this claim is scarce and not well documented, but the ‘improved’ model described below is based on the ‘belief’ that delayed dilution entry is possible and appears to be the basis for some significant changes from the conventional model of SLC.
The critical aspects of the ‘improved’ SLC model SLC Mines in Australia were breaking many of the classic rules but achieving better results than suggested by the ‘classic’ theory. The ‘new’ theory presented herein is based on observations from these mines in conjunction with experience from block caving operations. The ‘improved’ model is a common sense approach and is a reasonable basis for identifying the more critical aspects of SLC design. Many of the fundamental aspects of the ‘classical’ theory remain important but there are some new ones and some with a different emphasis in the ‘improved’ model. While some aspects are more significant than others in contributing to the overall results, what is clear is that the following are critical in achieving better results:
• Interactive Draw: The ‘improved’ layouts permit interactive draw to take place from a series of adjacent draw points that are retreated in a straight line. As material is drawn simultaneously or in rotation (taking a few buckets at a time from each) from these draw points, zones of low density are created between adjacent draw columns, which increases the width of the column of moving material and allows the rock to flow at much lower angles than is the case when draw points are drawn in isolation from its neighbours (Figure 5). In turn, this smoothes out the draw down profile of the ore/waste interface and delays dilution entry (Figure 6). (Draw data from two Australian mines showed large differences in extraction results between isolated and interactive draw). Interactive draw appears to be the single most significant factor affecting SLC performance.
• Loosened ore and compacted waste: Compacted waste material does not flow as readily as loose freshly blasted material. Sandbox modelling carried out by ‘classical’ theorists could not and did not take into account the compacting effect of the blast. The weight of the rock within the cave compacts the waste material and the action of the blast further consolidates this material immediately behind the ring. This more dense material will not flow as freely as the freshly blasted ore that has been ‘loosened’ by the blast.
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FIG 2 - SLC ‘improved’ layout.
• Temporary arching of the coarse material: Early research,
• Draw coverage: The wider the draw point development the
testing various size and shape particles in draw models (McCormick, 1968), showed that the larger the fragment size the wider the ellipsoid of draw became. What the ‘classical’ model failed to recognise (or researchers failed to identify) was the fact that the fragmentation size within a radial pattern (ring) blast generally varies from coarser towards the toe end and finer towards the collar end of holes (Figure 7) and that the coarser material forms wider ‘arches’ than the finer material. The wider temporary arching of the coarser fragments allows ‘loosening’ of the finer material below, which makes it flow more readily. Then, interactive draw limits the time this ‘temporary’ arching takes place and the ‘legs’ of the arch are loosened and fall in. This prevents or minimises the occurrence of major hang-ups and large voids into which surges of dilution may flow. The loosened ore, initially the fine fraction followed by the coarser fragments, is drawn preferentially to the compacted waste thereby delaying dilution entry and maximising ore recovery. Conversely, where independent draw is practiced, the ‘temporary’ ore arches hold up longer and allow dilution from behind the ring to surge into the void, causing premature dilution entry and reduced ore recoveries.
easier it is for the ore to flow into the draw point and the more even the ore/waste interface draw down, provided the draw point is worked across its full width.
• Differential fragmentation: Finer fragmented material flows more readily than coarse material. Dilution should be coarser than the ore (fine material can flow ‘through’ coarse material). If ore is finer than the waste it will be drawn preferentially delaying dilution entry. Drill and blast design can control the fragmentation of ore.
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• High-grade ‘dilution’: Having a mineralised envelope above and around the extraction volume ensures that dilution carries grade and therefore improves grade factors. In vertical, or very steeply dipping orebodies, a strategy of delayed draw on the first few levels sets up a thick high-grade ‘dilution blanket’ that ensures high recoveries throughout the life of mine. This finer blasted ore impedes the flow of the coarser waste rock through the blanket to the draw points. Unless very carefully controlled ‘over-drawing’ on the footwall side is practiced, this delayed draw strategy is not recommended in less steeply dipping orebodies as the high grade dilution material tends to ‘drag’ on the footwall side and may be cut off by waste from above.
• Dilution entry: The aim should be to preserve the high-grade dilution blanket as long as possible to keep ‘pure’ waste well away from an active draw point. Also, by having draw drifts as close as possible will lower the height of the interaction zone and improve interactive draw which in turn delays dilution entry (Figure 8). For block caving, Laubscher (1994) established that the dilution entry point was related to draw column height and height of interaction. A similar relation exists for SLC.
• Ground support: Ground support is installed primarily to maintain the integrity of the brow. The brow has to remain
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FIG 3 - Comparison of ‘classical’ and ‘improved’ SLC layouts.
stable for the short period of draw and subsequent charging up of the next ring. The support intensity will be much more than that required for normal tunnel stability since it has to accept the ring blast damage and has an extra degree of freedom.
• Blasting: The blasting must break the rock to the right fragmentation (uniform and not too fine) and without causing excessive damage, especially to the next row of holes. It should be sufficient to loosen the ore and compact the waste. Powder factors are generally more than 30 per cent greater than for unconfined blasting. Getting the right amount of explosive to the toe end of holes is critical. The overall objective for the SLC operation is delayed dilution entry. The items discussed above are very important when considering the layouts, conditions, equipment and procedures that will achieve or promote ideal draw conditions.
IDEAL LAYOUTS Everything in underground mining is a compromise between ideal layouts and what the orebody will allow you to do. In the case of SLC the compromise is one of:
• having smaller sized and wider spaced development (horizontally and vertically) in an attempt to lower operating costs, but in doing this, suffering loss of tonnes and grade through potentially earlier dilution entry and a lower grade factor; or
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• having more closely spaced, wider drifts to delay dilution entry and achieve better grades, but spending more on support and operating costs. Clarity is required about what the compromises are and their likely consequences. Often mining engineers focus on immediate development costs and are quick to conclude that less development is cheaper. In doing so they often overlook the effects this may have on head grade (which is mostly a function of planned and unplanned dilution) and loss of production. In other words, insufficient thought is given to the things that can go wrong that are brought about by non-ideal layouts. So it is vitally important to remember that dilution and production surety are the two most important engineering considerations and this must feature strongly in mine layout design considerations. There are a number of aspects that must be considered:
• Longitudinal versus transverse layouts: This decision is dependent on a number of factors such as: major stress magnitude and direction, dip and plunge of the orebody, orebody dimensions, and the frequency and orientation of major structures. As these vary from mine to mine no single prescription can be given. The following are more general comments. 1.
Transverse layout: This allows for more draw points and therefore more production flexibility. Tramming distances are shorter. Easier to ventilate than longitudinal layout, as main return raises are normally placed on the footwall or near side. Also cross-cuts are
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2.
shorter and therefore ventilation ducting losses are less. Ore pass layout simpler and shorter hauling distances on crusher transfer level. Long faces are more ‘relaxed’ and can be less stable; but if the face is retreated in steps of panels consisting of say five cross-cuts each, the stability problem will be minimised.
• Level interval: This is dependent on drilling accuracy (hole
Longitudinal layout: Longer tramming distances. Less production drives therefore less flexibility. More development and more complex layouts for ore passes and collection level development. Ventilation layouts more complex and requires very long ventilation ducts. If cross-cuts were used to allow for tramming to the footwall or near side, complicated ‘cross-over’ blasts would be necessary.
• Ring burden: In the past, the ring burden has been related to
• Size and shape of production drifts: The drifts should be as wide as possible to ensure good draw coverage. They should be as low as possible to allow for pillar wall stability and shorter muckpiles that enable easy access for charging. Also, have flat back drifts to widen the zone of moving material. This is an area of compromise as the stability of square brows is open to question and if arching the corners/shoulders results in more stable development then stability might take precedence.
• Drift spacing: Place production drifts as close together as possible. This is important in order to achieve interactive draw conditions (Figure 8). The spacing is dependent on pillar size and shape and loads. Pillar loads are high in the abutment areas and low beneath the cave.
deviation). Common sense suggests that the closer together the levels are the less likely that something will go wrong. Level interval has a significant affect on development cost per production tonne. With the improvements in drill technology the industry is rapidly increasing the level interval. Intervals of 20 – 30 m are quite common now. ‘dig depth’. This is doubtful though, as no buckets can dig anywhere near the back of a muckpile. The burden is dependent on the hole size, ring pattern and powder factor. The powder factor should be in the range of 0.9 - 1.1 kg/m3. The blast must minimise the incidence of freezing (an ever-present danger with choke blasting) and must ensure ‘loosening’ of ore and ‘compaction’ of waste. Overburdening is the common cause of ring ‘freezing’ and damage to the next row of holes. The design limits on the burden are probably around 3 m.
• Hole size: The largest possible holes are more economic, more accurate (so level intervals can be increased), are less likely to be closed (less re-drills) and can have larger burdens making it easier to access next ring of holes for charging. But they can do more damage to the next row of holes and can cause severe back-break affecting the collars of the next ring. ‘Sticky’ ANFO or emulsion can be used to charge larger hole diameters but the limit is probably 115 mm.
• Ring pattern: Fragmentation and damage are both determined by the ring layout. The hole spacing should be larger than the ring burden by a ratio of at least 1.3:1 at the
FIG 4 - Kvapil’s conventional model draw curves (modified graphics).
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FIG 5 - Draw points pulled independently or interactively.
toes of the holes. The ring pattern is also influenced by the rockmass conditions. In coarse blocky ground, smaller diameter closer spaced holes may be necessary to control fragmentation.
longitudinal layout), as it is very difficult to bring the rings through an intersection. In such cases multi-ring blasts are often necessary and there is a high frequency of freezing with multi-ring blasts in choke conditions.
• Ring inclination: Incline the ring towards the cave by 10° -
• Slots: A free-face is needed before ring blasting can start.
20° to adjust the shape of the moving material so that more ore and less waste is included in the mobile envelope. Inclining the rings also shields the draw point from the waste above and reduces the vertical pressure on the ore allowing it to loosen up more easily with draw. A further advantage is that it reduces the amount of back-break to the brow and makes it easier to access the next ring for clean outs and charging (Figure 9). There is very little information documented on ring inclination and the affect of different angles, but inclinations of 10° - 20° are quite common.
There are essentially four types of slot: (1) ‘individual’ slot with a slot raise at the end of each drift; (2) ‘continuous’ slot for a number of adjacent drifts with a slot raise at one end; (3) ‘slashing’ along the axis of the drift, or (4) ‘slashing’ along a perpendicular slot drift.
• Dilution blanket: Provided the ‘dilution blanket’ carries good grades, keep it as thick as possible to keep ‘pure waste’ well away from active draw points, and to keep grade factors high. The ‘dilution blanket’ can be drawn down towards the end of the life of the mine at marginal cost.
• Short hauls: Keep haulage distances for LHD as short as possible. Ore passes are ‘cheap’ in terms of cost per tonne compared to additional hauling costs, so have more well placed ore passes and shorter hauls.
• ‘Continuous’ development: Avoid having cross-cuts in combination with drives (this would only occur in
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MATTERS THAT CAN GO WRONG AND WHY THIS HAPPENS Before discussing the ideal conditions under which to operate SLC it is first useful to consider what can go wrong. The emphasis should be on achieving late dilution entry and high production rates per draw point. The items that have been identified as contributing to the success of the ‘improved’ SLC model can also detract from the results if they are not managed correctly. The areas that most commonly go wrong are listed below:
• Brow failure: Wedge failures in the draw point brow cause uneven narrow draw. Wedge failure or back-break also allows muck to rill out further than planned making it difficult to charge the next ring of holes. Shear failure of the brow area
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due to high horizontal stress will damage the holes making it difficult to charge and can necessitate reaming out of the collars.
• Ring freezing: Rings may ‘freeze’ due to blasting problems
other mineral, occurs and blasting causes excessive proportion of ‘fines’. Size of hole and type of explosive contribute to the amount of ‘fines’.
such as overburdening, or result from a delay in operations where the caved rock has compacted and over-consolidated, leaving no expansion space for the ring to break into (Figure 11). If a draw point has been left to stand for a long period of time and it is essential to loosen the ground by drawing off a few buckets of material prior to blasting the next ring.
• Pillar failure: High stresses, vertical structures running
• ‘Ribs’ between draw points: Unbroken toes on the side of
• Fines loss: This happens particularly where free gold, or
parallel to the pillar walls, or pillars that are too small can contribute to pillar failure. This may necessitate high support costs to remedy the problem. In highly stressed ground, development that is too far ahead can also suffer pillar wall failure due to unnecessarily long time exposures.
• ‘Bridges’: When toe end of holes remain unbroken, ‘bridges’ form (Figure 10). As it is bad practice to draw the draw point to a void, bridges can remain undetected until they have grown larger over a series of rings and thickened until they are down to the brow. Bridging allows waste to flow around the ‘bridge’ and cut off the ore remaining in the solid ‘bridge’. If bridging is serious it may require re-slotting to re-establish normal operations.
• ‘Walls’: ‘Walls’ are said to occur when the waste stands up as a ‘wall’ and does not flow down filling the void made by ore extraction. This may be the result of over compaction or the presence of ‘sticky’ material. A ‘Wall’ may also result when double rings are fired with the front ring freezing. Double ring blasts are not recommended.
the ring leaves unbroken ore between the drifts (Figure 12). This is a result of too large a spacing of the toes of side holes and/or incompletely drilled (or blocked) holes.
• Oversize fragmentation in pillar area: Hole deviation, misfires or incompletely charged holes especially in the toe end of the holes passing through the pillar area can cause significant oversize. This slows or interrupts the draw process and affects production rate. If the large rocks cause hang-ups it can result in premature dilution entry as waste surges into the void or waste fines flow down through the coarse ore (Figure 13). The coarse ore is cutoff by the waste and ‘lost’ into the cave.
• Wedge failure within the cave: Large, Steeply dipping continuous structures day-lighting on the top of the pillar and along the face of the retreating panel may result in large-scale wedge failure or sliding on the structure. This would shear drill holes and necessitate re-drilling and use of ‘sleeves’ to bridge gaps when charging.
FIG 6 - More even draw down of ore/waste interface with interactive draw.
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FIG 7 - Fragmentation profile through typical ring blast.
• ‘Overhangs’: Dipping ore, or domains of ‘strong’ rock that result in an overhang geometry, with delayed caving, can result in significant stresses being induced due to the overhang. This can result in crushing of development and/or loss of holes. This may happen where the retreat is from the footwall to the hangingwall of an inclined orebody, which makes this retreat direction undesirable.
• Incomplete slots: If the production drift starts with a slot that is not open to full height the rapid formation of bridges and ribs can occur. This would require re-slotting to be undertaken to get the area into production.
• Loss of holes: Drilling too far ahead and stress changes or relaxation can result in deterioration of drill holes (shearing or crushing). Daylighting of structure can cause block movement and cut-offs, loose material can rill into break-through holes and blasting damage from the previous ring can cause hole loss.
• Unplanned overcharging: Where high stresses have caused ‘dog-earing’ of blast holes or opening of joints/structures and the hole volume has increased overcharging of holes may occur unless sleeves are used. Overcharged holes usually result in excessive damage to the next row of holes.
• Delayed or no caving: If the plan area undercut by the initial production level is insufficient to cause caving of the overlying rockmass to occur, an airblast risk could develop if the matter is not correctly managed. Contingency plans, such
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as: increasing the undercut area, assisted caving through blasting or hydro-fracturing, or use of introduced fill materials will have to be considered. Ideal layouts for SLC have been reviewed and what can go wrong and why this happens during operations has been discussed. As mentioned earlier, mining methods are a compromise between what looks good on paper and what the orebody will accommodate. Determining what the orebody will allow is a ‘technical’ consideration. Achieving what the planning promises to deliver is a matter of ‘management’. The ‘ideal layouts’ are a result of the ‘technical’ considerations and ‘things that can go wrong’ are minimised through good operational ‘management’. A combination of a sound technical basis and good operational management should indicate the ideally achievable conditions for SLC.
IDEAL CONDITIONS FOR SLC What the orebody will allow you to do is a function of ‘mining difficulty’. Geometry, rock mass characteristics, major structures, stress conditions, grade distribution and rate of mining collectively determine what the degree of ‘mining difficulty’ will be for a particular orebody. For SLC these conditions should ideally be as follows:
• Strong rock: This enables the use of small pillars (small dimension between drifts), which allows good draw coverage and interactive draw that contributes to delayed dilution entry.
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FIG 8 - Effects of draw point spacing on improving draw characteristics.
• Competent rock: This is determined by degree of jointing and formation of unstable wedges. Competent rock implies few joints and strong joint surfaces (little infilling, irregular and rough surfaces), allowing for wide backs in production drifts. Where the presence of joints may cause wedge failures in one direction and not in the other, a decision may be made to align development in the favourable direction.
• Few major structures: If there are few major structures this will prevent massive wedges and/or hole cut-offs through relaxing and opening of the structures.
• Steep dip: Keeps the low-grade dilution source further away from the current draw points so most of the dilution comes in over the top of a mixture of ore and waste from caving at much higher levels than the current extraction level.
• Mineralised waste: It is a distinct advantage with caving operations to have the orebody surrounded by a mineralised envelope, especially if the grades in this zone are only just submarginal.
• Massive deposit: The orebody should present a sufficiently large footprint to ensure high production rates. Most dilution comes from the boundary between ore and waste - the more massive the deposit the smaller the proportion of material from the boundary. The development yield (ore recovered per waste development) is higher for massive deposits.
• Fragmentation: Ideally the caving waste rock fragmentation should be coarser than that of the blasted ore. The frequency,
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condition and orientation of joints and structures, the presence of micro-fractures and the intact rock strength will all affect fragmentation. A competent rock mass is usually the most suitable as larger, widely spaced blast holes will still result in good ore fragmentation but the caving waste will be very coarse. This results in blasted material being finer than the caved waste and a relatively even fragmentation for the ore with a minimum of both oversize and fines.
• Dry conditions: Dry roadways minimise tyre wear and a dry cave prevents accumulations of saturated fines/mud within the cave and the inherent risk of mud rushes. So the ideal is a minimum of groundwater and a positive drainage program ahead of the cave face.
• No ‘puggy’ or ‘muddy’ material: Ideally there should be an absence, or a minimum, of very weak or rapidly weathering material (especially clays or other puggy materials) to avoid problems of mud rushes. Also undesirable are ‘sticky’ cave materials with poor flow characteristics and over-compaction of the waste, which form ‘walls’ and hang-ups.
• Caving: The overlying rockmass must be able to cave freely when undercut over the mining footprint. Cavability is a function of rock mass conditions and the associated hydraulic radius (where hydraylic radius = area of footprint/perimeter of footprint) required to initiate, and sustain, caving (Figure 14). Massive orebody footprints generally provide a sufficiently large hydraulic radius that caving is usually not
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an issue. However, as the occurrence of voids can be an issue, through delayed draw and controlled draw, choke conditions must be maintained at all times. Caving is seldom an issue when starting immediately below an open pit, however, if the SLC starts well below surface then adequate footprint must be available to cause the overlying rock to collapse.
PROCEDURES AND PROCESSES TO MAXIMISE SLC EFFECTIVENESS Although mining engineers would like to have ‘ideal’ mining operations, mining is never ‘ideal’ and compromises are inevitable. There are a number of processes and procedures that can be used to help overcome the possible consequences of making various compromises and to enhance the effectiveness of SLC. There are a number of ways in which design (technical) and attention to detail (management) can be used to reduce uncertainty and maximise the effectiveness in sublevel caving, including:
are mined to specification. The excavation process can cause considerable damage through over-break, making cross-cuts less stable and more difficult to support. Therefore, care must be taken in the design and management of the development excavation process, including:
• Perimeter holes sufficiently close to cut a regular profile and allow for post-splitting.
• Perimeter holes to be charged with smooth blasting agents (definitely not ANFO) and timed to effect a post-split. If necessary the next row of holes may also have to be cautiously charged (again avoiding ANFO).
• Hole spacing to allow good breaking angles for all holes (at least ‘right-angled’ triangles). This reduces confinement and minimises damage through back break.
• Ensure correct timing of shots to prevent out-of-sequence detonations.
• Use tested and proven standard drilling patterns. Do not let jumbo operators ‘individualise’ their patterns. Variations may introduce unnecessary errors.
• Mark the patterns in detail on the face for the jumbo operator to follow.
Development excavation As the size and shape of the production excavations has a large impact on operational efficiencies it is important that these drifts
• Ensure jumbo operators use auto-parallelism wherever possible and keep the ‘look-out’ angle on perimeter holes to the minimum necessary.
Fig 9 - Inclining rings to give better production results.
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Ground support In SLC the support installed in production crosscuts has two primary functions:
• to protect personnel and equipment; and • to provide support of the brow to facilitate optimum operating conditions (draw profile and access for charging). The support has to work when the brow has freedom to move and has to resist strong blasting disturbance. Support requirements for the two above functions can be quite different. Establish whether precautionary or structural support is required. Precautionary support can be split sets and mesh, whereas structural support would normally include grouted bolts with strong plates (length of bolts between 2.4 m and 3.0 m), increasing support intensity through use of straps and/or fibrecrete (reinforced shotcrete). The structural support is generally concentrated at the brow position although, under certain circumstances the walls may need more support than the backs. In high stress conditions (abutment areas) bolting and/or cables with strapping might be required to hold pillar walls. Use of mesh as an areal support in production drifts should be avoided if possible, using fibrecrete instead, as the cutting of mesh at blast time and removal and handling of mesh in the draw muck pile is difficult and slows down production. Mesh causes problems on grizzlys, in the crusher and at conveyor transfer points. A point worth noting is that collaring of blast holes is far easier and more accurate when fibrecrete/shotcrete is used, and hole accuracy in SLC is very important.
The objective of the support is to retain brow shape, maintain stable collars, reduce back-break and produce safe charging conditions.
Ring blasting The objectives of ring blasting include: no misfires or partially charged holes (that cause bridges, ribs or frozen rings), loosening of the ore and compaction of the waste, good fragmentation that aids mass flow conditions and high productivity, and limited damage to brows and next row of holes.
• Spacing and burden: The blast layout is designed to minimise oversize that interrupts or slows draw. Oversize is a function of draw point width, LHD bucket size and grizzly dimensions. Shifts in fragmentation distribution are achieved with varying the burden and spacing of holes: finer fragmentation is obtained using smaller burdens and larger spacings and longer delays (50 millisec or greater between holes). The burden cannot be too small or difficulty will be experienced in accessing the next row of holes for charging. This requires a compromise to be reached.
• Drill and blast pattern: The lower side holes should be ’flattened’ as much as possible, but not too low that they choke and misfire and create ‘ribs’ (ie not less than 50°). This is done to reduce the length of the longest central holes and minimise the amount of drilling through the damaged pillar zone (Figure 15). The problem with relieving the flatter holes
FIG 10 - Formation of ‘bridges’.
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FIG 11 - Freezing of blasts.
is reduced by ‘interactive’ draw. Rustan (1982) showed that the limiting flow angle (a) can be roughly calculated: a = 45o + φi / 2 Where φi = friction angle, determined by fragmentation, rock strength, cohesion, degree of packing of rock and is a measure of the friction between the pieces of rock. Dry conditions (less cohesion) and interactive draw (loosening) will lower the angle.
• Timing of blast holes: There should be no overlaps in
• Hole deviation: Minimising hole deviation/deflection is
The objective of draw control is to draw as much ore and as little waste as possible. In other words, to maximise the draw before dilution enters the muckpile, to delay dilution entry as long as possible and to quickly correct anything that contributes to early dilution entry. The objective is achieved by giving attention to the following:
critical. This necessitates larger diameter tube drilling rather than smaller diameter rod drilling. Extra expense for stiffer drill assemblies (tubes), laser set-up, good roadbed to assist with set-up, will always pay dividends. As mentioned earlier, starting holes through fibrecrete/shotcrete also increases collaring accuracy. Hole should be surveyed for deviation and any hole more than three per cent off line should be re-drilled. Hole checking is critical, particularly to avoid the occurrence of bridging.
• Charging pattern: This must minimise sympathetic detonations, damage to the brow and damage to the next ring. This is achieved by keeping charging density in all parts of the ring pattern to below 1.5 times the design powder factor. Powder factor at critical toe end of ring should not be less than 0.9 – 1.0 kg/m3. Ideal powder factor for SLC is between 0.9 - 1.1 kg/m3. The uncharged collar lengths should not be less than 0.6 - 0.8 times the burden. If ledges are formed then cautiously blast the brow area with decoupled charges. The blast must be able to create swell by compacting the ‘choke’.
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timing of holes, allowing individual detonations with good free faces to break into. Delays should be sufficiently long (not less than 50 ms between holes) to get separation and movement.
Draw control
• Interactive draw: Ensure that panels of 4 - 6 draw points are retreated in a straight front and are drawn simultaneously. Draw between 50 – 150 tonnes from a draw point (always an even number of bucket loads so that loader operator draws evenly from both sides of the draw point) and then move to the next drift. In vertical or very steep orebodies, draw equal amounts from all drifts and rotate from draw point to draw point to ensure interactive draw conditions are maintained. Interactive draw is the single most important factor affecting SLC performance. In inclined orebodies, it may be advisable to ‘over-draw’ the footwall draw points and ‘under-draw the hangingwall side to get optimum ore recovery and minimal hangingwall waste dilution.
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FIG 12 - Solid ‘ribs’ due to toes of side holes not breaking.
• Even draw: Draw from right across the muck pile, do not draw continually from one side and do not draw around hang-ups or large blocks in the muckpile. Loader operator should aim to take alternate loads from opposite sides of the draw point.
• Prompt secondary breakage: As soon as oversize is encountered, if it cannot be quickly loaded away to a breaking bay, have it broken in the draw point. Do not delay in removing or breaking up the oversize and do not attempt to continue drawing around the oversize, or disruption to interactive draw and even draw will occur.
• Break hang-ups instantly: All forms of hang-ups, whether they are a result of very large oversize, bridges or ribs, should be immediately broken with hang-up drilling equipment. Hang-ups are not self-correcting and ignoring them by simply blasting the next ring usually makes the problem worse. Where extraction is interrupted for any length of time, the blasted ore (and waste) are likely to settle and compact, causing draw problems later. Total ore extraction from any blasted ring should be completed before any planned long interruptions.
• Production control: Production planning engineers will calculate the available tonnage and required tonnage factor for each blasted ring. The tonnage factor should take into account any ore left from the level above, especially if a high grade dilution blanket has been created and is being
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maintained above draw levels. A ‘bucket’ sheet should be issued to the loader operator that tells the operator how many buckets he can draw from a draw point before detailed visual/assay sampling takes place (this will be related to the desired tonnage factor). It is common to use load cells and electronic monitoring systems to measure the mass and quantity of loads taken from any draw point.
• ‘Visual’ checking: Geologists or draw control personnel should inspect muckpiles regularly (at least twice per shift) to assess proportions of fresh ore, ‘old’ ore and waste in the muckpile. In many instances ‘old’ ore from above, but not broken in current ring, may not be able to be distinguished from freshly broken ore.
• ‘Shut-off’: If drawing to a ‘shut off’ value, careful checking must take place from approximately 90 per cent draw to ensure that grade of muckpile is above shut-off value. Then pull to whatever extraction results in shut-off. If a dilution blanket is being preserved for final extraction, then a set tonnage factor will be applied and the draw point stopped for blasting of the next ring when this percentage of draw is reached. When mining to a shut off value, use marginal costing to continually adjust the shut-off value/grade and check that predicted average grade is being achieved. Early dilution entry and high dilution grades can result in very high extractions and recoveries over 100 per cent but at a low average grade.
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• Monitoring draw: It is most important that the draw is continually monitored by the geology or draw control personnel (once every two to four hours during draw) to check the point of dilution entry and how the proportion of dilution changes with extraction. ‘Dilution’ is anything that was not in the freshly broken ring. This monitoring can only be done effectively if fresh and old ore material can be distinguished and, if not, then simply distinguish between ore and waste, if this can be done. All draw information should be recorded so that, over time, a draw control database can be established and used for future draw estimation and grade control purposes. Use is made of a draw nomogram to calculate tonnage and grade to shut off (Figure 16). The nomogram is initially generic but with time and monitoring feedback it will become ‘real’ and representative of local conditions. Therefore, monitoring is essential.
• Grade control: Mineral characteristics and distribution can influence recovery, particularly where a significant proportion of mineral reports in the ‘fines’ in which case higher losses can be expected. Fines generation is largely controlled by the geotechnical characteristics of the rockmass, drill hole diameter and type of explosives used. Regular muck pile sampling should be undertaken and grades/values reconciled against planning figures. Over drawing can quickly reduce grade factors.
Floor conditions Cross-cut floor conditions should be dry and smooth with effective and rapid drainage. This can be achieved by laying an engineered road base (compacted gravel surface) kept in good condition by regular grading.
Excavation repairs Watch ground conditions and effect repairs as soon as possible before it deteriorates to becoming a major problem. Pay particular attention to walls. If repairs are becoming frequent increase the support density. Rehabilitation of development is expensive, unsafe and interrupts production.
Drilling ahead Drill ahead as far as possible (have at least three – six months drilling stocks), but if re-drilling becomes a major interference with progression of the blasting face then drilling might have be done on a ‘just-in-time’ basis. Where major structures are intersected by production drifts, leave rings either side of the structure for ‘just-in-time’ drilling. This will reduce the loss of holes due to potential relaxation and opening of structure.
FIG 13 - Oversize in the pillar area due to poor drill and blast.
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FIG 14 - Stability diagram (Laubscher, 1994).
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FIG 15 - SLC drilling layouts.
EVALUATION (BUDGETING AND FORECASTING) In a well-designed and efficiently managed operation, what will SLC deliver? Like all mining methods, the costs per unit are easy to calculate. The difficulty is in predicting the productivity per unit (primarily the production rate per draw point) and the head grade. The latter is a function of planned and unplanned dilution, which in turn is dependent on the point of dilution entry, the degree of mixing, and the dilution grade. To add to this difficulty, there are very few SLC operations worldwide and the results of one Australian operations seem to be ‘too’ good. However, this should not detract from what is potentially a very productive and cost-effective method. It is a method that can deliver the lowest cost-to-metal under certain circumstances. In spite of the lack of good data, estimations and predictions must still be made. With regards to grade prediction, a key element is to fully utilise the benefits of interactive draw. A suggested draw model is illustrated in Figure 16. This has been based on models developed for block caving (Laubscher, 1994) and from previous experience with SLC operations. The model shows the proportion of ore (below the line) and dilution (above the line) as the extraction increases (along the x-axis). For example, on the 50 per cent dilution entry curve, at an extraction of 80 per cent the material drawn consists of 62 per cent fresh ore and 38 per cent dilution material. It can be seen that very high extractions can be achieved dependent on the shape of the curve and the grade of the ore as well as the grade of the dilution material.
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Ideally we should also have a model that estimates the grade of the dilution material, which is a mixture of waste, mineralised waste and ore left behind from each recovered ring. A methodology to estimate the contents of the ‘dilution bin’ is illustrated in Figure 17. ‘Dilution bin’ is a label given to the accumulation of dilution material (ie material not sourced from the current rings being blasted and drawn) that is available to be drawn as dilution and which gathers immediately above the level containing the rings of ore being extracted. While the material in the ‘dilution bin’ at any point in time will generally be a mixture of coarse and fine fragments, as the mine gets deeper and as the cave matures, the fine material within the cave will migrate preferentially down through the coarse material, forming the larger proportion of the contents of the ‘dilution bin’. What the model illustrates is a ‘dilution bin’ that represents the average grade of the mixture of pure waste, mineralised waste and the ore left from each recovered ring. The amount of ore that is left behind after each ring is drawn to shut-off will depend on the effectiveness of the draw process. Less effective draw, say the 30 per cent dilution entry curve in Figure 16, will leave more ore from the current ring unextracted. But as more and more of the remnant ore enters and is mixed in the dilution bin, the grade of the bin may increase. This will also depend on the grade of the waste that is entering the bin from above.
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FIG 16 - Draw nomogram for calculation of extraction tonnage and grade.
What happens in practice is that less effective draw can result in very large extractions (more mixing of ore and waste but the resulting mixture still above shut-off grade). Very high recoveries of the mineable metal can be achieved (well over 100 per cent) but at lower overall grade factors. The dilution bin representation in Figure 17 will therefore depend on the actual in situ resource grade, shut-off grade, the extraction factor and the grade of the dilution before the ring was drawn. This bin is progressively topped up with waste from above and ore from below (remnant ore from drawn rings). Figures 16 and 17 are simplistic representations of what will occur in practice, but they do give a reasonable first estimate and illustrate the sensitivity of the results to the efficiency of the draw control and the presence of mineralised waste. It must be remembered that the models are initially ‘common sense’ but with time will become ‘real’ if a cave is correctly monitored and data fed back to fine tune the models. Monitoring is essential in this regard. This methodology was used on a recent project and allowed reasonable sensitivities to be run. The results from the model agree reasonably well with the achievements of two other existing operations as can be seen in Figure 18. However, neither of these operations (Mine A and Mine B) are technically or operationally ideal. The curves for Mines A and B are for all draw conditions (interactive and independent) and from both longitudinal (narrow) as well as transverse (wider) areas. Also, the three ‘Kvapil’ lines (constructed from Kvapil 1982 draw curves) represent what ‘classical’ theory would
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consider as being good, average or bad results. Therefore, it is believed that a well-engineered and operated SLC in ideal circumstances (geometry and rockmass conditions) can deliver better results than predicted by the model. Cut-off grades are determined by examining the cost-to-metal. The costs used are the full operational costs including capital. The shut-off grade should be continually re-estimated using marginal costs to ensure that the grade is sufficient to make a contribution to depreciation and amortisation. It is not advisable to use marginal costing for development and stope boundary planning as the metal price is not normally known this far in advance. It is much better to work to a strategic cost-to-metal. Hedging programs may change the cut-offs and shut-offs but should not change the common sense approach of planning on cost not on revenue. To estimate production rate, the draw point is used as the key operational indicator. The number of available draw points will determine the overall production rate. To determine the required number of production drifts, one must estimate the time required for each dependent operational element at the draw point. This includes drilling, charging, blasting, brow repairs, re-drilling, mucking and any secondary breakage and hang-up clearance. One must also consider the influence of keeping the face flat and achieving interactive draw. With all of these conflicting activities, it is difficult to estimate the ‘long-term’ production rate from a set (or face) of interacting draw points.
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SUBLEVEL CAVING — TODAY’S DEPENDABLE LOW-COST ‘ORE FACTORY’
FIG 17 - Mechanics of dilution bin formation.
The experience at SLC operations where conditions are reasonable and the plans and procedures are appropriate is possibly 500 - 600 tonnes/day for each draw point. Clearly, the haul distance should be kept as short as possible. In longitudinal retreat this conflicts with minimising the number of access cross-cuts. Although the fragmentation from SLC is usually very good, there may be problems with grizzlies, ore passes and chutes. The production LHDs should be as large as possible. The fragmentation, road conditions and haul distances should be such that a LHD should be capable of at least 25 - 35 buckets per operating hour assuming a straight run into the muckpile and the bucket is filled in a single movement. A large LHD will require some four to five draw points to keep it supplied with broken muck. If interactive draw is also considered and the draw is across a group of four to five draw drifts then at least ten draw points may be required per LHD (five for loading, five for drilling, hole clean-outs and charging). If trucks are being loaded, then there might be only one LHD operating in a face/panel. In such a case, the production rate will be a function of buckets per hour for that LHD. The overall maximum production rate of any mining method is very difficult to estimate on paper, but typical rates-of-fall through the deposit under good conditions for a large SLC operation would be in the region of 65 m per year.
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CONCLUSION In the past many people have considered SLC to be a development intensive, high dilution, and low-recovery mining method. However, results from recent operations, where the ‘new’ SLC model has been applied, tend to contradict this assumption. The clear advantage of SLC is that it is a very predictable ‘factory’ type method with high production potential, reasonable costs, ‘top-down’ approach (low up-front capital and quickly into production) and very little ore at risk at any one time (a few thousand tonnes in an individual ring). Currently, most massive deposits are either block caved or if they will not cave readily they are open stoped with cemented fill. Sublevel stoping is generally not considered, or if considered, rejected based upon dilution and recovery concerns. The response being that recovery will be higher and dilution lower with a filling method. This is partially correct but recoveries are usually well below 100 per cent and the dilution often runs at over 15 per cent. A mature SLC can achieve recoveries in excess of 100 per cent and grade factors well over 80 per cent (the proportion of pure waste is less than 20 per cent). This assumes that there is mineralised waste around the ore (not unusual) and that the dilution grade steadily increases. The latter is the expected attribute of SLC which must be maximised. In addition, SLC can be very productive in terms of rate-of-fall through a deposit. It is generally double that which can be achieved with filling methods.
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G BULL and C H PAGE
FIG 18 - Comparison of cumulative dilution.
In a recent project, SLC was shown to exhibit very superior economics over the more conventional methods. Recoveries of over 100 per cent, grade factors over 80 per cent, and operating costs less than $9/t were achieved. Only a very few orebodies are ideally suited to SLC. They need to be:
• strong and competent, • have a large footprint with a very steep dip, and • preferably have mineralised waste If these conditions apply, the only methods that can compete on a cost-to-finished-metal basis are possibly block caving and partial extraction. For SLC the major future advances are with (1) interactive draw, (2) the understanding of what contributes to the success of the method, and (3) ensuring that this success is achieved.
REFERENCES Cox, J A, 1967. Latest developments and draw control in sub-level caving, Trans IMM, Sect A, Vol 76, ppA149-A159. Janelid, I, 1968. Sub level caving: how to use it, what are the advantages, problems, World Mining, September, pp76-78. Kvapil, R, 1982. The mechanics and design of sublevel caving systems, Underground Mining Methods Handbook (Ed: W A Hustrulid), pp880-897, (SME). Kvapil, R, 1992. Sublevel caving, SME Mining Engineers Handbook, 2nd Edition (Ed: H L Hartman),Chapter 20.2, pp1789-1814, (SME). Laubscher, D H, 1994. Cave mining – state of the art, The Journal of the South African Institute of Mining and Metallurgy, October:279-293. McCormick, RJ, 1968. How wide does a draw point draw?, E/MJ, June:106-117, Rustan, A, 1982. Compendium in sublevel caving operation, Division of Mining and Rock Excavation, University of Luleå, February.
ACKNOWLEDGEMENTS The authors wish to thank the management of SRK Consulting for the opportunity to publish this paper and for encouragement given along the way. The excellent graphics are kindly produced by Kerry King Graphics.
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Gravity Flow of Broken Rock — What is Known and Unknown A Rustan1 ABSTRACT The knowledge in the field has been reviewed by literature studies, and using experience from personal research in sublevel caving. There is a lot of material regarding gravity flow of fine grained materials in bins, but the theory of flow for these materials are normally not applicable for coarse blasted rock. This study was therefore concentrated towards coarse rock fragments created after blasting in orebodies. The goal for this study was to make a report of known knowledge, and identify what we have to learn more about. The Bergmark-Roos theory, which was a paradigm shift in the science of gravity flow, is today the best mathematical theory for calculation of the drawbody for a homogenous material or when we wish to simulate the mean drawbody for a large number of draws. The Bergmark-Roos theory must however be improved, so it can take into account the flow pipe effect. When this is done it will be possible to construct a 3D computer model for calculation of ore extraction versus amount of loaded ore and waste. It is also necessary to find the 3D shape of the flow cone to be able to calculate the 3D ore rest that normally develops on footwalls. Calculations should be done for the draw on several sublevels including determination of the ore losses and temporary ore rests. First when this computer model is developed optimisation of ore recovery and waste rock dilution can start for a given orebody. The author points out the importance for international co-operation and standardisation of methods for determination of material properties. More field tests and model blasting tests should also be done. When enough knowledge is gathered, a handbook of the optimisation of sublevel and block caving should be written.
INTRODUCTION Despite enormous amount of research in the gravity flow field, 3D simulations of ore recovery and waste rock dilution can still not be done for conditions in a specific mine. Using model tests, only the outcome for a certain parameter combination can be determined. If any parameter like sublevel height or distance between cross-cuts are changed, several new model test have to be run until the optimum burden can be determined. This paper is based on a literature review and experience from personal laboratory and field research in sublevel caving starting in 1965 and the paper makes an overview of some existing relevant knowledge of gravity flow of coarse rock and model tests to simulate the flow of coarse rock in sublevel caving and block caving. The literature review includes the most important findings in the field of gravity flow starting from the 1940s. The result from the literature review will be presented in chronological order in each chapter and some comments are given to the findings. In the section entitled: ‘Important Questions for the Design of Sublevel and Block Caving’, I will consider some crucial questions to answer concerning gravity flow and in the section entitled ‘Material Properties and Quantification of Physical and Mechanical Properties of Blasted Or Cave Rock’ is shown how to characterise the material properties of coarse rock. It is concluded that it is a mistake to introduce constant properties in the calculation models for gravity flow, because the material properties vary within the gravity flow, both in space and time. However for the relative comparison of different materials some recommendations for international standardisation of properties are given.
1.
CENTEK, Luleå University of Technology, Sweden. E-mail: [email protected]
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The history of the drawbody shape is given under the heading ‘determination of drawbody shape’, and the most relevant theory for the shape to be used today is presented. The upside down drop shape is the accepted shape today. Calculations of the most important parameters for the design of gravity flow, the critical and maximum flow width are given under the heading ’calculations of the maximum and critical flow tube width in bins’. The paper ends with the conclusions that a 3D computer program should be developed where different shapes of drawbodies can be simulated for calculation of ore loses and waste rock dilution for orebodies with defined properties of the ore, and geometry of the hanging and footwall. The computer program can then be used to optimise the different parameters. The computer program must, however, always be calibrated to the field by full-scale tests.
IMPORTANT QUESTIONS FOR THE DESIGN OF SUBLEVEL AND BLOCK CAVING Before we start the review of the literature, we must define the critical questions for which answers still are needed. There are three important questions both in sublevel and block caving. 1.
Quantification of physical and mechanical properties of blasted or caved rock?
2.
How to determine the shape of the drawbody at different draw heights and different material properties?
3.
How to calculate the maximum flow zone width at different material properties like size distribution, shape- and surface friction of fragments?
The drawbody and other terms used for gravity flow in mining are defined in Figure 1. It is assumed that the terms in Figure 1 are known to the reader, and if it is not, the reader is recommended to read the paper by Janelid and Kvapil (1966).
FIG 1 - Definitions of commonly used terms in the field of gravity flow.
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MATERIAL PROPERTIES AND QUANTIFICATION OF PHYSICAL AND MECHANICAL PROPERTIES OF BLASTED OR CAVED ROCK A full description of a material, in our case blasted or caved rock, includes a large amount of parameters, that according to Wood (1986) can be grouped into three main groups: 1.
discrete particles,
2.
bulk assembly, and finally
3.
In practical work, however, seldom all these parameters are quantified, and therefore concentration must be focussed on the most relevant physical and mechanical properties. The most important parameters for the critical flow width Dcrit are as follows: Dcrit = f ( k0 − 100 , L, µ, A, ρ, σ, C , m)
(1)
where k0-100 is the symbol for the size distribution of fragments, L is the shape factor preferably defined as a sphericity, see Wahlström (1955), µ is the surface friction of the fragments, A is the attrition, ρ the density of the fragments, σs is the shear
other influences. The factors are listed in Table 1.
TABLE 1 Properties and processes determining the characteristics of bulk solids during storage and flow (after Wood, 1986). Discrete particles
Bulk assembly
Other influences
Attrition
Aeration and fluidisation
Contamination
Degradation
Agglomeration
Crust/pan formation
Elasticity
Angle of repose
Flow rate
Electrostatics
Angle of slide
Geometry of container
Hardness
Boundary effects
Geometry of flow path
Magnetism
Bulk density
Loading method
Mineral composition
Cohesiveness
Pressure
Relative density
Compaction and consolidation
Rate of loading
Shape
Contact stresses
Rate of unloading
Size
Distribution of solids
Time of storage
Surface chemistry
Distribution of void space
Vibration
Surface energy
Electrostatic effects
Volume changes
Surface roughness
Fabric
Wall material
Flow potential
Weathering
Internal friction Interparticle forces Moisture content Packing Particle to particle interactions Permeability Pore pressure Porosity Segregation Shear strength (Fragment) size distribution Specific heat Stress distribution Temperature Thermal conductivity Type of interstitial gas (coal) Unconfined yield strength Vibration kinetic energy distribution Note: the most important factors have been marked with bold by Agne Rustan.
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strength and C is the cohesion of the bulk material and m is the moisture content. All symbols except A were taken from Rustan (1998). Large fragments or fine material with high cohesion are normally responsible for the hang-ups. The shape and the surface friction of the fragments determine the mobility of the material, and the attrition determines the change of fragment size distribution during the draw. The compact density influences the flow velocity of the material compared to the surrounding material, which normally has a lower density. The breakage of hang-ups is mainly caused by shear forces, and therefore the shear stress and cohesion is of fundamental importance. These are therefore the parameters, which should be included in an international recommendation for classification of caving materials. Too much research has been devoted to try to find some properties that characterise the material like internal angle of friction, limit border angle, angle of repose at dumping or loading, etc to use for calculations of gravity flow. The problem is that in the gravity flow of coarse rock, the properties like swelling, packing, porosity vary so much in space and time that it is not possible to characterise the material with a constant property of the rock, and therefore simulations using constant properties will always fail. It can only work for a very fine and mono grained material.
FIG 2 - Photo taken from a sublevel caving blast at a scale of 1:75. Observe some white limestone fragments on the muck pile directly after blasting. After Rustan (1970).
Fragment size distribution Only one reference is known where sieving of full-scale fragments from sublevel caving blasts has been done. A representative volume of the blast was sieved, and the boulder dimensions were measured manually, and their volumes estimated. The work was done by Maripuu (1968) at Luossavaara-Kirunavaara AB (LKAB) in Kiruna magnetite orebody. The mean fragment size, k50, for the ore was determined to ~100 mm. The knowledge of the complete size distribution was very valuable for model tests performed for LKAB later on.
Accurate simulation of fragment size distribution by blast model tests in sublevel caving Rustan (1970) simulated blasting in sublevel caving by models at a scale of 1:75. Blasts were simulated in one cross-cut. Five burdens after each other were blasted and loaded. Rustan devoted lot of time in finding a model material, that would give a scaled fragmentation of the full-scale fragmentation. To achieve this, it was necessary to insert artificial weakness planes in the model material, because without weakness planes, the middle size fractions were lacking, see Rustan (1990). The final recipe developed by Rustan was two-thirds of coarse magnetite, one-third of fine magnetite with the ambition to create as high density as possible for the artificial orebody. The magnetite was mixed with cement, water and crushed microscope glass plates. The reason to use glass plates was to achieve more middle size fragments. The microscope glass plates acted as natural weakness planes in the magnetite. In the LKAB Kiruna and Malmberget magnetite orebodies there are namely three dominant sets of weakness planes. After blasting, (see Figure 2), the loading of ore (magnetite concrete) was started, and the magnetite was separated by a magnet from the white waste rock (limestone) to be able to determine the ore recovery and waste rock dilution. In Figure 3 a blasted fan is shown after blasting. Observe how real the surface looks if compared to a real ring blast underground. Many of the phenomena observed in the full-scale could also be seen in the model, for example backbreak and variation in waste rock content during the extraction, etc. It is therefore
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FIG 3 - Sublevel caving front after blasting the burden against waste rock in a model at a scale of 1:75. After Rustan (1970).
believed that model scale blasting is the nearest we can come to the real process in full-scale. Model tests with loosely poured monograined fragmentation should not be used at all. If blasting could not be done in the laboratory, it is better to have a scaled down fragmentation instead of mono grained material in caving models. Within some burdens, extra holes where drilled, and markers inserted, and thereby the origin of the ore falling down into the drift directly after blasting could be identified. Through archaeological digging in the blasted ore and caving material, the position of each marker after blasting could be found. As a matter of interest some waste rock fragments were found on the muck pile directly after blasting, (see Figure 2). The only
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explanation to this phenomenon given so far is that the dynamic movement of the burden towards the waste rock during blasting, opens up a slot between the burden and the front in which waste rock can fall down from the top of the blasted ore. Discussion with Stazhevsky in 1995, revealed that Russians like Swedes have made observation of waste rock pieces on the muck pile directly after blasting in full-scale. It is therefore believed that small-scale blasting is the best method to simulate real sublevel caving in physical model testing.
Fragment size distribution in the Chinese Hebei magnetite mine In the Hebei Mine, full-scale test were done with the goal to determine the drawbody shape. The size distribution of the blasted magnetite ore was measured and presented, see Gustafsson (1998). k50 was very close to ~100 mm, the same as for LKAB in Kiruna. The method used to determine the size distribution is, however, not known. Probably sieving was used.
Full-scale tests to determine the shape of the drawbody Full-scale testing and model blast tests are the two best methods to learn to understand gravity flow. The number of full-scale tests are, however, restricted due to the high cost. To reconstruct the shape of the drawbody it is necessary to insert markers in that burden, which is under study. A representative marker was developed by Fröström (1970) for model tests, and a good marker type was also developed for the LKAB Kiruna tests, see Gustafsson (1998). The latter marker will be described further on.
Strategy for selection of proper markers to follow the flow
Fragment size distributions were also determined in several block caving mines, see Juhlin and Tobie (1973). The mean fragment size (k50) for some block caving mines varied from 100 mm for Mather Mine to 1500 mm for Grace Mine. Both mines are located in the USA.
Investigations were undertaken by Fröström (1970) at the Royal Institute of Technology to find the best type of markers. He found that a brass ring of a wedding ring size was the best marker to follow the gravity flow in a 1:50 scale. The diameter of the brass rings were 20 mm. If the marker is too small it will move much faster to the draw point due to segregation. If it is too big, it will move too slowly, and it will be difficult to get enough measurement points for reconstruction of the drawbody. If it is too light, it will have a tendency to float on the heavier magnetite used in the tests, and arrive much later to the opening. In the model tests a wooden disk for example arrived, in mean, 33 scoops later than marked stones placed nearby the markers original position. Wood is floating in a denser material.
Shear stress and cohesion
Grängesberg tests in magnetite ore
Central to the breakage of loose material is the shear stress and cohesion determined in shear strength tests. Three examples of such kind of studies for coarse rock are as follows; Fumagelli (1969) tested materials for rock fill dams and Marachi (1972) presented strength and deformation characteristics of rock fill materials for dams. Forsman and Pan (1983) determined the shear properties of magnetite in the size range 1.5 - 80 mm at porosities varying between 29 - 33 per cent. Interesting in the last tests was a decline of shear strength at larger normal stresses (0.8 - 1.2 MPa). A decrease in the order of 12.5 per cent was measured.
Janelid (1973 a and b) reported the result from the first full-scale tests made at Grängesberg in Sweden. The sublevel height was 10 m, the drifts were small 2.5 × 2.5 m, and therefore these tests have a limited value for large-scale sublevel blasting. The draw height was 19.4 m. Totally 14 rounds and 13 000 rubber hose markers were used in the tests. Markers were placed only in the upper part of the expected drawbody. Distance between marker hole fans was 0.3 m. The markers were made of garden water hose cut in length of 30 cm, and with a diameter of 19 mm. The relation between draw height and flow width was almost linear, (see Figure 4).
Fragment size in block caving mines
7
The bulk density of magnetite ore and its dependency on the size distribution of magnetite and waste in different combinations was examined in model tests by Rustan (1970). The purpose of his investigation was to examine if the bulk density of ore could be used to measure the magnetite content in the LHD loading bucket in sublevel caving. The results from different combinations of size distributions were reported. The bulk density varied in limestone from 1.46 to 1.74 t/m3 and in magnetite ore from 2.54 to 3.30 t/m3 dependent on size distribution and packing of fragments. The specific gravity for limestone was 2.7 and for magnetite ore 4.5.
6 Flow width (m)
Bulk density
5 4 3 2 1 0 12
14
16
18
20
Draw height (m)
Moisture content This parameter is difficult to quantify in field test, because it varies from one spot to another, but the moisture conditions should be commented, for example if it is dry or wet and if it is local or evenly distributed.
FIG 4 - Relation between the draw height and flow width from the Grängesberg full-scale tests. Diagram based on primary data from Janelid (1973a).
Hebei Copper Mine Co tests in magnetite ore DETERMINATION OF DRAWBODY SHAPE The drawbody shape can be found by model test or use of formulas, but the best way is to determine it by full-scale tests. Of special interest is the maximum flow width because it determines the distance between the cross-cuts in sublevel caving or the distance between drawpoints in block caving.
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Later on, in 1976 - 1977, full-scale tests were done by the Hebei Copper Mine Co in China in magnetite. The draw height was 55 m, ~3 times higher than that used in the Grängesberg tests. The results were reported by Chen and Boshkow (1981). However, there were not enough facts given for the Hebei tests like kind of marker, marker density, fragmentation of rock, etc
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and therefore Gustafsson (1998) contacted the research leader of the Hebei field test, Professor Su Hongzhi. These tests are of fundamental importance due to the high draw height and therefore the complementary information is given in Appendix 1. The flow width versus draw height for the Hebei magnetite mine tests is shown in Figure 5.
16 14 Flow width (m)
12 10 8 6 4
of about 4500 kg/m3. The sublevel height was 30 m, and the distance between the cross-cuts 25 m. The markers were made of used electric cables, earlier used for electric LHD’s. The length of the cable or marker was 1 m, diameter 60 mm and weight 15 kg/marker. This type of marker is therefore the best of all being used so far in the field. It is large and heavy. A total of 24 experimental rounds were blasted. About 40 markers were placed in each marker fan. The reconstructed flow width for a specific sublevel caving round varied from 12.4 - 21.4 m at 11 m drift width, 3 m burden and 115 mm production blast holes. Nothing can be said about the draw height, because no markers were placed in the top of the draw because of limited resources to do the test. It is anticipated that the draw height is twice the sublevel height, 2 × 30 = 60 m, about the same as in the Hebei test. The elongation ratio of the drawbody varied from 60/21.4 = 2.6 to 60/12.4 = 4.8. Backbreak from one burden to another was a common cause to variation of the waste rock dilution.
2 0 0
5
10
15
20
25
30
35
40
45
50
55
60
Model scale tests to determine the shape of the drawbody
Draw height (m)
Ellipsoid theory FIG 5 - The influence of draw height on flow width in the draw of copper-magnetite at Hebei Mine in China. Primary data taken from Chen and Boshkow (1981).
From Figure 5, we can learn that the linear behaviour between draw height and flow width ceases at about 27 m draw height. From there and on mass flow will start. Mass flow is defined as a flow where all particles move in the same direction and with the same velocity. From Figure 5, we can also see that the curve is getting close to a horizontal asymptote at about 16 m flow width. If the results from the Grängesberg and the Hebei tests are combined the joined curved shows the trend of the development of the drawbody shape in blasted rock, see Figure 6.
Kvapil (1965) published one of the fundamental papers regarding gravity flow of granular materials in hopper and bins. This paper gives a good description of flow mainly for fine materials. The area for the drawbody to develop in was, however, restricted due to the vertical walls of the bin. There are illustrative figures of gravity flow of fine material in the paper. Kvapil approximated the vertical drawbody shape with an ellipsoid. In a horizontal cut the drawbody shape was of course circular. Janelid and Kvapil (1966) at the Royal Institute of Technology in Stockholm presented one of the classical works about sublevel caving and especially its design. At that time, the drawbody was still regarded as an ellipsoid, but in later research it has now been found that an upside down turned water drop is a better description of the drawbody shape.
Drop hypothesis (improvement of the ellipsoid theory) 16
Flow width (m)
14 12 10 8 6 4 2 0 0
10
20
30
40
50
60
Draw height (m)
FIG 6 - Combination of full-scale data from gravity flow tests of coarse blasted rock from the Grängesberg Mine in Sweden and from the Hebei Mine in China. Data fits very well to each other. Primary data was taken from Janelid (1973a) and Chen and Boshkow (1981).
LKAB Kiruna tests in magnetite ore Recently full-scale tests were done in large-scale sublevel caving, at LKAB in Kiruna. The result was reported by Gustafsson (1998). The Kiruna orebody consists of magnetite with a density
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Fröström (1970) at the Royal Institute of Technology in Stockholm did physical model tests at a scale of 1:50 and with a simulated LHD-extraction from a drift. He simulated draw heights corresponding to 38 m in full-scale. The tests were done for size distributions with three different maximum sizes of particles (2, 8 och 20 mm) representing three different scales, and three different distributions for each maximum size, and therefore, in total the gravity flow for nine different size distributions were examined. The distributions were made artificially by mixing different fractions, and the size distributions created were all linear when presented in double logarithmic diagram. The reason to this is that most blasted rock masses follow a near linear distribution in double logarithmic diagrams. Fröström found, that the maximum width of the drawbody always occurred above the upper half of the drawbody, and therefore the drawbody is not symmetric around is half height. The shape is more like a water drop turned upside down. The ellipsoid theory could therefore only be used as a rough approximation of the real shape. Nobody has later on been able to argue against Fröström’s findings, and therefore this is the valid theory today. Owing to the examination of drawbody shape for as much as nine different fragment size distributions, the validity of the result is large, and we can talk about a paradigm shift in the field of gravity flow. Still at the time these findings were done, we did not have any mathematical explanation for the ellipsoid nor the drop shape.
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Mathematical derivation and verification of the drop hypothesis The first mathematical derivation of the drawbody shape was done by Bergmark-Roos at LKAB Malmberget and reported by Hedén (1976) at LKAB in Kiruna. The basics of the development of this theory are the forces acting on a fragment, and its trajectory to the outlet from its original position. In the derivation of the formula (3) it was assumed that the fragment moves on a straight line, see Figure 7, but in the reality the travel path is a little curved due to the side component of the gravity force acting on the fragment. From basic laws of mechanics regarding forces on a single fragment above an opening, and the well-known distance formula (2) for a fragment in a gravity field, formula (3) was derived. The distance formula reads as follows: S =
at 2 2
(2)
where S is moved distance for an object (fragment), a is the acceleration (on earth a is equal to 9.81 m/s2), t is the time for the fragment to travel the distance S to the outlet. The final formula developed Bergmark-Roos explains the drop shape: S (α ) = Ho
sin α − sin α g 1 − sin α g
(3)
where S is the distance travelled for a fragment from the surface of the drawbody to the outlet. Ho is the height of draw, α is angle between the horizontal and the line defining the flow path for each single fragment. The limit border angle α g is the angle between the tangent to the drawbody at the outlet and a horizontal line. By these definitions α is therefore always >α g. By varying α g the elongation ratio of the drawbody can be changed. The elongation ratio is here defined as the ratio between the half axis of the height of the drawbody divided by half axis of the width of the drawbody at half the height of the drawbody. The material properties are characterised in formula (3) only by the limit border angle. When the shape of the drawbody is known a formula can be developed based on formula (3) to calculate the elongation ratio. It is therefore here not possible to calculate the drawbody elongation ratio from some specific material properties like shear strength and cohesion. In the field it is almost impossible to measure the limit border angle, and therefore formula (3) is limited in its use to calculate drawbodies for full-scale blasts. The Bergmark-Roos theory, formula (3), only fits well up to certain draw heights, compared to some model tests done. One explanation may be that the formula does not take into account the development of a flow tube. This is demonstrated in the model tests by the drawbody getting more and more elongated, when the draw height increases. McCormic (1968) explains that the flow pipe depends on the size of natural arches, which normally is formed within the gravity flow area. The larger the arch the greater diameter of the flow pipe. McCormic means that pipe flow only can exist if the mobility of rock is low. The mobility of a material depends on many properties like swelling (porosity), shape and surface friction of the fragments, etc and it is therefore difficult to define the mobility by one single value. If the mobility of the material is large, no pipe flow will be formed. The normal case in gravity flow of coarse rock is that a pipe flow will form.
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F
= Gravity force on a particle F = ma
Fo
= Force on a fragment caused by gravity component in the direction of the discharge opening Fo = F sin α = ma sin α
Fn
= Side force acting on a fragment. This force changes the linear trajectory of the fragment to a little bent downwards. Fn = F cos α = ma cos α
Fm
= Retarding force
FIG 7 - Force balance and trajectory for a fragment from its original position to the outlet, and definition of symbols in formulas (2) and (3). This is the basic setup for derivation of the Bergmark-Roos formula.
Other theories for simulation of gravity flow Coulomb plasticity The Coulomb plasticity theory has been used to explain movements in ore passes by Pariseau (1966). This theory is more valid for fine grained materials.
Void (diffusion) theory The void diffusion approach was first developed by Jolley (1968). The fragmented material is simulated by cubes where each cube is given a certain probability for movement in the gravity field, (see Figure 8). After one cube has been removed the surrounding cubes will have a certain probability to fall out. A void is created which will be filled by a new cube, and the void will progressively move upwards. This method of simulating a drawbody, was given the name void diffusion theory by Jolley. It is however not a good term because the term diffusion stands for the physical effect when gas molecules from one gas is penetrating into a
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FIG 8 - The void (diffusion) theory developed by Jolley (1968). a) Principal stress trajectories during steady drawdown. b) Probability of block filling void layer N and N+1.
FIG 9 - Drawbodies at four different extraction rates, a) six blocks, b) ten blocks, c) 16 blocks and d) 33 blocks removed. After Gustafssson (1998).
neighbouring gas by Brownian movement. This is a not a mixing process of two gases. It would be preferred to call it just the ‘void theory’ not to confuse it with diffusion. Disadvantages with this method are that the probability of fall out cannot be related to material properties. The void theory was modified by Gustafsson (1998) to better fit to real gravity flow observed in the sublevel caving mines. With this new theory, asymmetric drawbodies can be simulated and especially small thin extensions from a larger drawbody, (see Figure 9).
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The author named these non-symmetric drawbodies ‘palm and finger drawbodies’. It would be preferred to call it asymmetric drawbodies. The explanation to this occurrence is that some larger pieces prevented the extension of the drawbody in its expected direction, and instead smaller extensions developed due to higher mobility of that material. Rustans comments to these findings is that it is more of academic value than real mine planning value, because we can never start to plan a mine according to asymmetric drawbodies as long as we use symmetric blast plans.
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CALCULATIONS OF THE MAXIMUM AND CRITICAL FLOW TUBE WIDTH IN BINS For bins, formulas have been developed to calculate the critical and maximum flow width. This width should bear a certain relationship to the dimension of the opening. The larger the opening the larger the width should be. Also the fragment size distribution should influence the maximum width. In 1968 McCormick discussed how wide does a draw point draw. His conclusions are based on observations of ore flow from more than 1000 draw points in the field, laboratory tests with sand or coarse rock 25 - 75 mm. According to McCormic the result is the same with sand as when using coarse rock. This is astonishing because coarser material should create a wider flow tube, see Janelid and Kvapil (1966). McCormick (1968) also varied the width of the discharge opening in steps 25, 50 and 150 mm, but the width of the flow tube was about the same. This is also against the general rule, which says, that the flow pipe width should increase with the discharge opening width, see Janelid and Kvapil (1966). The critical flow tube width is reached when the flow starts to act a mass flow. This width will then be widened by erosion until the maximum flow tube width is reached. Two formulas to calculate this width in bins will be presented.
Formula for calculation of maximum flow tube width in bins Normally formulas developed for bins cannot directly be used in sublevel or block caving. However, we can get ideas what parameters are of interest and therefore these formulas should also be of some interest in sublevel and block caving. One formula for bins is given in the book written by Reisner and von Eisenhart Rothe (1971). Since the diameter of the outlet affects the flow pattern of a material in a bin, test were run to determine the maximum diameter of the boundary between flowing and non-flowing material for flat-bottom bins in relation to the diameter of the circular discharge opening. Provided the material is free flowing, the width of the flow tube could be calculated by the following formulas.
Dmax = dout
Dmax = dout
dout H − f1 2 + 2 tan θ 1 + tan θ
when
H > f1
dout 2
(4)
when
H < f1
dout 2
(5)
where Dmax is the maximum diameter of the flow tube, dout is the diameter of the circular outlet, θ is the angle between the vertical and the slope of the flow in the bottom part of the bin, H is the height of material in the bin, f1 a derived dimensionless flat-bottom bin factor. Formula (4) is, however, not very useful, because there are no material properties in the formula, and we don’t know the value of θ and f1. θ, however, relates to the material properties. f1 will be difficult to determine in the field.
Formula for calculation of critical flow tube width in bins A formula to calculate the critical flow tube width is of outmost importance for the design of sublevel or block caving. When the critical flow width is reached, the diameter can increase a little more if the draw is continued. This is due to the erosion of the side walls from the flowing material until the maximum flow tube width is reached. The erosion is according to Stazhevsky (1992) mainly dependent on coarseness of the fragments, the
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stresses acting in the flow tube and the initial packing density of the granular material (ie on its dilatancy properties) in the dilatency box. Stazhevsky (1992) reports that the contribution of density and shape of particles to erosion is small, and that the Mining Institute of Novisibirsk have failed to relate the mechanism of erosion either with the active stresses or with the physical and mechanical properties of granular materials. The goal for the co-operate research between Luleå University of Technology and the Mining Institute of Novosibirsk in Novosibirsk Russia was to find a formula to calculate the critical flow tube width before the erosion of the flow tube starts. Part of this work was published by Månsson (1995). The formula, based on laboratory tests, cannot directly be used for full-scale design, but we can learn what parameters are important for the flow tube width in the model, and assume that these parameters also will be important in full-scale. In the experiments, the discharge opening was rectangular and the discharge was made by a piston, which means a mass flow was created over the whole discharge area. This is not similar to the discharge in a mine where it is more like a point discharge if the cross-cut has a half circular shaped roof. The problem in finding a proper formula made it, however, necessary to work at laboratory scale. Because of the fact that the density is an important parameter for the flow tube width, three different densities were tested at Luleå, sand 0.1 - 1 mm with even shape (ten tests), magnetite 3 - 5 mm with angular shape (three tests) and lead shot (used for hunting) 3 - 4 mm with near spherical shape (one test). In Novosibirsk, marble 3 - 5 mm angular shaped, iron shot 3 - 5 mm with near spherical shape and polystyrene balls 3 - 5 mm with near spherical shots were tested in the same type of model test equipment. The following empirical formula was developed from the model test, and it can be used for the calculation of the critical diameter of the flow tube Dcrit. Dcrit =
2 anσ zoµ m ξ 2 δσ z aγ 1 − a − 2 nσ zoµ w δz
(6)
where a is the depth of the discharge opening which was equal to the depth of the box used in the experiments, n is the stress concentration factor which is defined as the ratio between vertical stress at the rim of the outlet after draw has started, and the vertical stress at the bottom of the bin before start of draw, σzo static vertical stress in the flow tube before start of the draw, µm friction factor between material at rest in the box the flowing material, ξ 2 lateral thrust coefficient in the box defined as: ξ 2 = tan( 45°− (ϕ µ + υmax ))
(7)
where ϕµ is the effective angle of friction between particles and υmax is the maximum dilatancy angle. The sum of these two angles defines the angle of internal friction. γ 1 is the bulk density of the drawn material, σz vertical stress in the bin, z is the vertical depth below the top surface of the material in the box and finally µw is the friction factor between the walls in the box and the material in the bin. The accuracy of formula (6), when compared to the performed model tests varies between - 31 per cent and + 24 per cent. If we assume that this formula should be transformed and used for full-scale tests, the accuracy has, however, to be improved. The inaccuracy is too large to be acceptable for full-scale design. Comments made by Agne Rustan to formula (6):
• If the nominator and denominator are divided by ‘a’ the last term in the denominator can be seen as a correction term if the bin is not deep enough. If ‘a’ is very large the last term will be negligible.
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• From all the parameters given in formula (6) there are two that directly informs about the material in the experimental box, the bulk density and the angle of internal friction. Because we cannot change the bulk density so much in real sublevel caving, we cannot affect the maximum draw width very much. Different fragmentation of the ore by blasting in sublevel caving or by blocking in block caving will, however, give different angles of internal friction, and this will affect the draw width.
• If we use a ‘silo profile’ in sublevel caving (a silo profile is
parameters for mining of a whole sublevel or block. The computer program must, however, always be calibrated to the field by full-scale tests.
• More full-scale tests should be done in close collaboration with mines. Reasonable and achievable goals have to be set up because too much research work is today focused on establishing physical modelling techniques, which often are not useful for the mines. Physical model testing will today create low interest from the mines because they won’t be interested in such kind of research.
defined as a blast fan where the two side holes in the fan are almost vertical) the third factor in the denominator will have some importance, because the wall friction will occur both at the blast front and at one or two side walls defined by the breakage from the corner holes in the blast.
• Undertake more full-scale tests in the mines under
• The stress concentration factor, in formula (6), is difficult to
• The ultimate goal will be, to determine better design rules for
measure in full-scale, and therefore also this formula is not very useful for the design of sublevel caving. From model tests done at Luleå University it was, however, found that the stress concentration factor varied as much as from 1.1 to 1.5 dependent on the individual model test, see Månsson (1995). This factor therefore affects the Dcrit by as much as 40 per cent. The development of a flow tube or the chimney effect depends on the packing of the material, a high packing means development of flow tube, and a loose packing means no development of flow tube according to the model test performed at Luleå. Månsson (1995) states, that in the loose state of packing, the wall of the flow tube is usually not stable enough to shape an explicit tube. According to all known literature about gravity flow of blasted or blocked rock, a flow tube normally develops above the draw point. Indirectly this means that the blasted or blocked rock normally is well packed. It is, however, urgent to find some better design tools. To be able to do this an international group has to be formed to start the discussion of how we can find good design tools. With the conclusions from this international group it should be easier for a specific nation to apply for research money, especially if a plan is worked out how each country best should contribute in international research. What we need is a formula for full-scale use to determine the optimum draw point spacing dependent on the material properties of the blasted or caved ore. A suggestion to properties that may be of interest for the classification of rock regarding gravity flow was already given in the section entitled ‘Material Properties and Quantification of Physical and Mechanical Properties of Blasted or Caved Rock’. Each science has to define what properties are of most interest for their purpose of interest. Rock mechanics needs some properties, blasting some other properties and gravity flow another set of properties.
CONCLUSIONS The following conclusions can be made:
• An international research leader group has to be formed to analyse the need for more research, develop ideas how the research should be done, standardise test procedures and apply for funding.
• The Bergmark Roos theory has to be improved so it takes into account the flow pipe effect.
• A 3D computer program should be developed where different shapes of the drawbodies can be simulated for calculation of ore losses and waste rock dilution at orebodies with defined properties of the ore and geometry of hanging and footwall. The computer program can then be used to optimise different
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standardised test procedures and gather the information in a database. Mines with different fragmentation, shape of fragments and surface friction of fragments should be interested to take part in such research. sublevel and block caving and document the rules in a handbook.
REFERENCES Bucky, P, Steward, J W and Boshkov, S, 1943. What is proper draw point spacing for block caving? Eng Min J, 144(6):70-75. Chen, J Y and Boshkow, S, 1981. Recent development and application of bulk mining methods in the People’s Republic of China, in Proceedings International Conference on Caving and Sublevel Stoping, pp 393-423, (SME-AIME: Denver). Forsman, B and Pan, Chang Liang, 1983. Shear properties of fragmented rock masses- model tests and theories. Research report TULEA 1983:40, Division of Mining, Luleå University, Sweden, STU project No. 82-5490, 52 pp. Fröström, J, 1970. Examination of equivalent model materials for development and design of sublevel caving. Master of Science Thesis E 80, Royal Institute of Technology, Stockholm, 26 p (in Swedish). Fumagelli, E, 1969. Test on cohesionless materials for rockfill dams, (ASCE) Journal of the soil mechanics and foundations division. 95 (SM1):313-330. Gustafsson, P, 1998. Waste rock content variations during gravity flow in sublevel caving. Analysis of full-scale experiments and numerical simulations. Luleå University of Technology, Department of Civil and Mining Engineering, Division of Rock Engineering. Doctorial Thesis No 1998:10, 194 pp. Hedén, H, 1976. Sublevel caving seminar in Kiruna. LKAB, Kiruna (in Swedish). Janelid, I and Kvapil, R, 1966. Sublevel caving, Int J Rock Mech Min Sci, 3(2):129-153. Janelid, I, 1973a. Mining with caving, The Swedish Rock Engineering Foundation (BeFo), Stockholm (in Swedish). Janelid, I, 1973b. Study of the gravity flow at sublevel caving in full- and model scale. Swedish Mining Research, B-series, No 180, Stockholm (in Swedish). Jolley, D, 1968. Computer simulations of the movement of ore and waste in an underground mine, CIM Bulletin, 61(675):854-859. Julin, D E and Tobie, R L, 1973. Block Caving, SME Mining Engineering Handbook, Chapter 12.14, pp 12-162 to 12-167. Kvapil, R, 1965. Gravity flow of granular materials in hoppers and bins, Int J Rock Mech Min Sci ,Vol 2, part 1: pp 35-41, part 2: pp 277-304. Marachi, N D, 1972. Evaluation of properties of rockfill materials. (ASCE) Journal of the Soil Mechanics and Foundations Division, 98, (SM1):95-114. Maripuu, R, 1968. Examination of size distribution and fragment shape for representative material from sublevel caving at LKAB in Kiruna. Master of Science Thesis, Royal Institute of Technology, Stockholm, 30 p (in Swedish). McCormick, R M, 1968. How wide does a drawpoint draw? Eng Min J, June, 169(6):106-116. Månsson, A, 1995. Development of body in motion under controlled gravity flow of bulk solids. Licentiate Thesis, Luleå University of Technology, 1995:19 L, Division of Mining Engineering, 101 p.
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Pariseau, W G and Pfleider, E P, 1968. Solid plasticity and the movement of materials in ore Passes, Transactions of the AIME, March, pp 42-56. Reisner, W and Von Eisenhart Rothe, M, 1971. Bins and bunkers for handling bulk materials, 280 p, (Trans Tech Publications: Clausthal Germany). Rustan, A, 1970. The theoretical basis of using bulk density as a measure of ore content and model blast tests to determine, swelling, gravity flow, in bench- and sublevel caving models. Technical Licentiate Thesis, Royal Institute of Technology, Stockholm, 237 p (in Swedish). Rustan, A, 1990. The importance of using joints to achieve scaled fragmentation in magnetite concrete used for sublevel caving blast models, Engineering Fracture Mechanics, 35(1/2/3):425-438. Rustan, A, 1998. Rock Blasting Terms and Symbols – A dictionary of symbols and terminology in rock blasting and related areas like drilling, mining and rock mechanics, 193 p, (A A Balkema: Rotterdam). Stazhevsky, S B, 1992. Gravity flow of disrupted rock - Report for the years 1991 - 1992. Summary of major research results achieved by the Inst of Min Novosibirsk in the Science of Gravitational Flow of Free-Flowing Bulk Materials, OSS. Wahlström, E E, 1955. Petrographic Mineralogy. New York. Wood, P A, 1986. Fundamentals of bulk solids flow. IEA (International Energy Agency) Coal Research, London, Report number ICTIS/TR31.
3520 markers were placed in a total of 177 marker holes. These markers consisted either of wood-filled plastic tubes, 45 55 mm in diameter, or ventilation pipes 25 or 50 mm in diameter. Both marker types were cut to 250 mm in length. Markers were placed in drill holes drilled upwards from the drift where the mucking took place and in the drift immediately above it. The marker drill holes were drilled with an YG-80 drilling machine. The marker planes consisted of 9 - 12 marker drill holes at the lower sublevel and 18 - 19 marker drill holes at the higher sublevel. The markers were held in place in the marker drill holes with clay and a drill hole plug below each marker. The markers were separated with 0.5 m wood lengths. Each marker drill hole contained four, eight or 12 markers. In Figure A2 seven marker planes are shown in a vertical view along the centre of the cross-cut.
APPENDIX 1 Additional information on the full-scale test performed at the magnetite mine in China, reported by Gustafsson (1998). The Hebei Copper Mine Co, in collaboration with Beijing Metallurgy and Mining University (now Beijing Science and Technology University) and Maanshan Institute of Mining Research, carried out an experiment similar to that of Grängesberg in 1976-1977. The experiment was carried out in orebody No 2 Lóngtan Iron Ore Mine, Shouwangfen copper mining district, Hebei Province. The ore was a low-grade magnetite with a density of 3800 kg/m3. At the experimental site the ore dip was 80° - 90°, and the ore width varied between 30 50 m. The drifts were approximately 3.3 m in width and 3.2 m in height. The blasted fan height was 50 m and the pillar width 40 m. The blast front inclination was vertical. The burden was about 1 m and 8 blast rings with a total burden of 8.4 m were blasted simultaneously. The amount of blasted ore was 32 000 t. A horizontal view of the experimental site is shown in Figure A1.
FIG A2 - Vertical section perpendicular to blast front showing marker hole planes in the Hebei magnetite mine. Numbers indicate height in m above the floor of the cross-cut.
FIG A1 - Horizontal view of the experimental sites in the Hebei magnetite mine. Cross-cuts showing eight burdens all blasted in one big mass blast.
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The observations were done wagon by wagon. Broken rock, approximately 2 m3 was loaded into rail wagons, and an observation sheet was filled in for each wagon. The data recorded were number of wagon, identification of markers found, lateral position of marker in the muck pile, position in the experimental site, time of loading, wagon weight and fragment size. At the time of the report writing, 15 200 t had been mucked at the experimental site. Sixty per cent of the installed markers have been collected. The fragment size distribution was 51 per cent < 100 mm, 24 per cent 100 - 300 mm, 15 per cent 300 - 500 mm
and ten per cent >500 mm. The blasting was less than perfect, with a large open volume in the broken rock formed in the beginning. The muck proceeded continuously, and the broken rock was dry. The waste rock percentage was not reported. (With such big burden it should, however, not be any waste rock at all.) The markers were used to reconstruct the drawbodies. The drawbodies developed from the tests are shown in Figures A3, A4 and A5. According to the supervisor for the experiment, the raw data from the Hebei field test do not exist any longer.
FIG A3 - drawbodies at the Hebei field test, vertical frontal section.
FIG A4 - Horizontal section of drawbodies in the Hebei field test at three different volumes extracted, I, II and III. Observe the near elliptical shape indicating that a flat roof shape has been used in the cross-cuts.
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FIG A5 - drawbodies at the Hebei field test. Vertical section along the centre of the cross-cut. Numbers indicates the height in m above the cross-cut floor.
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Drawpoint Design in Caving and Stoping Mines F O Otuonye1 ABSTRACT The design of drawpoints in caving and stoping mines has a significant impact on the productivity of the operation. Drawpoints must be designed to provide uniform drawdown and maximum recovery of the broken ore within the limits imposed by the dimensions of the stope, the stability of the pillars, the method of draw, material flow characteristics, etc. The design of drawpoints historically relied on the mining engineer’s experience and intuition. However, in many operations, empirical analyses have also produced important information regarding design geometry. Factors that affect the sizing and spacing of drawpoints are discussed and guidelines are provided for their design in order to obtain maximum recovery of broken ore and reduce hang-ups.
INTRODUCTION The success of large-scale caving and stoping operations depends largely on efficient and cost-effective handling of the fragmented material. Draw systems function as control stations that provide the means of recovering broken ore as the stope advances or when stoping is completed. Ore is typically drawn from the stope sill level through drawpoints, chutes or slusher pockets spaced along the sill. The design of draw systems requires careful consideration of a number of factors such as material flow characteristics, geometry and layout of drawpoints, extraction methods, stability of pillars, financial considerations, etc. While it is possible to estimate the effect of each of these factors in isolation, it is generally difficult to assess the precise effect of all these factors on draw behaviour, primarily because the interaction of these factors is difficult to formulate and evaluate. However, with extensive investigations using model studies, field tests and mathematical modelling, it is possible to predict draw behaviour and gain a better understanding of material flow characteristics. The design of an optimum configuration of drawpoints affects the cost of development and generally leads to more efficient allocation of personnel and machines and therefore increased productivity. Controlled draw from openings would result in less dilution of ore and better utilisation of equipment. The efficiency of draw, the maximum ore recovery and the optimisation of production are directly related to draw management.
very little control over the degree of ore fragmentation. Therefore, estimates of the degree of cavability of the ore and assessments of the size distribution are predicted from fracture spacing and orientation (White, 1979). In stoping operations, where drilling and blasting are employed to fragment the ore, a measure of control can be exercised. The degree of rock fragmentation can be estimated using any of the available models or modifications thereof (Otuonye, 1985).
Shape or angularity of ore fragments The shape of the ore fragments is as important as the size distribution in determining the mobility of the ore as it is drawn from the stopes. Two indices are useful in the quantitative description of shape. They are roundness and sphericity. While roundness or angularity measures the sharpness of the corners, sphericity on the other hand is an index of how closely the particle approaches a sphere. The ease of handling broken ore, the ability to flow readily and the resistance to break down when drawn are related to sphericity. The higher the sphericity, the less the tendency of the fragments to fracture or degrade into small particles as they are drawn. Ore fragments with a blocky shape will draw more readily than slabby ore fragments because the long axes of slabby fragments are suitable for producing hang-ups (McCormick, 1968). Particles with high angularity have the ability to interlock with each other and tend to form a wedge with themselves and the walls of the stope or drawpoint. This would contribute to the formation of an arch and thus reduce the mobility of the ore (Kvapil, 1965). The strength of the interlocking bonds depends directly on the particle strength and the type of interlock (Pietsch, 1969).
Frictional and cohesive properties of the ore
Extensive studies on the flow of bulk solids in bins, bunkers and silos have been conducted by several investigators (Jenike and Johanson, 1972; Janelid and Kvapil, 1966; Kvapil, 1965) in order to gain a better understanding of bulk material flow characteristics and to determine conditions for which a given bulk solid will flow without obstruction. Much of the under-standing of the flow of broken or fragmented ore through drawpoints, chutes and slusher pockets in caving and stoping mines is a direct result of these studies. The flowability of broken rock is affected by many factors, but the most significant are the following.
Apart from weight or inertial effects, the flow of ore through drawpoints depends on its frictional and cohesive properties. While surface forces dominate the behaviour of fine particles, gravity forces dominate the behaviour of large particles. The frictional resistance of particles is considered to consist of the mechanical resistance to sliding and the overriding of one particle past another. Cohesion, on the other hand, is considered the result of physico-chemical bond forces between particles. These forces are inversely related to the void ratio, confining stresses or the degree of compaction, shape and size distribution of the broken ore. Sticky, fine ores would tend to adhere to each other due to frictional and cohesive properties and would result in hang-ups if the fines have sufficient strength. Coarse ores would resist motion due to interparticle friction. The chance of an arch or a hang-up occurring depends on the span of the opening. If the opening is large enough, gravity forces will exceed the frictional and cohesive strength of the material and prevent a hang-up.
Size distribution of rock fragments
Degree of consolidation
The degree of rock fragmentation determines the success of any mining operation. It affects the production rates, the safety of the operation and the operating costs. In induced caving, there is
Some bulk solids have sufficient mobility under conditions of continuous flow, but develop strength and would not flow readily after storage. The degree of flowability of such solids would depend on the fragment size distribution, the pressure under which the material has been consolidated, the length of time it has been under pressure, moisture and temperature conditions.
MATERIAL FLOW CHARACTERISTICS
1.
Mining Engineering Department, Michigan Technological University, 1400 Townsend Drive, Houghton MI 49931-1295, USA.
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Generally, the factors which may constitute the main resistance to flow are the density, the interlocking of the particles and the resistance to shear between the particles. If broken ore has been in a stope for a long period of time, it may be consolidated due to pressure and develop higher shear strength, therefore the ore may not be as mobile as would be expected. Consolidation can be more of a problem in a shrinkage stoping than in sublevel stoping or block caving mines because part of the ore in shrinkage stoping operations is left in the stope to serve as a working platform for mining the ore above and to support the stope walls.
Moisture content and temperature of broken ore The flowability of the broken ore would decrease with an increase in moisture content until a minimum degree of saturation is attained when the bulk solid becomes mobile again. The minimum moisture content when flowability resumes depends on the size distribution and the characteristics of the fragmented ore. With plastic materials, such as clay, a high moisture content will reduce the effectiveness of the packing and interlocking of particles. For most broken or fragmented ores, depending on the mineral content and degree of consolidation, a significant increase in temperature would result in an increase in adhesion of particles and a reduction in mobility.
Draw rate
Method of draw The method used to draw ore from the stope directly affects the results of the operation. Ore can be drawn from a stope with load-haul-dump (LHD) machines, slushers or by gravity. The advantages and disadvantages of each method of draw for a caving operation are discussed by Ward (1981), and would also apply for a stoping operation. Generally, an economic analysis of each method of draw for a particular operation is useful in determining the most suitable and efficient method. It is apparent that for most stoping operations, the LHD method of draw would provide the production and flexibility required to yield a low-cost operation.
DESIGN OF DRAWPOINTS Theoretical and experimental work by Janelid et al (1966); Kvapil (1965); McCormick (1968); etc have shown that the volume of material that flows through a drawpoint in caving operations is equivalent to that which would be contained by an ellipsoid above the drawpoint before drawing occurred as shown in Figure 1. The draw shapes and the theory that describe them have been used by many researchers in actual caving operations to explain the behaviour of ore flow. It is postulated that draw shapes in stoping operations are similar to those of caving operations, thus theoretical treatments of ore flow in caving operations can be extended to stoping mines.
The intensity of ore extraction from the drawpoints plays an important role in material flow. Draw should proceed at such a rate as to increase the amount of ore recovered with a corresponding reduction in the amount of waste drawn. The minimum draw rate is determined by the number of active drawpoints and the horizontal area that has to be recovered at the maximum height of draw. If the ore is drawn from drawpoints uniformly or if adjacent drawpoints are worked simultaneously within the limits imposed by grade control and safety, it is less likely that arching or bridging of particles would occur than if the stopes were drawn empty in sequence starting from one stope down the length of the stope. According to Heslop and Laubscher (1981), it is not essential that drawpoints be worked simultaneously as long as sufficient drawpoints are worked frequently enough to preserve the interaction of stress fields above the drawpoints. In stoping with backfill where waste rock is introduced into the void area or stope in order to ensure the stability of the walls, the problem of waste dilution is exacerbated especially if draw is accomplished at the same rate as the waste is introduced into the stope. In this case, design concepts can be derived from draw theory.
Stope height The eventual stope height is generally governed by safety and economy, and for most practical cases would not exceed 120 metres. A high ore column or stope is economical to develop, but control over ore dilution and size distribution of the fragmented ore may not be readily achieved. If the thickness and dip of the orebody are fixed, a high stope may result in ore loss in the footwall and poor extraction. The mobility of the ore may also be reduced if the ore has been dormant in the stope for a long period of time. In caving operations, it may be advantageous to have a high stope, within the limits imposed by safety and economy, because more fines will be produced by attrition as the large ore fragments travel through large distances in the stope. However, the crushability of the rock, the distribution of fractures, the horizontal extent, thickness and dip of the orebody, and the percentages of fines produced are some of the factors that would determine the eventual stope height.
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FIG 1 - Draw shape based on ellipsoid theory.
It is assumed that the ellipsoid of draw shown in Figure 1 is cut-off at the base by a plane parallel to the base and corresponding to the width of the drawhole. The volume of the
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ellipsoid of draw, Vd, can be calculated, with the parameters shown in Figure 1, by the formula: Vd =
πr 2 hd2 h (1 − d ) a 3a
(1)
where r = semi-minor axis of ellipsoid of draw
4 which is the 3πrl2 al volume of the limit ellipsoid. When u = 0, w = 2rl, h = al, then 2 which is the semi-volume of the limit ellipsoid. Vl = 3πrl2 al When u = 0, hl = h, then Equation 4 can be written as: When u = 0, w = 0, hl = 2al, then Vl =
1 π 2 w2 w2 r1 h 1 + 2 + 1 − 2 2 4rl 3 4rl
Vl =
hd = height of ellipsoid of draw a = semi-major axis of ellipsoid of draw b = distance from centroid of ellipsoid of draw to the height of the drawpoint (a>b) Equation 1 can be expressed in terms of the parameters r, hd, and the opening width, w, of the drawhole by the formula: 1 π 2 w2 w2 2 Vd = r hd 1 + + 1 − 4 r2 4 r2 3
(2)
Equation 5 is the volume of the limit ellipsoid cut-off by a plane parallel to the base of the drawpoint. The volume is similar in form to that of the ellipsoid of draw given in Equation 2. The volume and height of the loosening ellipsoid are 15 and 2.5 times the volume and height of the ellipsoid of draw respectively (Janelid and Kvapil, 1966). The shape of the ellipsoid of draw is determined by its eccentricity and is given by the equation: ε=
4 which corresponds When w = 0 and hd = 2a, then Vd = 3πr 2 a to the volume of an ellipsoid. 2 When w = 2r and hd = a, then V = which corresponds to 2 πr a 3 the volume of a semi-ellipsoid. As the ore is drawn from the stope through the drawpoint, it is also loosened around the ellipsoid of draw. The loosening outline is similar in shape to the ellipsoid of draw and is called the limit ellipsoid. The contours of the loosening ellipsoid form a boundary between the zone of motion (inside the loosening ellipsoid) and the stationary zone (outside the loosening ellipsoid). The limit ellipsoid therefore defines the limit of the flow zone and the ellipsoid of draw defines the volume of the broken ore recovered from within the flow zone. The volumes of the loosening ellipsoid and ellipsoid of draw are increased as draw progresses, but the ore or waste outside the outline of the loosening ellipsoid remains stable. In stoping and in caving operations, the loosening ellipsoid may have the contour of an ellipsoid that is cut off at both ends by planes parallel to the opening width of the draw hole. The general equation for the volume of the loosening ellipsoid, Vl, can be written as:
(5)
1
(6)
a a2 − r 2
The particle size of the material is inversely proportional to the eccentricity of the ellipsoid draw, ie smaller particles will produce a larger eccentricity of the draw ellipsoid than larger particles. The wider the width of draw, the smaller the eccentricity and the less flowable the material. Ore fragments with higher mobility will tend to have an ellipsoid of draw of larger eccentricity. The composition of the ore, its mechanical properties, moisture content, etc will affect the shape of the ellipsoid of draw and its eccentricity. For the same material, the eccentricity of ellipsoid of draw is influenced by the diameter of the drawpoint raise. The larger the diameter of the opening, the larger the eccentricity and the more flowable the ore. An increase in the height of the ellipsoid of draw will result in an increase in eccentricity until the height reaches approximately 30 times the diameter of the drawpoint raise when the eccentricity remains constant. The rate of draw also affects the eccentricity, ie a higher rate of draw will produce a smaller draw width and therefore higher eccentricity.
SIZE OF DRAWPOINT
k3 + d 3 Vl = πrl2 ( k + d ) − al2
(3)
where rl = length of semi-minor axis of the loosening ellipsoid k = distance from the centroid of loosening ellipsoid to any arbitrary level above the centroid d = distance from the centroid of loosening ellipsoid to the height of the drawpoint al = length of semi-major axis of loosening ellipsoid Expressing Equation 3 in terms of rl, u, w, and hl, the following is obtained: 1 1 1 1 u2 2 w2 2 2 2 2 Vl = π r1 hl 1 − 2 − 2 ( u + w ) − 1 − 2 1 − 2 (4) 3 4 rl 4 rl 4 rl
In designing drawpoints, it is important to determine the size and spacing of drawpoints to ensure optimum flow, maximum ore recovery and minimum ore loss. The size of the drawpoint is determined by estimating the size distribution and the characteristics of the broken ore. Model experiments coupled with field experience indicate that the drawpoint size is three to six times the largest fragment size (Kvapil, 1965; Aytaman, 1960; Borquez, 1981). However, consideration must be given to the cohesion and frictional properties of the broken ore. Consequently, Coates (1981) presented an equation for determining the minimum width of opening to draw cohesive ore based on its density and degree of cohesion. The equation can be written as: σv =
A / P ( γ − CP / A ) (1 − e − Kz tan φ P / A ) K tan φ
(7)
where
where
σv = vertical stress acting on the top surface of an assumed trap door located on the drawpoint
u = width of loosening ellipsoid at any height, hl
A = area of the drawpoint opening
w = width of drawhole
P = perimeter of the drawpoint opening
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F O OTUONYE
γ = unit weight of broken ore
b.
C = Shear strength of broken ore K = ratio of horizontal to vertical stress σ − 2 c tan φ K= v σ v (1 + 2 tan 2 φ )
k3 + d 3 10 Vs = S d2 h − πrl2 ( K + d − al2 When ore is drawn from the stopes with inclined footwalls, the ore loss depends on the dip angle of the footwall. The value of the ore lost in the stopes can be determined. The drawpoint spacing desirable is the one which will provide a minimum cost for the value of ore left per square foot of stope (Bucky, 1943). The amount of development work depends on the drawpoint spacing. The adopted drawpoint spacing should ensure safety and economy of operation.
φ = angle of friction of broken ore z = vertical height of broken ore above the draw point When there is no obstruction to flow σv = 0, then Equation 7 reduces to: P/A=γ/C
(8)
For a circular opening diameter, D, Equation 8 becomes: D=4c/γ
(9)
c.
Based on the drawpoint spacing, plan of openings and geotechnical parameters of the ore, a stability analysis of the sill-pillar should be performed. The drawpoint spacing may be increased if it is determined that the openings are unstable.
d.
The adopted number and spacing of drawpoints should maximise net revenues from the stope, recover the largest possible amount of ore and ensure the stability of the pillars. A graphical approach could be used at this stage in determining the optimum number and spacing of drawpoints to ensure safety and economy of operation.
SPACING OF DRAWPOINTS The spacing of drawpoints depends on a careful analysis of the interaction of several factors such as material flow characteristics, method of draw, stability of pillars, financial considerations, etc. In planning the spacing of drawpoints, it would be valuable if the shape of the limit ellipsoid could be determined. The limit ellipsoid of adjacent drawpoints should touch or slightly overlap if uniform drawdown and maximum ore recovery is to be achieved within the flow zones from individual drawpoints. When drawpoint spacing is too close, too much overlap of the limit ellipsoids occurs and draw procedures become critical. Additionally, close spacing of drawpoints increase development costs, but reduces the amount of ore stranded on the sill. It can create ground control problems because the pillar size between drawpoints is reduced. When drawpoint spacing is too large, development costs are reduced, but ore losses increase. This results in static areas between the drawpoint which may transmit weight to the working areas below (Richardson, 1981). It is therefore critical to weigh the cost of the ore loss from an expanded drawpoint spacing and wider pillar sizes against the extra expense of more development, better recovery and the potential for ground control problems. Model experiments could be performed to determine the shape of the limit ellipsoid and therefore the optimum number and spacing of drawpoints which maximises net revenues within the limits imposed by recovery and stability of the pillars. Computer simulation of material flow based on empirical analysis will facilitate the consideration of a large number of alternative design layouts and operating schedules. The basic models can be developed by employing draw theory. Interactive graphics will permit the visualisation of material flow designs and reduce the time involved in physical model testing. The planning and operational problems created by the lack of precise scientific knowledge concerning the flow characteristics of broken ore can be resolved by a combination of theoretical and empirical analysis. In either case, the following procedures could be used to determine the optimum number and economic spacing of drawpoints for a given stoping operation. a.
572
The shape of the limit ellipsoid of each drawpoint should be determined by model experiments and computer simulation, and the tributary volume calculated using the formula given by Equation 3 or Equation 4. Drawpoints should be spaced so that the limit ellipsoids touch or slightly overlap in order to ensure an optimum recovery of broken ore.
The volume of ore stranded on the sill or lost between drawpoints can be obtained by subtracting the volume of the limit ellipsoid from the tributary volume of each drawpoint. If the spacing of drawpoints is Sd, and the height of the limit ellipsoid is h, then the volume of ore lost in the drawpoints is given by:
CONCLUSION The number, size and spacing of drawpoints for an existing mine depends mainly on the mining engineer’s experience, but for a new mine it is economical to employ model testing and computer simulation techniques. An approach to determine the drawpoint size and spacing has been suggested. There are many other factors not discussed in this paper that would influence the choice of an economic number and spacing of drawpoints. However, their exclusion from this paper does not diminish their importance and consideration in sizing and spacing drawpoints.
REFERENCES Aytaman, V, 1960. Causes of ‘hanging’ in ore chutes and its solution, Canadian Mining Journal, 8:77-81. Borquez, G V, 1981. Sequence of the analysis of a block caving mining method, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D Stewart), Chapter 21 (SME/AIME: New York). Bucky, P B, Stewart, J W and Boshkov, S, 1943. What is the proper draw-point spacing block caving, Engineering and Mining Journal, 144:70-75, 130. Coates, D F, 1981. Rock Mechanics Principles, Canmet Energy Mines and Resources Canada, Monograph 874. Heslop, G T and Laubscher, D H, 1981. Draw control in caving operations on South African Chrysotile Asbestos Mines, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D Stewart), Chapter 59 (SME/AIME: New York). Janelid, I and Kvapil, R, 1966. Sublevel caving, International Journal Rock Mechanics Mining Science, 3:129-153. Jenike, A W and Johanson, J R, 1972. Quantitative design of bins for reliable flow, Minerals Science Engineering, 4. Kvapil, R, 1965. Gravity flow of granular materials in hoppers and bins, International Journal Rock Mechanics Mining Science, 2:25-41, 227-304. McCormick, R J, 1968. How wide does a drawpoint draw, Engineering and Mining Journal, 169:106-118. Otuonye, F O, 1985. An assessment of size distribution of rock fragments from blasting in an underground hardrock mine, in Proceedings First Mini-Symposium on Explosives and Blasting Research, pp 69-80 (Society of Explosives Engineers: San Diego, CA).
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DRAWPOINT DESIGN IN CAVING AND STOPING MINES
Pietsch, W B, 1969. Adhesion and agglomeration of solids during storage, flow and handling: a survey, Journal of Engineering for Industry, Transactions ASME. Richardson, M P, 1981. Area of draw influence and drawpoint spacing for block caving mines, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D Stewart), Chapter 13 (SME/AIME: New York).
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Ward, M H, 1981. Technical and economical considerations of the block caving mine, in Design and Operation of Caving and Sublevel Stoping Mines (Ed: D Stewart), Chapter 11 (SME/AIME: New York). White, D H, 1979. Analysis of rock mass properties for the design of block caving mines, SME/AIME Preprint No 79-309.
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MassMin 2000 Sublevel and Longhole Stoping A Review of Sublevel Stoping
E Villaescusa
577
Mount Isa Mines — 1100 Orebody, 35 Years On
D Grant and S DeKruijff
591
George Fisher Mine — Feasibility and Construction
L B Neindorf and G S B Karunatillake
601
The Ventilation and Refrigeration Design for Australia’s Deepest and Hottest Underground Operation — the Enterprise Mine
R Brake and B Fulker
611
Bulk Low-Grade Mining at Mount Charlotte Mine
P A Mikula and M F Lee
623
The Mine Management System at Olympic Dam Mine
S Youds
637
Materials Handling at Olympic Dam Mine — Olympic Dam Aiming for ‘Gold’
P Bowman
645
The Automation of Western Mining Corporation’s Olympic Dam Underground Rail Haulage System
R Doubleday and D Mee
649
Open Stope Design at Normandy Golden Grove Operations
T M Calvert, J B Simpson and M P Sandy
653
Open Stope Mining in Canada
Y Potvin and M Hudyma
661
Open Stope Mining Strategies at Brunswick Mine
B Simser and P Andrieux
675
Evolution of Vertical Crater Retreat Mining at Mindola Mine, Zambia
E K Chanda and C Katonga
685
An Experimental Study on Large-Diameter Longhole Mining of High Stope at Anqing Copper Mine
Wang Renfa, Xia Qian and Jiang Zhiming
697
A Review of Sublevel Stoping E Villaescusa1 ABSTRACT A review of the factors controlling stope wall behaviour such as excavation geometry, rockmass strength, induced stress, ground support, blast damage and drill drive layout is undertaken. In addition, the advantages and disadvantages of single lift stope geometries linked to vertical crater retreat are also analysed. Finally, key components of multiple lift stoping such as cut-off slots, production rings, stope undercuts and drawpoints are studied in detail.
INTRODUCTION Sublevel open stoping methods are used to extract large massive or tabular, steeply-dipping competent orebodies surrounded by competent host rocks which in general have few constraints regarding the shape, size and continuity of the mineralisation. The success of the method relies on the stability of large (mainly un-reinforced) stope walls and crowns as well as the stability of any fill masses exposed. In general, open stopes are relatively large excavations in which ring drilling is the main method of rock breakage. Ore dilution consisting of low-grade, waste rock or minefill materials may occur at the stope boundaries. In addition, ore loss due to insufficient breakage can also occur within at the stope boundaries. 1.
Professor of Mining Geomechanics, Western Australian School of Mines, PMB 22 Kalgoorlie WA 6430. E-mail: [email protected]
The method offers several advantages including, low cost and efficient non-entry production operations, utilisation of highly mechanised, mobile drilling and loading production equipment, high production rates with a minimum level of personnel, furthermore, production operations are concentrated into few locations such as ring drilling, blasting and drawpoint mucking. The disadvantages include a requirement for a significant level of development infrastructure before production starts, thus incurring a high initial capital investment. However, most of the development occurs within the orebody. In addition, the stopes must be designed with regular boundaries and internal waste pockets cannot be separated within the broken ore. Similarly, delineated ore cannot be recovered beyond a designed stope boundary. Technical developments in the understanding of the rock mass and fill behaviour in conjunction with dilution measuring techniques, improved blasting, equipment, ventilation and ground support practices currently allows for a successful application of this method in increasingly complex geological and mining situations. In particular, an increased understanding of the method is required to facilitate improved stope access configurations and optimised extraction sequences leading to full orebody recovery while achieving dilution control. The complexity of the method and the current depth of the orebodies being extracted worldwide, suggest that adequate planning and control of the operations are critical to the successful
FIG 1 - A view inside an open stope.
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Unstable shapes
Stope Height
Stable shapes
Stope length FIG 2 - Stable shapes for sublevel stoping.
implementation of optimum stope sizes and sequences of extraction. The method is commonly known throughout the world as open stoping, sublevel stoping, longhole or blasthole stoping. The essential common elements of sublevel stoping are (Bridges, 1982):
• the stopes are open at some stage, without substantial wall collapse or caving;
• the stopes extend from sublevel to sublevel, with operations carried out only at these sublevels;
• the blasted rock moves by gravity alone to the stope drawpoints;
• • • •
the method uses long blastholes for rock breakage; the blastholes are located within planes called rings; the holes are drilled mostly downwards; the initial expansion slot is located on the side or bottom of each stope; and
• the method is non-entry, and personnel do not have access to the open portion of a stope (see Figure 1).
578
FACTORS CONTROLLING STOPE WALL BEHAVIOUR Excavation geometry In sublevel stoping, drilling and blasting is undertaken from drilling drives located at each sublevel along the height of a stope. Because of the limited cablebolt reinforcement that can be provided at the exposed stope walls, the excavations are usually designed to be inherently stable. In this regard, experience has shown that in most cases, it is possible to achieve stope wall stability (with minimal dilution) by either excavating openings having high vertical and short horizontal dimensions, or openings having long horizontal and short vertical dimensions (see Figure 2). The shape of the stability curve is hyperbolic and suggests that for multiple lift sublevel open stopes (excavations with walls that have high vertical and short horizontal dimensions) the critical spans are either the exposed horizontal lengths or the stope widths. Length and width are the critical stope dimensions, as they also control the dimension of the stope crowns. Bench stopes are excavations where the longest dimension is the strike length and the critical spans are usually the exposed heights as the orebody width is usually narrow.
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A REVIEW OF SUBLEVEL STOPING
stoping block) such as faults, shears and dykes usually have very low shearing strength and, if oriented unfavorably, provide a failure surface when exposed by the stope walls (see Figure 3). Such geological discontinuities largely control overbreak and stability around exposed stope walls. In particular, those including platy and water susceptible minerals such as talc, chlorite, sericite and kaolin (Bridges, 1982). In some cases, stoping activities have been linked to instability in concurrent voids along the strike or dip of main geological features such as fault zones (Logan et al, 1992). The location of these main geological discontinuities is well defined and most mines have a three-dimensional model of the local fault/shear network (see Figure 4). These features can also be seismically active, further increasing fall-off at the excavation boundaries. Overbreak is usually very difficult to control in this case, regardless of the blasting practices, and can only be minimised by stope sequencing. Stope wall behaviour is also a function of the number, size, frequency and orientation of the minor scale geological discontinuities. Such discontinuity networks usually control the nature and amount of overbreak at the stope boundaries. Rock mass characterisation techniques can be used to estimate the shape and size of blocks likely to be exposed at the final stope walls (Villaescusa and Brown, 1991; Villaescusa, 1992). The geometrical discontinuity set characteristics (size, frequency, orientation, etc) relative to the stope walls largely control the amount of dilution experienced at those walls (see Figure 5). Individual joints have a limited size and they may either terminate in intact rock (forming an intact rock bridge) or against another structure within a discontinuity network. These intact rock bridges are significantly stronger than the naturally occurring discontinuities and provide a higher resistance to failure within a rock mass.
Rock mass strength It is generally accepted that behaviour at the stope walls is largely controlled by the strength of the rock mass surrounding a stope, which in turn depends upon the geometrical nature and strength of the geological discontinuities as well as the physical properties of the intact rock bridges. Single or combinations of major discontinuities (usually continuous on the scale of a
52
55
º
M
N5 4
5500N
4500N
5000N
FIG 3 - Stope hangingwall behaviour controlled by bedding and joint frequency.
44
80 º
M
J 54 75º
5 T4
S4
8
80
52 O
º 50
53
80º
T
70
º
º
º
75 FIG 4 - Major structures affecting sublevel stoping at the 1100 orebody, Mount Isa Mines (from Alexander and Fabjanczyk, 1982).
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FIG 6 - A large stress-related footwall buckling failure.
In addition, failures due to stress changes of a tensional nature can also be experienced (Bywater et al, 1983). Stope extraction in a destressed orebody may lead to normal stresses of very low magnitude across some of the exposed walls. Buckling type failures may occur in depending upon the frequency of discontinuities parallel to a stope wall, the size and frequency of any cross discontinuities and the size and shape of the exposed spans (see Figure 7).
FIG 5 - Footwall and hangingwall fall-off.
Induced stresses Extraction within a stoping block, can generate large concentrations of stress around the excavation boundaries. If the local (induced) stresses increase beyond the shear strength of the rock mass, changes in the quality of the rock mass around the stope will occur, and localised failures are likely to be experienced either following discontinuity surfaces or directly through intact rock. Where movement through discontinuities occurs, stresses are relieved, however this may in turn lead to overbreak, dilution or large failures (see Figure 6). The change in rock quality (before failure) around the boundaries of a stope is called pre-conditioning of a rock mass. This is a very complex process, which in most cases results from a combination of stress re-distributions, near field blast damage and the effects of the excavation itself. In cases where stope wall failures do not occur due to the initial pre-conditioning by the stresses, vibration and gases from nearby blasting may damage the intact rock bridges, which define (and interlock) the in situ rock blocks, causing overbreak or dilution at the stope boundaries. Furthermore, the dynamic behaviour of an unsupported wall is directly proportional to the amount of intact rock available within the rock mass. The less intact rock available, the more cracking, stabbing and visible stope wall displacement will result from the blasting process. The larger the openings, the larger the excavation-related deformations that are expected at the boundaries, making the walls more susceptible to damage from stress re-distribution and/or blasting.
580
FIG 7 - A large structurally-controlled hangingwall failure.
Ground support Reinforcement by cable bolting provided at selected locations (usually constrained by the distance between drilling sublevels) can be used to reduce the deformations experienced at the final
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A REVIEW OF SUBLEVEL STOPING
stope boundaries (crowns, walls and rib pillars). The stope walls are pre-reinforced prior to any stope firings and, in most cases, the cablebolts are installed from rings drilled within the stope access drives. Thus, stope wall reinforcement tends to be localized in continuous bands that are separated by the distance between the sublevel intervals. The function of such arrangements is to divide the stope walls into a number of stable stope wall spans as well as arresting up-dip hangingwall failures (see Figure 8).
6m Bulbed cables
22B1 S PANEL CMS section
FIG 9 - Stope hangingwall cablebolting from a remote drive (Villaescusa et al, 1995).
Blast damage
FIG 8 - A large stope hangingwall failure arrested by a row of cablebolts installed prior to stope firings.
As an alternative to installing cablebolts from stope drill drives, special drives can be developed around a stoping block solely for cablebolt installation. To decrease cost and increase their reinforcing effectiveness, such horizontal drives are usually located at the same vertical horizon as the drilling sublevels and 10 to 15 m away from a planned stope wall location (see Figure 9). Support from minefill can also be used to minimise the deformations experienced by the stope walls and also to provide restraint of any adjacent rock masses. In general, cemented fill is needed to recover ore from secondary stopes, where stable fill exposures are required to minimise dilution. Cemented fill is essential in chequerboard extraction patterns within massive orebodies (Bloss, 1996), while uncemented rockfill is normally used in conjunction with bench stoping operations (Villaescusa and Kuganathan, 1998). An example of a bench stoping extraction strategy linked to backfill is shown in Figure 10 where the exposed walls are usually limited to a critical length value, which is defined by the distance between the minefill and an advancing bench brow (Villaescusa et al, 1994).
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Damage to a blasted rock mass refers to any strength deterioration of the remaining rock due to the presence of blast induced cracks and to the opening, shearing and extension of pre-existing or newly generated planes of weakness (see Figure 11). It is generally accepted that the damage is caused by expanding gases through the geological discontinuities and to the vibrations experienced from the blasting process. However, it is not easy to establish the approximate contribution to damage caused by the expanding gases, as it is difficult to measure their path within a rock mass discontinuity network (McKenzie, 1999). Nevertheless, significant backbreak may be regularly observed when the explosive gases are well confined within a volume of rock, and in some cases the gases can travel well beyond the explosive charges. Damage by the shock energy from an explosive charge close to a blast can be related to the level of vibrations measured around the blasted volume. Repetitive blastings also impose a dynamic loading to the exposed stope walls away from a blasted volume, and may trigger structurally controlled fall-off and ultimately overbreak. Conventional blast monitoring and simple geophysical techniques can be used to measure the effects of blasting in the near field. Vibrations and frequency levels from the shock wave can be measured reasonably accurately. This can be related to damage provided the contribution (to the total damage) from the shock energy can be estimated. Vibration and frequency levels at the mid-span of instrumented stope walls can be used to characterise the dynamic response to blasting at the stope boundaries (Villaescusa and Neindorf, 2000).
Drill drive layout Additional factors such as poorly located (or pre-existing drives) which undercut the stope walls, also contribute to dilution or fall-off at the stope boundaries. In general, the number and location of drilling drifts in open stoping is usually a function of the width of the orebody. In wide orebodies hangingwall and footwall drill drives can be used to minimise the impact of blasting at the stope boundaries. In such cases, drilling and
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Filling
Bench Limit
Product ion Blast ing
MaximumUnsupport ed Span (Crit ical St rike Lengt h)
Mucking
Backfill Previous Bench
Open stope
Hangingwall
FIG 10 - Continuous filling operations in bench stoping.
Blasthole
FIG 12 - Holes toeing into the stope boundaries.
FIG 11 - Structurally controlled damage around a hole in an open stope brow.
blasting can be done on a plane parallel to the final stope walls or to any exposed backfill masses. Suitable values of stand-off distance (parallel to the stope boundary) for the perimeter holes can be determined, depending upon the rock type and the hole size being used (Villaescusa et al, 1994). On the other hand, excessive wall damage, dilution and ore loss may be experienced in the cases where stoping requires drilling holes at an angle to a planned backfill exposure or a stope boundary (see Figure 12). In this example, hole deviation at the toes may create an uneven stope surface, thereby preventing effective rilling of the broken material to the stope drawpoints. In addition, hole deviation may cause excessive confinement at the hole toes, thus causing breakage beyond the orebody boundaries.
orebody. The method requires a ‘moving’ drawpoint system as the stoping progresses upward. Following the backfilling of a stope void, a drilling horizon becomes the next extraction level (see Figure 13). In order to optimise mucking productivity, up to two access
SINGLE LIFT STOPING A single lift design is the most basic arrangement for a sublevel stope extraction. The stope shape and size is constrained by two sublevels; the extraction and the drilling horizons. Access to the stope is via cross-cuts off a permanent access drive parallel to the
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FIG 13 - Three-dimensional view of single lift sublevel stoping (Potvin et al, 1989).
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cross-cuts per stope may be required at each sublevel interval. This actually increases the overall access development in waste to actual stoping ratio. The method requires very good control of the stope back and brow stability, especially in a highly stressed environment. Stress re-distribution due to the stoping sequence itself can create significant back failures, especially if shallow dipping discontinuities are present within a rockmass. Extended backs and pendant pillars and highly stressed brows are likely to be formed somewhere within the stoping sequence, and full cablebolting coverage is required to minimise the potential failures at each sublevel location. Full cablebolting coverage requires stripping the orebody access to the full stope width, thereby minimising the size of stopes that can be safely developed. As a result, single lift stopes tend to be relatively small openings, compared with multiple lift stoping. Primary development requires extending the access cross-cut to a proposed hangingwall location, where both the drill and the extraction sublevels are completely silled out to allow the installation of cablebolt reinforcement. In addition, the drilling of parallel blastholes is also facilitated with full stope undercut and overcut geometries. Parallel holes is the preferred way of drilling and blasting in vertical retreat stoping, which is linked to single lift stoping. The method requires a significant amount of remote mucking due to the flat-bottom nature of the single lift stope geometries, thereby increasing the overall mining cost compared to a multiple lift stoping geometry. In wide orebodies a number of stopes may occur across the strike in a given area, and in all cases, adjacent primary stopes are extracted to a level above that of a secondary stope. This type of sequence creates what is called a pendant pillar. A pendant pillar is a solid piece of ground that has many degrees of freedom for movement, as most stopes around it have been extracted (see Figure 14). Large pillar failures may be experienced in such stoping geometries (Milne and Gendron, 1990). FIG 15 - Vertical crater retreat within a single lift stope.
Pendant pillar
FIG 14 - Idealised stoping sequence for single stopes in a 1-4-7 extraction sequence.
Conventional vertical retreat stoping Vertical crater retreat (VCR) is a single lift stoping method where the stopes shape is defined by a lower (undercut) and upper (overcut) horizon. Large diameter holes are drilled in order to minimise deviation, and the holes are charged from the overcut and blasted by means of horizontal slices of ore progressing from the bottom level to the top level (see Figure 15). The separation between the undercut and overcut is a function of stope wall competency, nature of the orebody and drilling accuracy.
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Following blasting, only a slight amount of broken ore is mucked, so that enough room is available for a subsequent blast to break into. This keeps the stope full of broken rock, thereby providing passive support to the exposed stope walls, until blasting to the stope overcut is complete. Once blasting is completed and all the ore within the stope is mucked, the undercut accesses are closed off and the stope is filled. As mining progresses upwards, the stope overcut becomes the next mucking horizon in the sequence. The method has a number of perceived advantages including the requirement for few large diameter blastholes, likely to reduce the overall in-the-stope drilling. Large holes enable a larger sublevel interval, thus reducing the overall sublevel development cost. The cost of raising and slashing to create a slot is eliminated, and all the drilling and loading operations are carried from the overcut, thereby increasing safety. The disadvantage of this method is the potential for blast damage from crater blasting at the stope boundaries. Small diameter holes are not used due to hole closure caused by ground movment following the individual stope blasts (Hills and Gearing, 1993). In addition, this method may be susceptible to poor fragmentation (fall-off) from the unsupported areas defined by blasting, especially if an uneven back is formed and high stresses are subsequently redistributed upwards. Blast damage from cratering is even more detrimental when shallow dipping geological discontinuities are present within a rockmass.
Modified vertical retreat stoping A modified vertical retreat method uses a winze or a raisebored hole, which is located near the middle of the stope, into which a radiating pattern of blastholes is sequentially fired in horizontal
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lifts. The raise is used to overcome the limited free face available in a conventional vertical retreat stope. In order to facilitate the initial blasting, the method requires close spacing of the holes near the raise (see Figure 16). All the holes in a horizontal lift are fired, and some danger of collar damage exists when the inner holes near the raise do not perform. In addition, hole damage (closure, requiring re-drilling) within the last lift in the stope may be continuously experienced with this method (Hills and Gearing, 1993). On the other hand, the method is considered a relatively safe method because no vertical opening is made within the stope until the last firing.
• • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • •
Parallel rings of vertical blastholes to be fired in a radiating pattern from the raise
Plan View
• • • • • • •
Raise bored hole 1.1m diameter
Position after one ring firing Section View
FIG 16 - Typical blast layout for a modified vertical retreat stope in the Porgera Mine (Hills and Gearing, 1993).
HANGINGWALL DRILL DRIVES
ENDWALL DRILL RINGS
ACCESS CROSSCUT ACCESS DRIVES EXTRACTION LEVEL
TROUGH UNDERCUT TIPPLE
FIG 17 - Idealised three-dimensional view of a multiple lift sublevel stope.
up, firing rings towards an open slot, from sublevel to sublevel. The blasting sequences can be designed to minimise undercutting the individual sublevels during the stope blastings. A straight face is kept along the entire stope height by firing a similar number of rings at each level. The firing sequence advances upward as shown in Figure 18. Maintaining a straight retreating face along the entire open stope height minimises the creation of brows, which can be highly stressed and contribute to stope fall-off, which in turn can severely affect productivity during the subsequent mucking operations.
Cut-off slot
MULTIPLE LIFT STOPING Multiple lift stopes extend vertically over a number of sublevel intervals, in some cases exceeding hundred of metres in vertical extension. Stope extraction starts by creating a cut-off raise, which is expanded to a cut-off slot and extended over the full stope height. Main production rings are then fired progressively into the void created by the cut-off slot until the stope is completed. Ore breakage is achieved by rings of parallel or fanned blastholes, depending upon the type of drilling access used. Trough undercuts are developed at the base of the stopes in order to direct the broken ore into the drawpoints for extraction. The number of drawpoints is usually a function of the stope size, but in most cases at least two drawpoints are designed. Because the drawpoint location is fixed, permanent reinforcement can be afforded at minimum cost per unit of ore extracted. Access to the stope on each of the other sublevel locations is required for drilling, blasting and backfilling purposes. Usually, a single crosscut access is required at each sublevel, significantly decreasing development in waste (see Figure 17). In general, multiple lift stopes minimise back cable-bolting within the intermediate sublevels because a permanent back (full area) is only exposed at the actual crown of the stope. Consequently, stope crown stability is facilitated as reinforcement may only be required within a finite area on the top of the stope. Cablebolting coverage at the stope crown is a function of the degree of development within the top sublevel. In addition, the requirement for permanent reinforcement within any intermediate sublevel is minimised by the fact that all the back exposures within the drill drives are consumed by the stoping process itself. The stopes are usually sequentially sliced
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Mucking FIG 18 - A long section view of a multiple lift stope extraction.
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The outline for multiple lift sublevel stopes in tabular orebodies is usually associated with the use of long blastholes drilled from drives parallel to the orebodies. Depending upon orebody width these drill drives are either full orebody width or located at the boundaries of the orebodies. In such orebodies the stope boundaries are usually well defined by the orebody itself. Crown, hangingwall, footwall, end-walls and a drawpoint can be defined for each stope. The stability of stope crowns and hangingwalls is usually the most critical factor in the design and extraction sequences. A conventional design usually consists of multiple drilling sublevels with a single mucking horizon at the bottom of the stope as shown in Figure 19. One of the advantages of this design is that drilling and blasting can be done on a plane parallel to the final stope walls. Hangingwall and footwall drill drives are used to minimise the impact of blasting at the stope boundaries greatly decreasing the likelihood of dilution due to blast damage. In addition, the method reduces stope development in waste, given that except for the mucking horizon, a single stope drilling access is actually required at each sublevel location. The stability is enhanced when all the sublevel accesses are excavated after the adjacent (previously extracted) stopes have been backfilled. Open stoping in large, massive orebodies consists of a mining sequence that requires several stages of stoping in conjunction with the application of delayed minefill methods to enable pillar recovery. Usually, this requires that a number of stopes are designed between the orebody boundaries. In such cases, stoping comprises a number of stages which include primary, secondary and tertiary stopes which are usually extracted using a checker board sequence (Alexander and Fabjanczyk, 1982). The number of fill exposures ranges from none (in a primary stope) to up to three exposures within the late stages of stoping.
Large vertical dimensions can be designed with the height of the stopes usually constrained by the orebody thickness or by the stability of the exposed fill masses required for secondary and tertiary stope extraction. Stope dimensions in plan view are usually constrained by stope crown instability. The broken ore is also extracted in the bottom part of the stope (see Figure 20). In cases where the ground conditions are favourable, stope dimensions can be very large in plan, with full orebody height extraction achieved in a single stope (Bloss, 1996). Drilling and blasting is carried out from a series of sublevel locations ranging from 40 - 60 m apart. Blastholes are mainly drilled downwards, with some short upholes drilled within the trough undercuts and sometimes at the stope crown when a top access is not available. Following pillar extraction (secondary and tertiary stopes), a number of backfill exposures are created depending upon the location of the stope in the mining sequence. Pillar stope nomenclature is usually based on the number of exposed backfill masses. As an example a two SLOS stope has two fill exposures. Early in the life of a massive orebody primary stopes usually account for a significant part of the production. As an orebody extraction increases, the shifting to pillar mining as the primary method of extraction is evident. In such cases, the stability of the fill exposures is of primary importance to achieve target production (Bloss and Morland, 1996).
The stope cut-off slot Sublevel stopes are created by the sequential blasting of production rings into an initial opening that is called the cut-off slot. The cut-off slot is located on a side of the stope either transversally (across) or longitudinally with respect to the strike of an orebody. The critical point relates to whether the cut-off blastings will expose a critical stope wall (such a hangingwall or a backfill mass) very early in a stope blasting sequence. The slot
Cut-off Slot
Cut-off Slot
b) Plan View - Intermediate Level
H/W
rein
Cut-off Slot
f or c
eme
nt
a) Plan View - Mucking Horizon
Cut-off Slot
c) Cross Section View - Production rings
d) Long Section View
FIG 19 - Sublevel stoping in a tabular orebody.
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7 5
1
3
4
1
2 8
6
7
FIG 20 - Sublevel stoping in a massive orebody.
•
•
•
•
•
• •
•
• •
•
• •
•
• •
•
• •
•
•
•
•
• •
•
• •
•
• •
• • •
FIG 21 - Sequential blasting of a cut-off slot.
raises are blasted upwards from sublevel to sublevel in order to expose a full stope height. At each level the slots are formed by sequentially blasting parallel holes into a long-hole winze (LHW) or a raised bored hole. The slot must be expanded to the full width of the production holes that will be subsequently blasted into this initial opening (see Figure 21). High powder factors are normally used during slot blasting in order to ensure breakage and thus to have a free face and a void available where the remainder of the stope is to be blasted. The choice of slot location depends upon rock mass conditions, stope access and the extraction sequence chosen. In a steeply-dipping orebody, where the critical stope boundary is usually the inclined hangingwall, transversally oriented slots are used to ensure a
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sequential hangingwall exposure by the production rings. In large, massive orebodies, the choice of slot orientation is also controlled by factors such backfill exposures, stress regime and pre-established access. In general, a slot must be designed so that failure within the main production rings is minimised. In highly stressed pillars a slot can be oriented normal to the main principal stress to shadow the main production holes. This is likely to minimise hole squeezing or dislocation due to stress related damage. In cases where a stope access can be re-designed, the slot can be placed normal to geological features likely to fail and damage the main ring geometries (see Figure 22).
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P4 1 º
Fa u
65
º
P4 1
lt
º 65
Fa u
º
lt
Cut-off slot
Cut-off slot slot Cut-off
65
65
/W 1F P4
ay spl
ay spl
/W 1F P4 FIG 22 - Exposure of weak geological features by a cut-off slot (After Rosengren and Jones, 1992).
Cleaner ring
Production rings
Blasting Sequence 1. Holes near raise
Extracted (filled)
2. Mid cut-off 3. Complete cut-off + Cleaner ring Cut-off slot Plan View FIG 23 - Cleaner ring geometry to minimise fill damage from blasting.
Damage to fill masses from cut-off slot blastings can be minimised by placing a cleaner ring between the cut-off and the fill boundary (see Figure 23). The rockmass adjacent to a fill mass is usually pre-conditioned by stress re-distributions and it is likely to fail following the cleaner ring blasting. This practice, however, minimises dilution from fill failures, as the cut-slots are blasted away from the fill masses. Experience suggests that (except for highly stressed pillars) stope hangingwall failures are unlikely to occur during a transversely oriented cut-off slot formation, because the exposed hangingwall planes shapes are likely to be within a stable range of exposures (see Figure 2). The risk of a hangingwall failure during slot blasting increases when the hangingwall is undercut by the stope development or when the stope is undergoing large stress changes. In cases where longitudinal cut-off slots are
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located parallel (and adjacent) to a stope hangingwall, the slot exposes the full hangingwall plane early in the stope life. This usually limits the size of exposures that can be safely excavated, as this critical wall of the stope may fail when subjected to repetitive dynamic loading by the rest of the stope firings as shown conceptually in Figure 24. In addition, such stoping geometries are likely to toe holes into the adjacent backfill masses, thereby increasing the likelihood of dilution. On the other hand transversely oriented cut-off slots create a full exposure at one of the stope end-walls early on the stope life. In cases where the stope is located within a highly stressed environment, failures within such walls may occur. Although, strictly speaking, fall-off from stope end-walls does not constitute dilution, the stope mucking productivity can be severely affected, especially if the size of a failure is such that secondary breakage is required (Anderson and Leblenc, 1995).
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Q426 fill mass
Half cut-off slot
Main Rings
R422 fill mass
slot
Expansion Rings
P422 fill mass
Filled
Filled
········ ··· ···
Hangingwall FIG 24 - Dynamic loading of a fully exposed hangingwall plane. Diaphragm ring drill access (new development)
Expansion ring drill access (existing development)
The production rings
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FIG 25 - Influence of existing development and adjacent stoping on a stope firing sequence (Bloss and Morland, 1996).
140mm blasthole
· · · · · · · · · · · Remainder of
stope extracted
Fill mass
A designed stope shape is achieved by sequentially blasting rings of blastholes into the opening created by the cut-off slot. The main rings are sequentially blasted on each sublevel attempting to minimise undercutting of the internal solid portion of a stope. Avoiding undercutting usually reduces fall off from retreating stope brows as damage from stress re-distribution is minimised when a straight face is maintained along the entire stope height (see Figure 18). Conventional multiple sublevel stoping requires the sequential exposure of high vertical, short horizontal stope walls likely to remain stable and provide undiluted ore. The strike lengths exposed during the initial stoping extraction are unlikely to exceed the critical unstable stope spans. As the excavations are enlarged and several rings are sequentially blasted into the void formed by the cut-off and the initial production rings, confining stresses are reduced, and excess strain energy is induced and displacement of the stope walls is experienced. Depending on the structural nature of the exposed walls, the rock may tend to displace following a sheet-like behaviour, in which a group of layers move together (bedded rock), or the movement may be isolated to individual blocks which partially rotate and slide against each other. Once enough room is available, the most appropriate way to blast the rest of a stope can be considered, depending on the circumstances such as the level of the induced stresses or production requirements and access constraints (see Figure 25). As production blasting continues towards the final stope geometry (shape and size), the excavation becomes more unstable. Geological discontinuities as well as the dynamic impact of blasting begin to affect stope wall stability and contribute to dilution. Diaphragm rings consist of rings drilled parallel to a fill exposure. The purpose of a diaphragm ring is to prevent fill failure from a known weak cemented fill mass, to contain uncemented fill in adjacent stopes or to prevent fill failure from exposures of a greater dimension than is considered stable. Experience has shown that although parts of a diaphragm against backfill do fall off, this rarely results in excessive backfill dilution, as the fill mass remains comparatively undisturbed, compared to when blasting takes place next to the backfill (see Figure 26). A diaphragm is not capable of load bearing capacity and it is likely to deform considerably. However, when a large portion of the diaphragm remains intact, this enables clean stope extraction until the diaphragm is either fired or the stope is completed.
Main ring drill access (new development)
· · · · · · · · · · ·
2m from fill
3m burden on diaphragm ring FIG 26 - Idealised sketch showing a diapraghm ring.
The stope undercut The lower portion of a stope is shaped using trough undercut (TUC) rings in order to facilitate the draw of fragmented ore to the stope drawpoints and to minimise remote mucking. A TUC ring consists of parallel upholes drilled inclined towards the cut-off slot. Usually the toes of the TUC ring interlock with the
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toes of the main ring downholes from the sublevel above (see Figure 27). Drilling and blasting of the trough undercuts is usually carried out using relatively small diameter holes (70 89 mm). An improved explosive distribution likely to minimise rock mass damage around the stope drawpoints is achieved by using such small diameter holes. A disadvantage is the limited drilling length achieved, and the inability to match the burden drilled for the production ringholes immediately above
Drawpoints
Orebody
14 13 10 7 4
Hangingw
all
FIG 28 - A plan view of a fixed drawpoint geometry in sublevel stoping.
CONCLUSIONS
3
An understanding of the different factors controlling stope wall behaviour, as well as the response of a rockmass to the different styles of stoping geometry is critical to the successful application of sublevel stoping. In particular, single lift stoping usually requires significant lateral development where extended highly stressed backs and pendant pillars are likely to be created. Consequently full cablebolt coverage may be needed at each sublevel location, as the method requires moving drawpoints for mucking. In addition, a significant amount of remote mucking is required due to the flat-bottomed nature of the single lift stopes. On the other hand, multiple lift stoping requires significant vertical development while allowing the use of specialised trough undercuts and drawpoint geometries. Furthermore, reinforcement within the stope itself is minimal and limited to the stope crown, hangingwalls and drawpoints, but not necessarily required within the intermediate sublevels. More importantly, this method minimises stope wall dilution and fall-off from rings within the stope itself by firing a similar number of rings on each sublevel and creating inherently stable excavation shapes.
14 12 9 6
2
11
8
5
1
Longitudinal View FIG 27 - Firing sequence of a trough undercut with main rings in an open stope.
Because the TUC rings are drilled with a different burden to the main rings, the lower portion of the stope is usually blasted ahead of the main rings. This leads to undercutting of the main rings, which can lead to fall-off, especially in cases where large geological discontinuities are present or in regions of high stress re-distribution.
The drawpoints Mucking in multiple lift sublevel stoping is usually carried out transversely across the strike of the orebodies. This requires the introduction of fixed and specialised drawpoint geometries that may be located outside the orebody boundary (see Figure 28). The factors considered during drawpoint design, include size of equipment, tramming distance from access drives, gradient and orientation with respect to a stope boundary. The drawpoint dimensions must be sufficient to suit the equipment, but kept as small as possible to minimise instability. Drawpoint access should be straight and restricted to 15 - 20 metres in length from the stope access drive to the stope brow. This will ensure that auxiliary ventilation will not be required while mucking, and also that rear of the mucking unit is inside the drawpoint. Drawpoint spacing is determined by ground conditions and stope geometry. In most cases the minimum spacing used is 10 - 15 m between centre lines.
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ACKNOWLEDGEMENTS The author would like to gratefully acknowledge Mount Isa Mines Limited for the permission to publish the photographs shown in this paper.
REFERENCES Alexander, E and Fabjanczyk, M, 1982 Extraction design using open stopes for pillar recovery in the 1100 orebody at Mount Isa, in Design and Operations of Caving and Sublevel Stoping Mines, (Ed: D Stewart) pp 437-4458 (SME: New York). Bloss, M, 1996. Evolution of cemented rockfill at Mount Isa Mines, Mineral Resources Engineering, 5:(1):23-42. Bloss, M and Moreland, R, 1995. Influence of backfill stability on stope production in the Copper Mine at Mount Isa Mines, in Proceedings Underground Operators’ Conference, pp 237-241 (The Australasian Institute of Mining and Metallurgy: Melbourne). Bridges, M C, 1982. Review of open stope mining, AMIRA Project 81/P138, 80 p. Bywater, S, Cowling, R and Black, B, 1983. Stress measurements and analysis for mine planning, in Proceedings Fifth ISRM Congress, Melbourne Australia, D29-D37. Hills, P and Gearing, W, 1993. Gold ore mining at Porgera, Papua New Guinea, in 11th CIM Underground Operators Conference, Saskatoon, Saskatchewan, 18 p. Logan, A S, Villaescusa, E, Stampton, V, Struthers, M and Bloss, M, 1993. Geotechnical instrumentation and ground behaviour monitoring at Mount Isa, in Proceedings Australian Conference Geotechnical Instrumentation and Monitoring in Open Pit and Underground Mining, Kalgoorlie, pp 321-329.
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McKenzie, C K, 1999. A review of the influence of gas pressure on block stability during rock blasting, in Proceedings Explo 99, pp 173-179 (The Australasian Institute of Mining and Metallurgy: Melbourne). Milne, D and Gendron, A, 1990. Borehole camera for safety and design, in Proceedings 92nd CIM Annual Meeting, Ottawa, 13 p. Potvin, Y, Hudyma, M and Miller, S, 1989. Design guidelines for open stope support, CIM Bulletin, 82(926):53-62. Rosengren, M and Jones, S, 1992. How can we improve fragmentation in the Copper Mine. Unpublished Mount Isa Mines Limited Internal Report. Villaescusa, E and Brown, E T, 1991. Stereological estimation of in-situ block size distributions, in Proceedings 7th International Congress on Rock Mechanics, Aachen, West Germany, pp 361-365. Villaescusa, E, 1992. A review and analysis of rock discontinuity mapping methods, in Proceedings 6th ANZ Conference on Geomechanics, Christchurch, New Zealand, pp 274-279.
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Villaescusa, E, Neindorf, L B, and Cunningham, J, 1994. Bench stoping of the Lead/Zinc orebodies at Mount Isa Mines Limited, in Proceedings Joint MMIJ/AusIMM Symposium, New Horizons in Resource Handling and Geo-Engineering, Ube, Japan, pp 351-359. Villaescusa, E, Karunatillake, G and Li, T, 1995. An integrated approach to the extraction of the Rio Grande Silver/Lead/Zinc orebodies at Mount Isa, in Proceedings 4th International Symposium on Mine Planning and Equipment Selection, Calgary, (Ed: R K Singhal) pp 277-283 (Balkema). Villaescusa, E and Kuganathan, K, 1998. Backfill for bench stoping operations, in Proceedings Sixth International Conference on Mining with Backfill, pp 179-184 (The Australasian Institute of Mining and Metallurgy: Melbourne). Villaescusa, E, and Neindorf, L B, 2000. Damage to stope walls from underground blasting, in Proceedings International Conference Structures Under Shock and Impact VI, Cambridge (Eds: N Jones and C A Brebbia) pp 129-140 (WIT Press).
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Mount Isa Mines - 1100 Orebody, 35 Years On D Grant1 and S DeKruijff2 ABSTRACT The Mount Isa Mines Copper Mine in Mount Isa started producing copper in 1966. The Copper Mine gradually increased production rates to nearly six million tonnes per year. Over the last three years the mine has been in a gradual ramp down due to the reduction of primary ore sources and head grade. The majority of ore sources are predominantly in pillar retreat sequences bounded by multiple fill masses. There also a substantial lower grade resource delineated to the south of the main orebody. These remaining ore sources present significant engineering challenges to obtain a safe and cost-effective extraction. Stope designs have become increasingly complex and require additional engineering and operations input and analysis. This paper describes the major planning and operational challenges and issues associated with the stope design and extraction of an orebody in the later stages of mine life.
INTRODUCTION This paper describes sublevel open stope design at the Copper Mine and the operational issues considered to improve the overall mine planning process and hence the viability. The Copper Mine commenced production in 1966 and gradually ramped up to in excess of five million tonnes per year (mty) over a ten year period. During the last three years the mine production has reduced to approximately 3.5 mty because of the increased dependence on pillar extraction sequences and the increased reliance on truck haulage. The Copper Mine orebodies extend for nearly 3000 metres north to south, up to 500 metres east to west and vary in height from 30 metres to in excess of 250 metres. Figure 1 shows a plan of the Copper Mine orebodies. The current challenge is to economically mine the remaining Copper Mine orebodies with a continuing drop in head grade and increasingly complicated stoping geometries.
GEOLOGY The geology of the Mount Isa Copper Mine has been described in detail in numerous technical papers (Perkins, 1984; Bell, Perkins and Swager, 1988). The total past production and published reserves from the Isa copper orebodies as of 1981 was approximately 255 million tonnes at 3.3 per cent copper (Gustafson and Williams, 1981). These orebodies are hosted within Urquhart Shale Sedimentary rocks of Precambrian age. These rocks have undergone structural deformation and chemical alteration resulting in a siliceous and dolomitic halo around the mineralisation. Copper ore occurs as disseminated and massive chalcopyrite. The Urquhart Shale sequence is a 1100 m thick package of thinly bedded black, pyritic and dolomitic shales that typically strike north south and dip to the west at 65 degrees. The main Copper Mine orebody is the 1100 orebody, which commences in the north and extends approximately 2000 metres to the south where the orebody splits into a hangingwall and footwall lens. The 1900 orebody is positioned to the footwall of the 1100 orebody towards the north end. Figure 2 shows a typical cross-section through the 1100 orebody. The Paroo Fault 1.
MAusIMM, Copper Mine, Mount Isa Mines Limited. E-mail [email protected] or [email protected]
2.
Copper Mine, Mount Isa Mines Limited.
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FIG 1 - Plan view of the Copper Mine showing shaded extracted stopes and future stopes to be mined.
(Basement Contact Zone) separates the copper orebodies and the basement volcanic rocks. The basement contact zone consists predominantly slatey shales and buck quartz with some carbonaceous mylonite. This zone varies significantly in thickness and orientation from a thick nearly horizontal plane to a thin near vertical tight contact. The basement contact zone is typically found in the drawpoint and footwall development areas of the Copper Mine (Brook and Struthers, 1990).
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FIG 2 - Typical geological cross-section in the Copper Mine.
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Two main groups of post ore faults intersect the orebody in stope crowns and throughout the height of stopes. One set dips to the south west between 40 and 70 degrees while the other dips to the north west and has a similar range of dip angles. These faults are continuous and range in thickness from two to twenty metres. The impact of these faults on the operation is discussed throughout this paper.
MINING METHOD The mining method utilised at the Mount Isa Mines Copper Mine is sublevel open stoping (SLOS), which has evolved over the last 70 years of operation, in the copper and lead orebodies, to the
present day design standards. Stope dimensions are typically 40 by 40 metres in plan and are extracted vertically to the full height of the orebody. The sublevel interval used is approximately 40 metres which limits the length of blast holes. The mining method is discussed in detail in numerous technical papers (Hall, 1992; Horsby and Sullivan 1978). Figure 3 shows the typical stoping arrangement used at the Copper Mine. Variations of these dimensions are becoming more common as the complexity of the orebody increases and difficulties that arise when adapting existing development into the stope design. Stope drawpoint configurations vary greatly depending on existing development and the stope extraction strategy.
FIG 3 - Typical sublevel open stoping arrangement used at the Copper Mine.
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Initial Copper Mine production commenced at the northern end of the orebody with primary stopes followed by secondary extraction of stope pillars with multiple fill mass exposures (Alexander and Fabjanczyk, 1981). The mining front moves in a north to south direction. The significant changes in stope design have largely resulted from the development of new technologies, utilisation of higher performance mining equipment and a better understanding of how the ground responds to mining. The main advance in mining technology has been the design of large, low-cost, stable fill masses. As many as three fill masses 40 metres wide and up to 250 metres in height have been exposed concurrently. There has been numerous technical papers written about the filling system at Mount Isa Mines (Dayton, 1978; Neindorf, 1983; Grice, 1989; Bloss, Cowling and Meek, 1993).
STOPE SEQUENCE The stope sequence in the Copper Mine moved in a north to south direction with development continually stepping out to access primary stoping blocks. Secondary stope pillars with one, two or three fill mass exposures were then produced constrained by access and stope turn around time (Kirby, Blair and Hutton, 1978). The sequence continued to step out in a southerly direction leaving large, 40 metre by 300 - 400 metres, east-west traverse pillars for access, ventilation and services. Figure 1 shows the remaining unmined areas including the status of the transverse pillars. Stress analysis computer models are used to model various stoping sequences to minimise and/or delay effects on infrastructure. In recent years the critical stopes in the sequence have been the first stopes to break the transverse pillars, which results in large changes to the stress field. Slow gradual changes to the stress field are believed to be easier to manage. The last major traverse pillar was successfully ‘broken’ last year with a stope located near the hangingwall of the orebody. Current stope turnover times in the pillar retreat areas are approximately six months between the completion of production and the adjacent stope starting up. This includes time for building fill bulkheads, stope filling and drainage, primary mining, fill mining, rehabilitation, production drilling and constructing ventilation controls. This varies depending on the stope size, amount of rock mass damage incurred from stope firings and location of future development. There are three main transverse pillars remaining in the 1100 orebody and three in the Footwall Lens. The Hangingwall Lens will be mined as a south to north retreat sequence with a few early primary stopes. There are a number of other minor stoping areas in the 1100 and 1900 orebodies and Footwall Lens that help supplement production, but these sources only make up a small percentage of total production requirements. Where possible activities are scheduled to minimise interaction. This requires close contact with operations to take advantage of opportunities as they occur and to make the appropriate adjustments arising from delays. The future stope sequence will be largely constrained by the turn around time achievable in the pillar retreat areas. Other areas will be constrained by difficult mining geometries, complex geological structures and development.
STOPE DESIGN CURRENT PRACTISE The current stope design practises at the Copper Mine have been described in detail (Hall, 1992; Evans, 1988) and have not changed significantly. The major changes have been to adapt stope designs to accommodate existing development, the increased proportion of poor ground conditions and the number of fill mass exposures as mine life has progressed. Drive size to adjust to new advances in mining equipment have also changed during the last ten years and will be discussed in the stope extraction section of this paper.
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Development The presence of pre-existing horizontal and vertical development within the area of a new stope design presents the challenge of how to best utilise the development if it is accessible and safe to use. The presence of existing development can affect the entire stope design process from a safety perspective through to development, production drilling, stope firing and finally stope filling. If the development is accessible, then the stope design will be adjusted to optimise the existing development. This may require the location, size and orientation of the cut-off slot to change on each sublevel of the stope. As a result, production drilling and stope firings can become more complicated. Figure 4 shows an example of a stope (Poniewierski, 1998) with a different cut-off development on each of the stope’s four accessible sublevels (19L, 18B, 18E and 17D). The cut-off slot is shown as a dark shaded area on each sublevel. Two additional sublevels were not accessible between the extraction level 19L and mid height sublevel 18B. Production drilling was adjusted to extract this region of the stope from above and below. Figure 5 is a three-dimensional isometric view looking in a north east direction. Figure 4 shows the four sublevels accessed for drilling and the two sublevels (19A and 19C) that could not be accessed. Inclined oriented uphole and downhole cut-off slots have also been successfully adapted to existing development. When new stope development needs to pass through old vertical development a high degree of caution needs to be taken as the development approaches the ventilation rise or ore pass. Probe drilling is conducted to locate the position prior to mining the breakthrough. Once the drive has been mined into the vertical opening, concrete plugs are placed in the floor and the back to secure the excavation before continuing with the development. Horizontal and occasionally inclined development is encountered. Stope development in this situation needs to intersect the existing development (at the same elevation) or pass above or below leaving a sufficient pillar between the excavations. Additional ground support may be required to help reinforce the pillar area. Horizontal pillar dimensions also need to be maintained as large as possible to minimise deterioration during the stope firing.
Production drilling A number of difficulties arise with production drilling in pillar retreat stopes. Difficult drilling conditions are encountered through faults that have undergone blast vibration damage from repeated stope firings and fill water permeation along these fault zones. The remaining stope pillars are commonly in areas of low stress which promotes further opening of faults. Large 140 millimetre holes are used to reduce the effect of hole closure. Where possible drilling is oriented perpendicular to the structure and horizontal drill holes are minimised. The increase in drilling shadows caused by existing vertical and horizontal development increases the amount and complexity of drilling. This requires an increase in out of plane drilling (dipping and dumping holes) and purpose placed holes that bypass inaccessible development to a specific target area. An increase in the number of uphole cut-off slots has been used to reduce the length of down holes from the above sublevel. Shorter down holes are more accurate and more likely to intersect target areas. In areas where development has been mined perpendicular to the direction of the production drilling, difficulties are created for collaring of flat drill holes due to the loss of the wall. The ring may have to be drilled off angle in order to obtain a wall to collar into, or the construction of a wall maybe required to provide something solid to collar low angle drill holes above the floor.
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FIG 4 - Shows an example of a stope with a different cut-off slot layout on each sublevel in the stope (19L Extraction, 18B, 18E and 17D). The cut-off slot is shown as a dark shaded area, blast hole rings retreating towards the stope access.
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FIG 5 – Three-dimensional isometric view showing the four accessible sublevels (19L Extraction, 18B, 18E and 17D) and the two inaccessible sublevels (19A and 19C).
Stope firing The stope firing sequence is related to the location of the cut-off slot development and the ring firers access to the cut-off and rings. In the pillar retreat stopes, the cut-off is positioned as far from the retreat access as possible. This is not always possible and toes can be fired to minimise the complexity of the final cut-off firing. However, in some cases up to four rings have been fired behind the cut-off due to access constraints. The firing sequence is designed to minimise undercutting and maintain a simple retreat type firing sequence for the stope. Vertical walls
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within the stope generally provide the most stability and reduce the amount of in stope fall off. Operational experience has found that smaller firings tend to produce better fragmentation than larger mass blasts. Undercutting has been successfully used in the later stages of a stopes life, where a major geological structure is located in the stope crown. The delayed exposure of the fault can defer the handling of oversize material until the end of the stoping cycle, thereby reducing the impact on production. Blast vibration levels are generally not a problem in the pillar retreat stopes. The surrounding fill masses tend to dampen the peak particle velocity to acceptable levels. Charge weights per
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delay are limited to 1000 kilograms and are generally substantially less. The Orica MS series 1-25 in conjunction with the ConnectaDET provides an extended delay series to reduce charge weights. However in the Footwall Lens stoping is still stepping out to the south and primary stopes tend to generate higher vibrations from relatively smaller firings. This is believed to be caused by the interaction of blast vibrations with major geological structures and the virgin state of the rock mass around primary stopes. The charge weight per delay criteria has been modified to generate lower peak particle velocities. Stope designs also work towards minimising the need for large stope firings.
single or twin strand Garford bulb cable bolts. The cable bolts are either installed manually or mechanically with the use of the Tamrock Cabolter. The drive size of development has become larger to accommodate the new bolting equipment. The use of shotcrete (either plain or steel fibred) to rehabilitate and stabilise areas with high backs has increased significantly, providing timely temporary support for production drillers and ring firer’s. Shotcrete is also used as a longer-term support in areas that intersect the basement contact zone and orebody faults. Ground support is matched with current and future predicted ground conditions. The use and the longevity of the excavation are also considered when selecting the ground support.
Location of old stopes and fill masses
Ground conditions
During the design process, the position of previously produced stopes and the quality of the fill material are taken into account when determining extraction options for the new stope. Prior to 1994 stope surveys were performed using various methods, which gave an approximate three-dimensional estimate of the stopes final shape and location. After 1994, Optech surveys have been conducted, producing a more accurate three-dimensional stope survey. This information is critical to determining the various extraction options, location of stope development and location of temporary pillars to minimise fill dilution and maximise ore recovery. In the case of older stopes, probe drilling is required to obtain the final overbreak profile of the stope; however this may provide only limited information, depending on the access available to perform the probe drilling. Sometimes probe drilling is not possible and stope development goes ahead, resulting in design changes if fill masses are encountered. The cemented fill masses exposed in development are stable and generally do not require additional support. Design changes also occur at the production drilling stage where drilling intersects fill prior to reaching the target depth. The impact on the stope design is usually lessened when Optech surveys have been performed; however stope fall-off can occur during the filling process. Stope fall-off that occurs after the Optech survey alters the final three-dimensional stope shape and usually occurs where major geological structures intersect the walls or crown of the stope.
A series of southwest and northwest dipping (40 to70 degrees) faults cut through the Copper Mine orebodies. These faults are continuous and range in thickness from two to twenty metres. Figure 6 shows a cross section through one of the remaining transverse pillars with the faults intersecting cross-cut development. Typically, mining development has attempted to cross-cut the structures to reduce the affects of the poor ground conditions. In the transverse pillar areas, these structures have been repeatedly rehabilitated after adjacent stope firing and filling events. Some areas have formed into large cathedral shaped openings with wedge geometries peaking at seven to nine metres. Maintaining access through these areas for subsequent stoping has been accomplished using a combination of steel fibre reinforced shotcrete and cablebolts.
STOPE EXTRACTION CURRENT OPERATING PRACTISE Ground support During the last 30 years a wide spectrum of ground support systems and standards have been in place at the Copper Mine. Previous ground support systems have included splitsets, swellex, fully encapsulated resin and cement grouted rebar (with and without plates), plain strand and Garford cablebolts, shotcrete and concrete reinforced pillars and brows. Primary ground support practices at the Copper Mine (indeed across all the Isa and George Fisher operations) adopt fully mechanised one pass support systems, providing immediate support (Potvin et al, 1999). The system also includes the installation of sheet mesh when required. Development headings are bolted every cut and the practise of hand bolting has been significantly reduced. The primary ground support currently in use consists of split sets and ungrouted PAG bolts (point anchored rebar with specially designed shell for short term support requirements. Fully encapsulated cement grouted PAG bolts are used for the long-term support. The equipment used to mechanically install the primary ground support is dedicated bolting machines or development jumbos. Secondary reinforcement consists of either
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Stope filling The stope filling component of the mining cycle is critical for reducing stope turnaround time in the pillar retreat sequences. The fill requires draining and curing prior to being exposed in the subsequent stope. This is incorporated into the stope schedule and constrains the overall production rate. During the filling process large volumes of water are released into the rock mass which can pick up corrosive ions from the in situ copper ore and attack any exposed ground support (Robinson and Tyler, 1999) and air and water services. The time of ground support installation (before or after the stope firings) will determine the likelihood of cracking in the grout column. An assessment of the rehabilitation required is made during the fill mining stage. Efforts are being made to increase the amount of dry fill used and thereby reduce the quantity of water introduced to the rock mass. Containing the fill during the stope filling process can also be difficult in areas. In the transverse pillars, large extension cracks have developed during previous stope firings to the north and south. The extension cracking provides conduits for the fill to flow along. The leaking areas need to be identified and then grouted before the stope can re-commence filling. Fill water may also migrate to producing areas causing the muck to become saturated and somewhat mobile. The position of the fill bulkhead will determine the amount of fill mining required to re-establish access for production drilling. The quantity of fill mining can be significant and is further complicated by the rehabilitation of the ground support.
Mine services The access to pillar retreat areas and the availability of services will impact on the scheduling of mining activities. In pillar retreat sequences there is a high level of activity required to reduce stope turnaround times. This is complicated by the demand for resources from the various mining departments to complete their tasks. Compacting numerous activities into
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FIG 6 – Cross-section through a transverse pillar showing extracted stopes and stopes to be mined, geology (structures and lithology), stoping access and infrastructure.
confined areas is difficult because the supply of services has been reduced as cross orebody accesses have been broken. Air, water, drain, power and ventilation can only be supplied from the one access point. This reduces significantly the possibility of tasks being performed concurrently, adding to downstream production pressures. The increased use of drill holes to link power cables betweens sublevels has been successful for fast tracking critical activities. A recently implemented scheduling computer package will allow a detailed operational focus on the activity schedule. This schedule can be updated as changes occur at the operational level and output critical path activities.
MINE INFRASTRUCTURE The mine infrastructure at the Copper Mine is composed of ore handling, fill reticulation, ventilation, mine services and man and supply shaft systems. These systems are critical to ongoing production of the Copper Mine. The ore handling, reticulation and man and supply shaft components will discussed in this section.
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the the the fill be
Ore handling The Copper Mine ore handling system currently consists of two main components. The ‘Spider’ ore pass system is comprised of five ore passes that feed directly to the Copper Mine primary crusher. The remaining production is chute loaded into trucks and hauled to two of the Spider ore passes where it is tipped into the crusher system. Approximately half of the ore is produced from each component. The ore is then crushed and feed onto a cable belt conveyor that transfers the ore two kilometres to the copper hoisting shaft. Figure 7 shows the Copper Mine ore handling flow diagram. The mining of the pillar retreat sequences is beginning to encroach on the ore pass pillars and additional design considerations are required to maintain the integrity of these ore pass pillars. An adequate pillar, depending on ground conditions, is required to maintain the integrity of the ore pass pillar. When the stope is filled, strict procedures are adhered to monitor the bulkheads and surrounding rock mass for fill leaks. The adjacent ore passes are monitored at various access points and in the control chains area of the Spider.
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FIG 7 - Shows a schematic diagram of the Copper Mine process flow throughout the mining cycle.
Large tonnages of ore have been tipped into the ore passes. The amount of wear is usually estimated by back calculating tipped amounts into an empty pass. It is difficult to quantify areas of major wear, due to the geometry of fingers to the main pass and the tipping history of the ore pass (choke fed versus tipping into an empty ore pass). The ore pass camera has been a useful tool for identifying areas of high wear. Sometimes probe drilling in areas where key access is close to the ore pass may be required. When an ore pass does come close to the main access, monitoring is performed and development bypasses may be considered.
Fill reticulation As the extraction of the 1100 orebody approaches completion a massive crown surface is being exposed. As the transverse pillars are retreated the vertical support of the 1100 ore body will be largely removed. Figure 1 shows the location of the mined out area of the 1100 orebody and the unmined stopes remaining. Major periods of subsidence in the northern 1100 orebody have occurred in the past resulting in damage to infrastructure and production delays (Logan, 1992; Tyler 1999a). This subsidence has been largely related to lack of tight filling of stope voids and the proximity of major geological structures. The current practise is to tight fill stope voids as soon as possible. Tight filling of the pillar retreat stopes will be critical in order to maintain regional stability and sustain high production rates from these areas. Precise levelling surveys are performed routinely to monitor regional movement.
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Man and supply shaft The man and supply shaft that services the Copper Mine orebodies has a shaft pillar that exceeds 200 metres in diameter. As Copper Mine production has extended south the amount of extraction to the west of the shaft has increased. Initial movement in the shaft lining commenced about five years ago and is believed to be related to the extraction of a group of stopes approximately 200 metres to the south west of the shaft (Tyler, 1999b). Figure 1 shows the location of shaft relative to 1100 orebody. The ground movement that has been translated through the rock mass into the shaft lining and steel guide rails shows both extension and compression in the shaft. The different modes of movement are believed to be linked to the location of major geological structures relative to different areas in the shaft. The movement tends to occur in a gradual fashion rather than large or sudden steps associated with stope firings. Movements in the shaft have been managed during the routine shaft maintenance program to adjust the steel guide system and scale loose material. The amount of movement experienced to date is well within engineering guidelines and an order of magnitude less than shaft movement experience in documented South African mines of similar scale and extraction (Stacey and Page, 1986; Spearing, 1995). Numerous studies have been conducted to predict the future movement the shaft is likely sustain as mining proceeds. This is a complex problem because a large percentage of the shaft lining
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deformation is related fault slip movement. The basic approach which has been adopted, is to monitor ground movements in and around the shaft using closure stations, extensometers, guide measurements and a small micro seismic system. The measured movement is then correlated to stoping activity and calibrated with numerical models to predict future deformations in the shaft. This information can be used to look at opportunities to alter the stoping sequence and determine if and when future decommissioning of the shaft may be required.
CONCLUSIONS This paper has described some of the challenges faced in designing and operating in a mine that is approaching full extraction. The Copper Mine at Mount Isa is now confronting increasingly complex stope designs that need to adapt to changes during the operating life of the stope. Copper Mine large tonnage open stopes are a combination of fill constrained pillar retreat sequences and other production sources which require re-handling ore with trucks. Efforts to reduce mining costs and stope cycle times are vital initiatives required to counter the effect of a dropping head grade. The Copper Mine is also presently focusing on mine development in the Footwall and Hangingwall Lens areas at the southern end of the mine to establish infrastructure. Production from these areas will tip to chutes or side load directly into trucks. The trucking system is currently being upgraded to match the tonnage rates required. Areas currently being investigated are narrow, flat lying orebodies with faulted material in the crown areas. These may require alternate mining methods to achieve safe economic extraction.
ACKNOWLEDGEMENTS The authors would like to thank Mount Isa Mines Limited for permission to publish this paper and the assistance with the figures. The authors would also like to acknowledge the effort and team work provided from the operating and technical personnel of the Copper Mine over the years.
REFERENCES Alexander, E G and Fabjanczyk, M W, 1981. Extraction design using open stopes for pillar recovery in the 1100 orebody at Mount Isa, Design and Operation of Caving and Open Stoping Mines, (Ed: D R Stewart) pp 437-458 (AIME Society of Mining Engineers. New York). Bell, T H, Perkins, W G and Swager, C P, 1988. Structural controls on development and localisation of syntectonic copper mineralisation at Mount Isa, Queensland, Econ Geol, 83:69-85. Bloss, M L, Cowling, R C and Meek, J L, 1993. A procedure for the design of stable cemented-fill exposures, MINEFILL 93 (The South African Institute of Mining and Metallurgy).
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Brook, M D and Struthers, M A, 1990. The basement contact zone at Mount Isa - its impact on rock mechanics and mine design philosophy, in Proceedings Mine Geologists’ Conference, pp 43-48 (The Australasian Institute of Mining and Metallurgy: Melbourne). Dayton, S, 1978. Mount Isa mixes multiple concepts of mining with advanced fill technology, Engng Min J, 179(6): 94-107. Evans, R D, 1988. Current drilling and blasting practice in the 1100 orebody - Mount Isa, in Proceedings Explosives in Mining Workshop, pp 63-67 (The Australasian Institute of Mining and Metallurgy: Melbourne). Grice, A G, 1989. Fill research at Mount Isa Mines, Innovations in mining backfill technology, (Ed: Hassani et al) pp 15-22 (Balkema: Rotterdam). Gustafson, L B and Williams, N, 1981. Sediment-hosted strataform deposits of copper, lead and zinc, in Economic Geology, Seventy-fifth Anniversary Volume (Ed: B J Skinner), pp139-178 (The Economic Geology Publishing Company: El Paso, TX.) Hall, B E, 1992. Copper ore mining at Mount Isa Mines Limited, Mount Isa, Qld, Proceedings of the Underground Mass Mining Conference, pp 173-177 (South African Institute of Mining and Metallurgy). Horsby, B and Sullivan, B J K, 1978. Excavation design and mining methods in the 1100 orebody: Mount Isa mine: Australia, in North Queensland Conference 1978, pp 171-181 (The Australasian Institute of Mining and Metallurgy: Melbourne). Kirby, R W, Blair, J R and Hutton, R C, 1978. Development of a pillar extraction strategy for 1100 orebody, Isa Mine, Mount Isa, Australia, in Proceedings of the Eleventh Commonwealth Mining and Metallurgical Congress Hong Kong (Ed: M J Jones) pp 177-184 (The Institution of Mining and Metallurgy: London). Logan, A S, 1992. 1100 Orebody Rock mechanics review. Internal Technical Report Neindorf, L B, 1983. Fill operating practices at Mount Isa Mines, Third International Symposium on Mining with Backfill, (Ed: Swoboda) Pp 179-187 (Balkema: Rotterdam). Perkins, W G, 1984. Mount Isa silica dolomite and copper orebodies: the result of a synctectonic hydrothermal alteration system, Econ Geol, 79:601-637. Poniewierski, J M. P446 Stope Filenote. Internal Filenote. Potvin, Y, Tyler, D B, MacSporran, G, Robinson, J, Thin, I, Beck, D and Hudyma, M, 1999. Development and implementation of new ground support standards at Mount Isa Mines Limited, in Proceedings of the International Symposium on Ground Support, Kalgoorlie, Western Australia, 15-17 March 1999 (Eds: E Villaescusa, C R Windsor and A G Thompson) pp 367-372 (AA Balkema: Rotterdam, Brookfield). Robinson, J and Tyler, D B, 1999. A study of corrosion in underground reinforcement at Mount Isa Mines, Australia. in Proceedings of the International Symposium on Ground Support, Kalgoorlie, Western Australia, 15-17 March 1999 (Eds: E Villaescusa, C R Windsor and A G Thompson) pp 77-82 (AA Balkema: Rotterdam, Brookfield). Tyler, D B, 1999a. Subsidence on 4340XC, 15 Level. Internal Memorandum Tyler, D B, 1999b. X41 Guide rail deformation – an update. Internal Memorandum Stacey, T R and Page, C H, 1986. Practical handbook for underground rock mechanics. (Trans Tech Publications: Clausthal-Zellerfield). Spearing, A J S, 1995. Handbook on hard-rock strata control, 152 p (The South African Institute of Mining and Metallurgy: Johannesburg).
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George Fisher Mine — Feasibility and Construction L B Neindorf1 and G S B Karunatillake2 ABSTRACT
INTRODUCTION
Following a pre-feasibility study, Mount Isa Mines commenced a feasibility study on the deposit of the George Fisher Mine in May 1997. Board approval was granted in May 1998 to construct a 2.5 million tonne per annum mine and to upgrade existing concentrator and smelter facilities. The world-class George Fisher zinc-lead-silver deposit is located approximately 22 kilometres north of the city of Mount Isa, in north-west Queensland and is 2.6 kilometres north of the existing Hilton Mine. It will be a large-scale underground mine, sustaining a profitable zinc-lead-silver business in Mount Isa for at least ten years. This paper will outline the details of the mine that was developed during the feasibility and design phase. It will include a description of the mining methods, ventilation, fill, ore handling, services and infrastructure. The utilisation of existing assets and the development of innovative ideas were essential to the successful economic development of the George Fisher Mine.
Syd Carter, chief geologist for Mount Isa Mines Limited (MIM), discovered the George Fisher deposit (previously known as Hilton North) in 1948, approximately 22 kilometres north of the city of Mount Isa, in north-west Queensland and is 2.6 kilometres north of the existing Hilton Mine (Figure 1). Recognition of jasper and ironstone gossans overlying the mineralisation led to exploration along the strike and the discovery of the Hilton orebodies. Extensive drilling from the surface commenced in the mid-1980s (see Table 1). In 1991 underground development commenced to access the George Fisher deposit on 12 Level (750 metres below surface) from Hilton. Additional drilling for the pre-feasibility study was completed between December 1993 and December 1996 from underground drill sites (see Table 1). Work on the George Fisher pre-feasibility study was completed in March 1997. The George Fisher feasibility study commenced in May 1997 and was completed in February 1998. Following the completion of the bankable feasibility study in April 1998, MIM announced that it would proceed with the $270 million George Fisher development, comprising a new mine, and
1.
MAusIMM, George Fisher Project , Mount Isa Mines, Hilton Admin Building, Private Mail Bag, Mount Isa Qld 4825.
2.
George Fisher Mine, Mount Isa Mines, Hilton Admin Building, Private Mail Bag, , Mount Isa Qld 4825.
FIG 1 - George Fisher Mine location map.
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FEASIBILITY STUDY SCOPE
TABLE 1 Exploration drilling at George Fisher 1948-1996. Location/Type
Total (metres)
During mid-1995 to 1996 (metres)
Underground
85 900
67 810
Surface diamond
84 400
4500
Surface percussion
10 000
0
Total
180 300
72 300
An essential ingredient in progressing the pre-feasibility to bankable feasibility status was to have an initial well-defined feasibility scope. The mining section for the George Fisher feasibility study was prepared on the basis of the following agreed parameters. Production Rate
- 2.5 Mtpa
Ore Sources
- C and D orebodies 6900N to 7800N between 9 and 14 levels (Figure 2).
Mining Methods upgrades and life extension work to the existing Hilton Mine and lead-zinc processing facilities at Mount Isa in April 1998. Since then an integrated project team has been established, comprising of MIM and Bateman Brown and Root. The mine infrastructure development contract was awarded to Brandrill Ltd and underground development commenced in September 1998. Production was scheduled to commence in the mid second half of the year 2000. The development of some initial trial stopes permitted 250 000 tonnes to be produced in 1999/2000. At full production the George Fisher Mine will hoist 2.5 million tonnes of ore a year, producing 170 000 tonnes of zinc in concentrate and 100 000 tonnes of crude lead containing approximately five million ounces of silver. The projected life of the mine will be over ten years.
- Sublevel open stoping for orebody widths greater than ten metres and benching for orebody widths six to ten metres.
- A cemented fill system with no surplus
Fill
water and having a maximum curing time of 28 days. Ore Handling
- Two possible systems will be evaluated a) Underground trucking of ore to the Hilton crusher and hoisting via the P49 shaft. A dedicated haulage level will be developed on 14 level (Figure 2). b) Crushing and hoisting via new facilities at George Fisher.
GEORGE FISHER PROJECT Longitudinal Section Looking West Showing Hilton Mine* and Key Development
Existing P49 Hoisting Shaft
South
L72 Ventilation Shaft
Existing decline access
K74 Ventilation Shaft
Lim it of proposed development
North
7C
10L 12L 15C
14L
16L Ore Limit
Proposed haulage
0
250 (Metres)
500
GEORGE FISHER
7800 N
HILTON MINE
6900 N
Ore Limit
* Some Hilton Mine development omitted for clarity
FIG 2 - Longitudinal section showing Hilton Mine and key development for George Fisher Mine.
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Servicing
- An efficient men and materials transport
system will be re-engineered using the P49 shaft and existing decline connection from 10 to 12 levels (Figure 2). No decline to the surface will be included. The objective was to put in place a state-of-the-art mine, which will deliver a quality product to the concentrator at minimum cost. Other essential ingredients in completing the bankable quality feasibility study were:
• a dedicated experienced focussed team with minimal work done outside the team;
No adverse groundwater related problems have been experienced underground at George Fisher and no major water inflows have been encountered in development. Limited water inflows have been experienced in diamond drilling where holes intersect the Hangingwall Fault. There is significant exploration potential within the George Fisher deposit. This includes the northern continuation of the deposit up to the Spring Creek Fault and extensions of the same mineralised sequence at depth (below 16 level) and to surface. More regionally there is a considerable economic opportunity at depth below the existing Hilton workings.
• early establishment of an economic model and evaluation of
RESOURCES AND RESERVES
reserves;
• utilisation of independent experts in key areas; • well documented information that supported and/or verified conclusions;
• • • •
emphasis on always adding value to the study; a mine to market approach; regular reality checks and reviews; and risk assessments conducted at different decision points.
GEOLOGY Of the 180 300 metres of exploration drilling completed, 86 000 metres were drilled from underground. All drill-core was geologically sampled and assayed and geotechnically logged. Underground cross-cut and drive development on 12 level was mapped for rock type, ore distribution, structure and geotechnical condition. The deposit has now been interpreted to extend from the surface to over 1000 metres in depth and for 1700 metres along strike. Direct access to the orebodies has enabled confirmation of structural interpretation and metallurgical bulk sampling prior to the feasibility study. Mineralisation at George Fisher is similar to the proterozoic shale style of the zinc-lead-silver deposits at Mount Isa and Hilton (Forrestal, 1990) with orebodies occurring in bedded-parallel tabular orebodies, hosted in well-bedded Urquhart shale. The deposit strikes approximately 18 degrees west of magnetic north. The dips of the orebodies vary and are generally between 40 - 55 degrees. The depth of mining at George Fisher will be between 420 and 1060 metres below the surface. Eleven separate parallel orebodies have been identified at George Fisher (Figure 3). Two central orebodies, C and D, show the best potential in terms of continuity, thickness, grade and metallurgical recovery. The average true widths for C and D orebodies are 15 and 25 metres respectively. Sphalerite-pyritegalena-pyrrhotite type sulphide mineralogy is present in all George Fisher orebodies. Major block faults are prevalent, consisting of steep south-east dipping transverse faults on the southern end and steep east dipping strike orientated faults in the northern end. These faults cut the orebody into blocks ranging in strike length from 30 – 150 metres. The faulting at George Fisher, unlike that which occurs at Hilton, does not constrain the use of lower cost bulk mining methods. Within the ore-blocks minor faults commonly occur which may be splays off major block-bounding structures. The significance of these minor structures on the mine stability is not known. The Hangingwall Fault marks the western boundary of the deposit. Exposures in the underground development indicated that this fault may be up to 40 metres wide and heavily sheared.
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The combined Proved and Probable Ore Reserve of 24.2 million tonnes at 9.1 per cent zinc, 5.6 per cent lead and 128 grams per tonne silver was estimated for George Fisher from a comprehensive geological block modelling exercise. Although 11 orebodies have been defined, the reserves only include ore from C and D orebodies. Grade distribution varies within and between orebodies. Currently orebodies are grouped according to metal grade, pyrite/pyrrhotite distribution, gangue mineralogy and metallurgical assessment. C and D orebodies in general contain the highest value material. They have higher lead and silver credits and a lower pyrite content (eight to ten per cent). Tabulated in Table 2 are the current Ore Reserves and Identified Mineral Resources derived from the geological block model (published in the MIM 1999 Annual Report to Shareholders). The information contained in the table has been compiled in compliance with the ‘Australasian Code for Reporting of Identified Mineral Resources and Ore Reserves’, July 1996, for metals.
MINE PRODUCTION AND METHOD Integration The George Fisher development, timing, production build up and rate needed to be investigated in view of the existing lead-zinc mining and processing facilities. The potential impact of production profiles on the George Fisher feasibility project required investigation of the following key questions:
• Will lead smelter capacity be met and/or exceeded with production profiles?
• Do the profiles impact on the proposed George Fisher project?
• Is it possible to defer the expenditure of capital for the George Fisher project?
• If so, are there potential economic gains in deferring the capital expenditure for George Fisher? An economic model reflecting expected improvements in recoveries due to the implementation of planned concentrator upgrades for the treatment of George Fisher ore was essential in optimising the integrated production profile (Gregor et al, 1998). This was completed for the remaining reserve in the Isa Lead and Hilton Mines and the George Fisher Mine.
Mining method A detailed mining system that satisfies the following objectives was selected for the extraction of C and D orebodies:
• provide a safe and quality working environment for the workforce;
• maximise the head grade and ore recovery;
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• achieve a production rate of 2.5 million tonnes per year; • maximise the existing capital investment at the Hilton mine; • achieve a mining cost that will place the mine in the lower half of the cost curve; and
• maintain the integrity of the surrounding rock-mass so that other orebodies can be successfully extracted at a later stage. Geotechnical information collected from drill core logging, underground mapping and observation was utilised to establish stope parameters. Stope hangingwall stability was assessed by using both the Hangingwall Stability Rating (HSR) and the Modified Stability Graph (MSG) methods (Potvin, 1988). Panel Open Stoping in the wider parts (generally greater than ten metres) has been chosen as the mining method for the mining of C and D orebodies. It will be based on a 30 metre sub-level
interval with 15 metre long primary and 20 metre long secondary stopes. Maximum height of C and D orebody panel stopes will be limited to two lifts (60 metres). In the primary/secondary stoping sequence proposed for panel stoping, it is essential that the primary stopes lead the secondary stopes by at least one sublevel. This is required to avoid the formation of a large continuous sill at any sublevel, which will be difficult, if not impractical to support in wider sections of the orebodies. For this reason, it is planned that the initial primary stopes above 13 level will have a height of 60 metres (two lifts), while the secondary stopes will have a height of 30 metres (one lift). Thereafter, both primary and secondary stopes will always be leading the secondary stopes by at least 30 metres (Figure 4). Bench Stoping (Cunningham, Neindorf, and Villaescusa, 1994) will be used in orebodies up to ten metres maximum width with some variation in application in the six - ten metre range.
FIG 3 - Typical geological cross-section.
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TABLE 2 Ore reserves and identified mineral resources. Tonnes (Mt)
Zn (%)
Pb (%)
Ag (g/t)
C and D Orebodies only Ore Reserves Proved
16.2
9.1
5.8
136
Probable
8.0
9.0
5.1
116
Total Reserves
24.2
9.1
5.6
129
Measured
2.1
10.1
6.2
119
Indicated
2.0
10.0
6.3
114
Inferred
20
11.9
5.7
73
24.1
11.6
5.8
80
69
Identified Mineral Resources
Total Resource (C and D)
Other Orebodies Identified Mineral Resources Measured
12.3
11.3
4.0
Indicated
7.6
11.4
4.2
70
Inferred
33
12.3
5.5
77
52.9
11.9
5.0
74
Total Resource (Other Orebodies)
FIG 4 - Longitudinal section looking west showing stoping sequence.
Benching is the method used for narrower orebodies and is widely used at Mount Isa and Hilton. A high recovery can be achieved without using cemented fill. Variation in ground conditions can be handled by the timing of filling, with delayed filling in good ground and immediate filling in less stable areas. Benching is not applicable to wider orebodies as the process automatically sills out the orebody to full width at the lower extraction level. All major infrastructure will be located in the footwall of D orebody. Stopes will be accessed from footwall drives developed off the declines (Figure 5 shows a plan of a typical level arrangement). Access to ore passes, return air raises and fresh air raises is from a footwall drive driven parallel to D orebody on each sublevel.
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Drilling and blasting A cut-off slot and main ring drilling and blasting technique will be used for extracting these stopes. This involves establishing a single central access with a vertical cut off slot with a footwall orientated cut-off raise. Fanned rings parallel to the slot will be drilled from the developed hangingwall drive. A trial panel open stope is planned with 89 mm blast holes. Figure 6 shows the development and drilling arrangement of the first 714D trial stope. Continuous bench stoping in C and D orebodies will also be trialed. Bench mining will be used where the orebody width is less than ten metres.
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FIG 5 - 13C sublevel plan.
FIG 6 - 714D trial stope.
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Trial stoping permitted earlier and the ability to program, which methods.
using the two main mining methods has production from the George Fisher orebodies obtain valuable information from a monitoring will permit analyses and refinement of the
Sequence and scheduling Initial primary stopes in D orebody are mined using a one, five, nine sequence followed by mining primaries in three and seven, followed by mining secondaries in two, four, six and eight (Figure 4). Following the extraction of D orebody, C orebody will be extracted on a particular horizon in a stressed shadowed environment. An NFOLD (the collective name for a suite of computer programs which permit the modelling of ground behaviour following an excavation) stress analysis was carried out to investigate stress levels around likely stoping sequences. This showed that high stresses developed at the base of secondary stopes if they lagged the primary stopes by an excessive amount. However it was assessed that these high stress levels would dissipate readily by small-scale movements on the numerous transverse faults. For this reason, it was concluded that regional pillars are not required for stress management purposes. However, it is planned that two 35-metre long temporary regional pillars will be left between 7077.5 N and 7112.5 N and between 7567.5 N and 7602.5 N as major accesses to C orebody. These pillars would be recovered progressively in the later stages of mining the deposit. Stress monitoring will be conducted as mining progresses as a means of refining planning assumptions. The production rate of 2.5 million tonnes per annum is achieved after a build-up over three years. The time between the completion of one stope and the commencement of the next scheduled stope in the sequence is critical for the production rate in both C and D orebody.
MINING INFRASTRUCTURE During the feasibility study it became clear that optimising the use of some of the existing assets added greater value than building new infrastructure. These assets included using existing or upgraded mine access, ore handling and hoisting facilities, main pump station and main ventilation systems.
Mine access and services infrastructure The main access to George Fisher will be via the existing eight metre diameter P49 Shaft at Hilton Mine and the LE52 decline from 10 level to 12 level. This shaft is concrete lined from the surface to the sink below 16 level, (1040 metres). The shaft is currently used as a service facility for men and material, and for rock hoisting. During the initial two year mine development phase, over eight kilometres of horizontal development will be completed. Figure 5 shows the typical development being mined during construction, including ore passes, return air and fresh air rises. In additional to this mobile equipment service bays, cribroom, fill station, pump station and fuel bay are being built during construction.
Ventilation
Miscellaneous mine infrastructure In addition to the above infrastructure a number of new facilities have been installed at George Fisher. This includes an underground fully air conditioned integrated office and lunchroom facility. This area caters for mine planning, tele-remote loader operation, supervision, training and emergency refuge. Located in the same complex is the mine communication system centre, which primarily collects and distributes data and information throughout the mine via an optic fibre and copper cable network. The communication system supports voice and data in the form of telephone, radio, LAN (Local Area Network) and SCADA (Supervisory Control and Data Acquisition). A PED (Personal Emergency Device or Productivity Enhancement Device) system has also been employed to further improve communication links and initiate mine blasts. An innovative mine management system is also being developed to best utilise real time information for improved decision making. Another facility that is being constructed is a mine refuelling system that supplies diesel into the mine via a dedicated fuel hole directly from a surface fuel farm. This facility will reduce the logistical issues of getting diesel into the mine via the P49 shaft. A mobile equipment service facility is also being constructed at George Fisher. All major workshop work will be completed in the existing Hilton workshop facilities. A new explosive magazine and tyre bay is being constructed at George Fisher. The existing P49 pump station at 10 level has also been upgraded with a positive displacement system that can pump dirty water directly to the surface. A new positive displacement system is being installed on the 15C sublevel haulage horizon at George Fisher. This will pump dirty water to the 10 level pump station.
Mullock and ore handling
The primary ventilation system at George Fisher is developed around two shafts; the seven metre diameter intake shaft L72 (570 m3/s), and the 7.6 metre diameter exhaust shaft K74 (700 m3/s), approximately 200 metres to the north of L72 shaft. In addition to these shafts, existing Hilton shafts will provide additional intake and exhaust.
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The primary ventilation system is developed in three stages. The initial stage where the existing J53 shaft provides exhaust and fresh air down the existing P49 shaft via the LE52 decline from 10 level and 14 level for the truck haulage drive and the L72 fresh air shaft. Secondly following the completion of the K74 shaft exhaust fan power is progressively installed to meet requirements. The K74 shaft was initially raisebored to 2.4 metres in diameter to 12B sublevel and is then stripped to 7.6 metres in diameter using the horidiam shaft development technique. This proposal includes transferring some of the existing installed plant from J53 to K74. The timing of relocating the fans will be well coordinated with Hilton Mine to minimise interruption to production requirements. Finally the 3.1 metre L72 shaft is stripped to seven metres in diameter again using the horidiam shaft development technique. This will now provide an independent fresh air supply to the George Fisher Mine. L72 and K74 shafts will be connected to a series of fresh air and return air raises (3.1 metre diameter) with dedicated intake and exhaust horizons developed respectively on 12C and 12B sublevels. With these raises, three independent ventilation domains will be established of the footwall drive on each sublevel. This ventilation flexibility is essential to enable the various mining activities to be performed while building up to and maintaining the final 2.5 million tonne production output.
During construction the amount of mullock produced exceeded the capacity to dispose of mullock in the Hilton mining operation. To cater for this, a mullock handling system including a raise-bored mullock pass (NB 49 10B on 12 level) was constructed to allow mullock to be disposed of through the existing Hilton ore and mullock handling system.
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FIG 7 - George Fisher process flow diagram.
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Large (7 m3 - 17 tonne bucket capacity) load haul dump units will transfer ore from stope draw-point to the footwall ore passes. Three of the raisebored (2.1 metre diameter) ore passes are being developed now as a part of the initial project infrastructure. Later, these ore passes will be extended to transfer ore above 12 level to 15C sublevel. Fifty tonne diesel articulated dump trucks will haul ore from the ore pass chutes on 15C sublevel to 14 level Hilton, via the 2.5 kilometre truck haulage drive and the developed truck loop system. Ore will be tipped from 14 level through a new grizzley/rock breaker arrangement to a new ore pass and feeder system. This provides a third ore pass and a second feeder to feed the existing LE49 jaw crusher on 16 level. The ore is then conveyed to an upgraded loading station and hoisted up the P49 hoisting system (also upgraded). The hoisting capacity was estimated as 1.9 million tonnes per annum with the existing two 12.5 tonne capacity skips. The annual hoisting capacity has been upgraded to 2.5 million tonnes with the payloads of the skips being increased to 14 tonnes. Modification to the existing crushing, conveying and hoisting system is in progress and will be used to handle George Fisher ore. On the surface ore will be transported by diesel on off-highway trucks to the metallurgical plant in Mount Isa. The same trucks will be used to back-haul aggregate backfill material from Mount Isa to the George Fisher mine.
MINE FILL The chosen fill system for filling the initial primary open stopes is Cemented Slurry Aggregate Fill (five per cent Portland Cement) using Heavy Media Reject (HMR) as aggregate. The HMR consists of 17 mm free flowing product and will be delivered using the same trucks that are used to deliver ore from the P49 truck loading facility. The HMR is unloaded into a surface bunker and flows underground via a 740 metre vertical fill pass from the surface. A surface slurry fill plant will be constructed adjacent to the aggregate fill hole. A Portland cement slurry will be mixed on the surface and placed down a dedicated 100 millimetre diameter cement slurry fill hole. The Portland cement will be transported by road to the mixing station in sealed bulk road transporters and pneumatically transferred into storage silos. At the underground mixing station a 50 tonne ejector truck will proceed under an aggregate filling point with the driver initiating the mixing cycle. This will start a timer, which will open the fill pass discharge chute and start a slurry transfer pump. The timer will be set to give a controlled amount of aggregate and cement slurry to the truck. First of the five fill mixing stations will be commissioned by September 2000 on 12 level. Articulated trucks with ejector trays will deliver the fill to the stopes. Non-cemented development mullock supplemented with HMR will be used in secondary stopes. The HMR is supplied at the same underground mixing facility via a diverter. The overall George Fisher mining process is summarised in Figure 7.
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MINE CONSTRUCTION A preliminary engineering study was completed to assist in firming up on the mine construction scope, budget and schedules. This study established the necessary controls to manage the construction to be brought in on budget and on schedule. One of the main risks identified during the feasibility risk assessment was the brown field nature of the project construction and the affect on the existing Hilton production during construction. The risk has been minimised through the good communication between operations and construction. Designs have been continually improved through regular formal and informal communication between the construction team (the supplier) and the operation team (the customer). At all times the feasibility and engineering study has been an excellent reality check. As the construction activities are completed and handed over to operations, there is a stronger ownership and better understanding of the operational and maintenance requirements of the mine.
CONCLUSIONS The George Fisher Mine will be a world class underground zinc-lead-silver mine. It has been developed on the basis of utilising the best of the existing Hilton infrastructure and introducing new or refurbishing existing infrastructure. New and innovative systems have been introduced to support best operating practices. Theses systems have been introduced on a brown field site using the feasibility study as a basis and a consultative process with the operators to continue to improve existing operations. This has created an environment that has contributed to improving the existing Hilton operation and provides a positive outlook to the future of the George Fisher Mine.
ACKNOWLEDGEMENTS The authors wish to thank the management of Mount Isa Mines Limited for the permission to publish this paper.
REFERENCES Forrestal, P J, 1990. Mount Isa and Hilton Silver-Lead-Zinc Deposits, in Geology of Australia and Papua New Guinea (F E Hughes) pp 927-934 (The Australasian Institute of Mining and Metallurgy: Melbourne). Pease, J D, Young, M F, Rosengren, M D and Gregor, G, 1998. Finding a Mine Within a Mine, The Mount Isa Lead/Zinc Case Study, in Proceedings 1998 AusIMM Annual Conference, pp 327-331 (The Australasian Institute of Mining and Metallurgy: Melbourne). Potvin, Y, 1988. Empirical Open Stope Design in Canada, PhD Thesis, (The University of British Columbia, Canada). Cunningham, J B, Neindorf, L B and Villaescusa, E, 1994. Bench Stoping of Lead/Zinc Orebodies at Mount Isa Mines Limited, in Joint Japan and Australasian Institute of Mining and Metallurgy Symposium, (Uke, Japan).
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The Ventilation and Refrigeration Design for Australia’s Deepest and Hottest Underground Operation — the Enterprise Mine R Brake1 and B Fulker2 • Prevention: removing or reducing the risks from heat
INTRODUCTION The Enterprise Mine (EM) is located at Mount Isa and is wholly owned by Mount Isa Mines Limited (MIM). This high-grade copper mine extends from about 1000 m below surface to almost 2000 m below surface. A recent $400 million upgrade has taken the operation from a production rate of 1.3 Mtpa to a capacity of 3.5 Mtpa. The effects of high surface ambient temperatures in summer, combined with the depth of the mine and high virgin rock temperatures results in heat stress in the working place that, without intervention, would exceed the levels that human physiology can withstand. It would also result in significant decreases in productivity and high accident rates even where work is possible. Moreover, with a very hot and deep mine, the needs of ‘egress’ and entrapment of workers during a mine emergency must be carefully considered. Enterprise Mine has taken a ‘systems approach’ to the issue of working productively, healthily and safely in heat. This involves setting limits of thermal stress for workers, understanding productivity decrements when working in heat, establishing productive and healthy environmental targets, developing engineering solutions to meet these targets, creating a well-educated workforce with respect to working in heat and then developing health protocols to ensure workers’ health is protected. Refrigeration must be added to the mine to provide acceptable and productive conditions underground. By the time the mine is commissioned in 2000, Enterprise mine will have an installed refrigeration capacity of in excess of 40 MW(R). The ventilation engineering principles proven over many years at the other Isa underground mines are tailored to ‘flooding’ the mine with air to remove the heat, which is the main contaminant. This principle will not work at the Enterprise Mine, for reasons discussed below. With the underlying ventilation principle no longer satisfactory, a complete review of the entire ventilation principles and engineering solutions was required for the Enterprise Mine.
hazards;
• Monitoring: monitoring the level of risk created by the hazards; and
• Contingency: developing contingency plans that minimise harm if the risk from the hazard rises to a dangerous level. In other words, if Plan A fails, what is Plan B? These are shown diagrammatically in Figure 1. Equipment/ Engineering
Procedures
Competency
Prevent Monitor Contingency
FIG 1 - The elements of a system to control risk.
Enterprise Mine then developed a series of statements (Figure 2) that embodies this philosophy in terms that management and the workforce could relate to.
Engineering the Environment We choose airflows, the amount of refrigeration and other engineering solutions to bring the heat to acceptable levels
Job Design We use air-conditioned cabins, air movers, labor-saving devices and other ways to improve the local environment or reduce the physical effort required for work
Ventilation and Refrigeration Standards HAZARDS, RISKS AND A SYSTEM TO MANAGE THE RISKS The control of hazards related to in working in heat were reviewed under three categories:
• Engineering: what engineering solutions can be put in place to reduce and maintain the risks at acceptable levels?
• Procedures: what procedures are required to ensure the risks can be managed at acceptable levels?
• Competency: what level of education, training and competence is required by all persons in the system to ensure the risks are sustained at an acceptably low level? Hazard management was reviewed in three respects:
We develop standards for the way we install and operate equipment, or go about our work, so that our engineering solutions will be effective
Health and Safety Medical Protocols We provide health procedures and medical tests to ensure that heat illness is avoided or picked up at a very early stage and treated
Education We educate our workforce so they understand what happens when they work in heat and can do so safely and without damaging their health
Emergency Response 1.
FAusIMM, CPMin, MMICA, Principal, Mine Ventilation Australia, 12 Flinders Parade, Sandgate Qld 4017, formerly Project Manager-Ventilation and Refrigeration, Enterprise Mine Project.
2.
MAusIMM, Enterprise Mine Manager, Mount Isa Mines Limited.
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We provide a means of ensuring that anyone who develops a heat illness can get effective, immediate treatment. FIG 2 - The EM philosophy on working in heat.
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All the elements of this system are required to be in place and working effectively if the system as a whole is to operate properly (see Figure 3).
QUALITY OF AIR REQUIRED Local statutory regulations, codes of practice and other standards (such as Australian and international standards) can either set mandatory requirements or provide technical guidance as to the quality of air required in the mine. In general, because people must be able to traverse most areas of the mine (to fix equipment, to inspect ground conditions, to muck out shaft spillage), the relevant limits are those that apply to people rather than equipment. However, where the requirements for equipment are more onerous, then these would apply. Statutory and codified requirements are usually of two types:
• maximum concentrations of harmful substances, eg carbon monoxide, aldehydes, dust, NO and NO2 etc; and
• minimum airflow requirements, either in terms of volume (m3/sec) or velocity (m/s).
By estimating the production of harmful substances in the mining process, the required dilution rates with fresh air can be calculated in a relatively straightforward manner, although care needs to be taken; for example, not every item of diesel equipment in the mine is operating all the time, nor is it operating at full engine power. Obviously, it is rare for the design quality of air to change over the life of the mine.
THE PARTICULAR CASE OF HEAT
FIG 3 - How the system fits together.
Heat is a particular case of an unwanted air-borne contaminant. There are three almost ‘unique’ features of heat:
THE ENGINEERING REQUIREMENTS OF THE SYSTEM Certain minimum amounts (volumes) of air are required to dilute noxious and nuisance gases and particulates to acceptable levels. A major pollutant in the underground air in the Enterprise Mine is heat. In general, it is the heat in the mine workplace at EM that is the determining factor in the amount of air and refrigeration required. The management of heat in the Enterprise Mine is expected to be the single most significant success factor in the on-going operation of the mine, although there are other innovations being introduced to the operation, especially in the area of paste fill. Special protocols have been developed to ensure all levels of heat can be handled safely and without damage to health (Brake, Donoghue and Bates, 1998). In the final analysis, there are three key issues to be resolved:
• What quality of air is required in each area of the mine, over time?
• How much (quantity of) air is required in each area of the mine, over time?
• What is the ‘best’ way of delivering the required quantity and quality of air to each area of the mine, over time? A key concept is that optimising the ventilation cannot be done simply for a single ‘snap shot’ of the mine’s life: it must be done ‘over time’. The ventilation consequences of the normal ‘ebb and flow’ of a mine operation over its life, from construction through to remnant mining, must be considered. Ventilation design is an iterative process even when the mine is not considering refrigeration – it requires optimising airflows and airway sizes and fan duties. In a refrigerated mine, there are several more design parameters to be optimised. It is therefore essential to treat the engineering of the environment as a system, and to use a systems approach. Taking a narrow view means that one problem may be optimised, but at the expense of producing an unsatisfactory or suboptimal system as a whole.
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• With other contaminants, usually the only practical control mechanisms are to dilute the air so that the contaminant is reduced to a safe level, or to prevent the contaminant entering the air. With heat, it is also possible to ‘re-condition’ the air, ie remove some of the heat by refrigerative air-conditioning. It is also possible to make the cooling impact of the air more effective, at least for humans, by increasing its velocity over the human skin, ie without removing any heat at all. These factors are not true of other contaminants.
• With most airborne contaminants, use of a respirator can ensure the inspired air is safe for the lungs; however, a respirator will not remove ‘heat’.
• Higher levels of heat result in significant decrements in productivity, in addition to impacts on health and safety. Generally, there is a critical threshold below which health and safety is not affected, although productivity is. This is because there is a crucial nexus or ‘link’ between heat stress and health and safety problems. This link is the state of hydration, acclimatisation and health of the individual. Workers can work safely in hotter conditions, up to limiting values, providing more regular and longer breaks are taken and providing they are healthy, acclimatised, well hydrated and self-paced. Work can be done safely, but it can be very unproductive, with very high rest components in the work/rest cycle.
MEASUREMENT OF HEAT In terms of the impact on humans, many indices have been developed to measure heat stress. However, most of these require measurement of some or all of the following parameters:
• Dry bulb (DB): is the temperature in common use and referred to in weather reports.
• Meant radiant temperature (MRT): is the heat load due to radiation from a hotter body (eg the Sun) to a cooler body (eg a person). In an underground mine, radiant loads come from rock surfaces (especially fresh rock), diesel engines and the
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THE VENTILATION AND REFRIGERATION DESIGN FOR THE ENTERPRISE MINE
dry bulb temperature of the air (which, itself, is a ‘hot body’). It is usually measured with a thermometer inside a blackened copper globe, with suitable corrections made for DB and wind speed over the globe.
• Wet Bulb (WB): is the temperature at which water evaporates into the air (at a particular dry bulb temperature) once equilibrium between the water and air has occurred. It is very much more important than the dry bulb temperature to physiologists as the evaporation of sweat is related to the partial pressure of water vapour in the air (in effect, the humidity). Knowing any two of dry bulb temperature, wet bulb temperature or humidity (along with barometric pressure) will allow calculation of the third.
• Wind speed: as discussed earlier, the wind speed over the skin is a key issue in the rate of evaporation of sweat and therefore of body cooling. In an underground mine, it is common to assume the radiant temperature is the same as the dry bulb temperature. This is not always true, especially for operators of hot diesel equipment if not in air-conditioned cabins; however, equipment to measure radiant loads was not, in the past, suited to the underground environment. With new technology, better heat stress measurement devices are now becoming available (Bates and Matthew, 1996).
SOURCES OF HEAT Heat is both an input to and a ‘by-product’ of mining. Heat is produced or derived from a number of sources in any mine (Pickering and Tuck, 1997). These include:
• autocompression∗ • machinery: plant and equipment, • geothermal heat from strata (ie influenced by virgin rock temperature (VRT) and other factors),⊕
• • • • •
fissure water, oxidation,
FIG 4 - Surface climatic data for Mount Isa.
well-established procedures (Hemp, 1989). The approximate proportions of heat from the various sources at the EM are shown in Figure 5. The size of the ‘pie’ is about 40 MWr. Note that these values exclude the impact of the Isa climate, which is very substantial.
Heat flow from rock walls 24%
Diesel equipment Electric 7% equipment 7% Broken rock 8% Groundwater 3%
explosives, broken rock, and miscellaneous sources, such as lighting, personnel, service water.
A further, and major, issue on heat is the heat contained in the air even before it enters the mine. This is a function of the surface climate. The Mount Isa climate is shown in Figure 4. At Mount Isa, the maximum (or limiting) summer surface condition is 26.5° WB†. The reference surface condition (exceeded about 220 hours or 2.5 per cent per year) is 25°, the average summer (December to March inclusive) condition is 21° WB. Heat loads can be estimated with reasonable accuracy using ∗ Autocompression is a term which denotes the fact that the air going underground heats up merely by virtue of the conversion of the ‘potential’ energy of the air at surface into heat when the air moves closer to the centre of the earth. At EM, autocompression will increase the wet and dry bulb temperatures by about 6.5° and 8° respectively at the average depth of the mine, which adds about 22 MWr to the required refrigeration loads. ⊕ At EM, the geothermal gradient is 20°/1000 m from a surface rock temperature of 28°. † In future, wet bulb/dry bulb temperature combinations will be written as 25/35: the wet bulb always preceding the dry bulb. All conditions are at 97.5 kPA. Note that the wet bulb temperature can never exceed the dry bulb temperature, so there is no such thing as 35 WB/25 DB
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Autocompression 51% FIG 5 - Heat loads at EM (excl surface climate).
At the summer reference condition of 25/35, and assuming a ‘design’ temperature at the workplace of 28/36, the surface air contributes 13 MW of cooling (ie the surface air is still cooler than the required air temperature in the workplace). At the mid-winter temperature of 10/15, the surface air contributes 61 MW of cooling, totally changing both the size of the ‘pie’ and the relative proportions of each component. As expected, with climatic heat load swings of this size, the surface bulk air coolers can be turned off completely for about 20 per cent of the year.
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Heat is also a special case when it comes to optimising airflows. This is because of an effect known as the critical depth. Critical depth is the depth below surface at which air at the underground workplace will exceed the maximum (design) condition solely through autocompression without any other heat gains. Critical depth is a function of surface temperatures and the depth of the mine only. Critical depth is therefore deeper in winter (because surface air is colder) and shallower in summer. The point about critical depth is that if the workings are above the critical depth, then the cheapest control measure on heat is usually to ‘flood’ the workings with air. However, if the workings are below the critical depth, then the air itself adds to the heat load; in this latter case, the primary airflow should be kept to the minimum practical and the heat controlled via refrigeration. It is normal to design the refrigeration requirements for a notional ‘average’ workplace at the ‘average’ depth of the operation. Local refrigeration arrangements are then used to ensure that the deeper (and hotter) areas are not under-cooled. The average depth of the Enterprise mine in its first six years is 26A sub or 1350 m below surface. The critical depth at a surface temperature of 25° C WB is 1150 m below surface or 23C sub.
HUMAN PHYSIOLOGY AND HEAT Heat illness in underground, Australian miners has been recently described (Donoghue, Sinclair and Bates, 2000). This paper is the first to document the clinical features, haematological and biochemical profiles and thermal stress on these workers. The ability of the body to dissipate heat is a function of many factors. The key ones are:
• Cardiovascular fitness: a strong, fit, heart (the pump, including provision of essential oxygen from the lungs to operate the pump) and good circulation are essential.
• Obesity: this effectively insulates the body and lowers the surface area/body mass ratio, reducing heat loss.
• Acclimatisation: this is the sum total of all the adaptations the body has to extended exposure to heat.
• Hydration levels: the body has to remain well-hydrated when working in heat. Even small degrees of dehydration cause significant decrements in the ability to work in heat.
• Clothing and Personal Protective Equipment (PPE) requirements: the more clothing and PPE that is worn, the greater the insulation effect, the lower the water vapour permeation through the clothing and PPE, and the more difficult it is to stay cool. In an underground mine, there is no direct exposure to radiant heat (except at newly blasted rock faces), so there is nothing to be gained from using clothing to reduce radiant heat loads.
• Environmental conditions: evaporation is affected by several environmental factors. However, the two most significant in an underground mine are humidity and air speed over the skin. Obviously, at zero humidity it is very easy for water (sweat) to evaporate from the skin into the air. When the air is saturated (100 per cent humidity), it is much more difficult for evaporation to occur‡ At intermediate points, the difficulty in evaporation increases as humidity increases, which is why ‘hot, humid’ air is much more stressful than ‘hot, dry’ air at the same temperature. Air speed over the skin improves the rate of evaporation under all circumstances (except at 100 per cent humidity or very high wet bulb temperatures) and therefore facilitates work in heat. Of equal significance is the amount of heat that the body needs to reject. Clearly, this is primarily a function of the work rate. The maximum efficiency of the human body is about 20 per cent, but it is rare to be this high. The harder the body works, the greater is the metabolic rate. To ensure conservative designs, it is
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usual to ignore any ‘real’ (ie useful mechanical work done) and assume the entire metabolic rate (energy expenditure) must be released as heat. The relationship between Thermal Work Limit (Brake and Bates, in press), (called Air Cooling Power (ACP) at Enterprise Mine), and air velocity for the higher wet bulb temperatures typical of thermally stressful situations underground is shown in Figure 6. The units of TWL are watts per square metres of body surface. A 70 kg person has a surface area of about 1.8 m2, so that a metabolic rate (energy expenditure) of 200 W/m2 is equivalent to 360 W. Note the significant increase in TWL up to about 0.5 to 0.7 m/s. Beyond this point, although TWL continues to increase, it does so almost linearly rather than exponentially. Therefore most of the benefit of higher airflows is harnessed by a minimum velocity of about 0.5 to 0.7 m/s. Note also from Figure 6 that if an average cooling power of the underground workplace was set at 180 W/m2 (say), which would ‘cover’ work up to a metabolic rate of 180 W/m2, then this could only be achieved by WB temperatures of about 27 to 28 degrees. Higher temperatures than this become asymptotic to the required cooling rate, or require unrealistic or impractical wind speeds. From inspection of Figure 6, it can be seen that whether the mine is above or below the critical depth, and whether it is relying on refrigeration or ‘flooding’ as the heat control mechanism, there is a lower limit on airflow that must be achieved. However, this can be achieved by:
• stronger primary ventilation, ie more fresh air; or • less primary ventilation but augmented by increased velocity of air over the skin, eg using an ‘airmover’. Using an airmover to increase the velocity of air over the body does not constitute ‘recirculation’ of air, which is effectively banned in most jurisdictions. Because TWL is measured in W/m2, it can easily be compared to watts of refrigeration ‘coolth’. The impact of localised cooling using refrigeration or an airmover can therefore be measured directly. For example, consider a workplace being ventilated with 10 m3/s of air at 30° WB, 40° DB, 40° Globe, 100 kPa barometric pressure and a wind speed of 0.2 m/s. The initial TWL is 105 W/m2. Local refrigeration of 100 kWr [kilowatts of refrigeration effect¥ is installed. Standard psychrometric equations can be used to calculate that temperatures will drop to 28. 0° WB and 31.4° DB, which results in an increase in TWL to 150 W/m2. The capital and operating costs of this engineering intervention (refrigeration) can be directly evaluated against the cost benefit of improved productivity Using the above example, the TWL could have been increased to the same 150 W/m2 by increasing the wind speed over the skin to 0.8 m/s without any addition of refrigeration This is achieved almost cost-free. On the other hand, the introduction of refrigerated air has had a noticeable effect, as predicted, on required air volumes at the working place. Typical comments such as ‘we can now get away ‡ Evaporation is still possible while the skin temperature exceeds the dew point temperature. However, once the dew point temperature exceeds skin temperature, then moisture will condense from the air onto the skin, ie, the reverse of evaporation, and the latent heat of condensation will be added to the skin, resulting in rapid hyperthermia. § An ‘airmover’ is a compressed air-operated, venturi device with no moving parts. ¥ The unit of refrigeration output is a megawatt of ‘coolth’ or refrigeration, written MW(R). The amount of electrical power required to generate a MW ( R) varies according to several factors. As a general rule of thumb, for a large, industrial, purpose-designed plant, an overall ‘coefficient of performance’ of about 3 is typical and means that a 20 MW ( R) plant requires about 6.7 MW (E) (megawatts of electrical energy) to operate.
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[WB from 26 to 32 deg, DB=WB+4 deg, Trad=DB] 300
250
TWL
200
150
100
50
0 0.0
0.2
0.4
0.6
26
0.8
27
1.0
1.2
Air Velocity (m/s)
28
29
30
1.4
1.6
31
1.8
2.0
2.2
32
FIG 6 - Thermal work limit versus air velocity.
with half the air we used to require’ are borne out by Figure 7, which shows the volume of air to a typical 40 m2 cross-section underground workshop going from 50 m3/s before refrigeration, to half this amount by only reducing the wet bulb temperature a little over 1°.
[DB=WB+6 deg, TWL fixed at 180 W/m2 ]
30
WB
29
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28 27 26 0
10
20
30
40
50
60
Air volume (m3 /s)
FIG 7 - Reduction in airflow required in underground workshop after introduction of refigerated air.
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When humans work in heat, their deep body core temperature will not rise to dangerous levels while the cooling power of the environment is greater than the metabolic heat generated as they work. Once the cooling power is exceeded by their metabolic rate, the heat can only be stored in the body, which can quickly rise to dangerous internal ‘core’ temperatures. Productivity is also seriously affected by excessive thermal stress. For example, if the metabolic rate (work rate) for a particular type of work in an environment of low thermal stress is 180 W.m-2, and assuming a ‘resting’ metabolic rate of 60 W.m-2, then the productivity when working in an environment with a TWL of 120 W.m-2 is given by: Productivity = (120 - 60) / (180 - 60) = 50 per cent, where (120 - 60) is the actual residual work capacity (the working rate less the resting rate) in this environment and (180 - 60) is the required residual work rate for full productivity. This then allows simple calculations of the cost of lost production and other economic impacts of environmental conditions
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A sound upper design criteria, in terms of WB, for a typical workplace is considered to be 28° WB, as this should allow a metabolic rate of at least 180 W/m2 with realistic airflows. This metabolic rate is above the average work rate in both metal and non-metal mines (Van Rensburg et al, 1991; Tranter, 1998). Good practice in terms of the ‘spread’ of temperatures around this design maximum is considered to be 2° C WB (refer to Figure 8).
METHODS OF CONTROL OF HEAT IN UNDERGROUND MINES Methods of control of heat in mines is a substantial issue. As mentioned previously, there are two major control mechanisms:
• higher airflows: either by way of more primary airflow through the mine, or by achieving higher local airflows using airmovers or other such devices, or both; and
• refrigeration. Many types of refrigeration have been proposed and trialed in underground mines. The relative advantage of these depends on such factors as: depth of the mine, local power costs and mining method. However, at EM, along with many other major new operations today, refrigeration is achieved by a mix of the following:
Reference Summer Conditions Avg temp = 28 WB, Std Dev = 2 WB
25%
• Surface Bulk Air Cooling. Enterprise Mine uses an 8 MWr
Probability
20% 15% 10% 5% 0% 20 22 24 26 28 30 32 34 36 Wet Bulb
FIG 8 - Temperature distribution at underground workplace.
With the estimate of heat loads underground, the climatic profile and a ‘typical’ spread of underground temperatures, the average and actual distribution of temperatures in the workplaces can be established for any given refrigeration scenario. By then putting a cost on the lost productivity (direct wages plus production [revenue]) and a cost (capital and operating) on provision of more refrigeration, an ‘optimum’ or design maximum wet bulb temperature at the working place underground can be established. For the EM operation (and most other mines), the optimum (maximum design) temperature at the work place is 28° WB @ 0.5 m/s airflow (170 W/m2 TWL at 36° DB). Hence 66 per cent of jobs should lie between 26 and 30° C WB; 95 per cent of jobs should lie between 24 and 32° C and 99 per cent of jobs should lie between 22 and 34° C From Figure 8, if the average workplace temperature is reduced, the percentage of shifts spent working in more extreme conditions (WB>30°) is also reduced, but there are still some. But reducing the average requires much higher refrigeration capacity and operating costs. If the spread (standard deviation) is reduced, with the average left unchanged, the percentage of hot shifts is reduced and productivity improved without any extra capital or operating cost. Note that even if there is a ‘normal’ distribution of temperatures within the workplaces, the burden of ‘hot’ shifts through the mine will not fall ‘evenly’ but will be biased against the jobs that are deeper (more autocompression, higher VRTs, etc) and/or more difficult to ventilate (forced versus flowthrough, etc).
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existing surface refrigeration plant generating cold water at about 3°C, in combination with a new 24 MWr plant generating cold water at 7°C, to feed separate cooling towers at the collar of two major intake shafts. Surface cooling has the advantage of easy maintenance on the plants, keeping the ammonia refrigerant out of the mine, lower condensing temperatures than underground, ‘free’ refrigeration due to the difference between low wet bulb temperatures on surface and the condensing temperature, and high plant Coefficient of Performance+ .
• Underground Bulk Air Cooling. This uses a 10 MWr dedicated surface refrigeration plant to generate cold water (10°C), which then drops underground into an energy recovery device (at EM this is a Pelton Wheelφ ) and then feeds underground Cooling towers and Spray chambers. Some underground bulk air cooling is essential to provide the flexibility the ventilation engineer needs to overcome the impact of higher heat loads on the deeper areas of the mine. Achieving this solely by surface bulk air cooling would result in some areas of the mine being over-cooled, an expensive solution. This is particularly the case if the fresh air intakes are shared with a separate mine, which is the case with EM, where intakes are also shared with the existing Isa Lead mine.
• Chilled service water. Chilled service water is widely used in South Africa. However, it is been found to be of low practical benefit at Enterprise Mine. This is because of the low usage of service water in areas where workers are exposed to heat. Very little ‘hand held’ mining is now done in Australia, and many drilling machines (which are the largest users of service water) now have air-conditioned cabins.
• Spot coolers: these are electrically-powered conventional split refrigeration systems installed underground, with the evaporator in fresh air and the condenser in return air. These can be used effectively for small heat loads (eg underground offices or cribrooms) up to about 300 kWr, where condenser heat can be rejected directly and locally into the return air (exhaust) circuit. Where close access to the exhaust circuit is not possible (most locations), conventional split systems + COP of a surface refrigeration process is about 5 to 6. The pre-cooling tower before the plant has a COP of between 20 and 50. After taking pump and other power consumers into account, typical surface plants have an overall COP of about 3. φ The Pelton wheel at EM recovers about 1 MW of electrical energy. It also reduces the effects of autocompression on the water itself. Without an energy recovering device, water will increase in temperature by 2.33° per 1000 m. With energy recovery, the temperature increase would be typically 0.8° or less.
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cannot be used. Because of the cold temperatures at the evaporators (resulting in condensation of air with dust on evaporator surfaces), these units are best suited to good quality intake air or the installation of an easily maintained dust filter, to ensure effective operation of the evaporator over time. A refrigerant that is safe for underground use must be employed.
ESTIMATING QUANTITIES OF AIR: THE PRIMARY VENTILATION REQUIREMENTS The most crucial steps in estimating the quantity of air required for the mine are:
• Determining the activities that will need to be ventilated: development, drilling, ground support, blasting, hauling, etc. This must be a complete listing and include allowances for fixed plant, plus areas where persons need to visit or travel.
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• Examine the mine development and production schedule over the life of the mine. This allows a schedule to be put together, based on the ventilation standards for each activity, showing ventilation requirements over the life of the mine. It is crucial to recognise that the volume requirement is a function of the number of workplaces to be ventilated and the standards at each workplace, not the number of persons working in the mine. An additional allowance should then be made. This merely recognises that:
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• Creating a design ventilation ‘standard’ for each activity. This standard must, having considered the required quality of air and the noxious substances produced in this working place, include a required volume of fresh air. Note that a good ventilation standard will include at least the following items:
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type of activity; method of secondary or auxiliary ventilation (forced duct, flowthrough, overlap, single or twin duct, etc); development sizes; limiting distances (note that this may not just be such the obvious ones as distance to the face, but also distances to truck re-muck bays); amount of diesel or other equipment in the workplace: types, kW, air-conditioned cabins (or not), etc; personal protective equipment (PPE) requirements of men in the work place; metabolic work rates of men in the work place (W/m2); leakage and losses of air: this is important as, even under good practice, these are real and relevant. In fact, in some cases, limited leakage from a ventilation duct may be desirable or even essential to ensure temperatures in the ‘outbye’ side of each workplace do not exceed the maximums allowed; a separate standard may need to be produced for some activities which have very different ventilation requirements in their ‘cycle’ (for example, a development face during boring versus during mucking); optimum volume and velocity of air (m3/s and m/s) in each workplace. This is a large topic in itself. However, for practical purposes, the following guidelines hold: where the mining operation is above the critical depth for all of the year, refrigeration is not required, and the optimum airflow is the minimum required to achieve acceptable air quality, or that required to provide acceptable air temperatures in the return airways during the summer period, whichever is the higher; and where the mining operation is below the critical depth during summer, refrigeration is required, and the optimum airflow will be the minimum required to achieve acceptable air quality, or about 0.5 to 0.7 m/s, whichever is the higherϒ. note that, even where diesel equipment is not operating
ϒ Note that at low refrigerated air temperatures, ‘losses’ start to increase significantly as these are related to the ‘driving’ force’, or ∆T, between the air and other surfaces, such as the rock and mobile equipment.
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but men are expected to work, it is considered essential to provide for some airflow. In many instances this will be related to the need to provide for cooling of any persons working in the area; and after developing ventilation standards, a summary of ventilation requirements can be produced.
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There are inefficiencies: a development crew may be allocated six ventilated workplaces, but sets up duct and fans for ten workplaces. In theory, only six fans should be on at any one time; in practice, double or more working places may need to be ventilated at any one time for this crew. A contingency must be provided to allow for estimation errors in the mine characteristic curve and/or the required air volumes: no-one designs a concentrator or hoisting facility to ‘just’ achieve the required tonnage. Some contingency is prudent and accepted as good practice to ensure the typical ‘peaks and troughs’ of production can be accommodated. Contingency also reflects the fact that the design is based on imperfect knowledge and that some change will always occur from the original assumptions or design criteria. These changes almost always impact negatively on ventilation requirements (ie need more not less). However, contingency only comes at a cost. A good and prudent design will allow for modest contingency that adds value, but not so large that it reduces the project value. It is a ‘balance of probabilities and consequences’ issue - a matter of managing risks. Large contingencies reflect either insufficient information being available or insufficient detail design having been completed.
• It is also crucial that the construction and pre-production phase of the project be fully taken into account. It is well recognised that the construction workforce in a project usually exceeds the on-going operating workforce. This results in there being more workplaces to ventilate during construction than during subsequent production. Sometimes the ventilation requirements during construction are more onerous than during production, and this is aggravated by the fact that a key part of the construction program is putting in the very ventilation system required to construct or operate the mine. Moreover, any delays during the critical and cash-consuming construction phase are likely to seriously impact on the construction cost, and hence project return. Any construction program for a new underground mine needs to be built around the four key obstacles to be overcome during the construction period itself: ventilation, mullock disposal, emergency egress, and logistics (getting men and materials in and out of the mine). The most logical construction program, from an engineer’s point of view, will rarely be the most effective or cheapest once these four key factors are taken into account. In addition to the impact of the construction workforce, it is also vital to recognise that a project in its construction phase has three other adverse impacts on ventilation:
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many of the rock surfaces are ‘new’ and therefore much hotter than they will be as they ‘age’. Old airways are cooler than new airways, disadvantaging a mine in its construction phase; - after the ventilation circuits are completed, the mine can operate more on flowthrough air, rather than ducted air. Flowthrough air is more reliable and effective than ducted air, again disadvantaging a construction project; and - during construction, the ore handling system is not operational (crushers, hoisting shaft, etc). Therefore waste and ore must often be trucked in diesel equipment. This large diesel loading is not required after the ore handling system is commissioned (if a hoist is being installed). Therefore diesel loads are usually higher and more concentrated in construction than during production. Note that various ‘rules of thumb’ also exist to estimate airflow requirements in mines. In the Australian context, with large-scale, highly mechanised mines and relatively large orebodies, a figure of 180 m3/s per Mtpa (plus 150 m3/s), or 3 to 3.5 m/s per kt per month is typical. Figures at the lower end of this range apply to cooler climates where heat is not a major contaminant OR to hotter climates or deep mines where refrigeration is required; figures at the higher end of the range tend to apply to moderate climates and depths where the main mechanism for controlling heat in the mine is by ‘flooding’ the workings with air. It is also important to recognise, particularly on deeper mines, that the density of air at the working place underground will be different to the density on the surface. Therefore volume requirements underground (which is where the requirements are set) will differ to those at the surface (which is where the air intakes are located), and affects the fan duties and the mine resistance curves (more dense air results in higher frictional pressure drops for any given airway).
OPTIMUM AIRWAY SIZES Most development sizes will be dictated by the size of equipment required to operate in the area. However, major fresh and return airways may need to be larger than this, to accommodate the necessary airflows. The trend to using larger equipment has generally been beneficial for ventilation personnel, as the larger development has had a major reduction on mine airway resistance. It is desirable, and in some cases due to statutory requirements, essential to have at least two fresh air intakes into the mine± and at least one exhaust to the surface. In most mines, it is convenient and most effective to bring fresh air into the mine on the same side of the orebodies as the major service accesses (usually the footwall), and to use the other side of the orebody (usually the hangingwall) to collect the return air and exhaust it to the surface. The basic design of the ventilation circuits must be matched to the mine production plans, stoping methods and the like. Linear programming or simulation exercises can be used to optimise the trade-off between surface refrigeration, airflows and airway sizes. However, the elaborate assumptions used in some refined optimisation exercises have a habit of coming undone in practice, after the mining operation has started and data has come to light that was unforeseen at the feasibility study stage. The danger with sophisticated optimisation protocols is that they are
± Most mining jurisdictions require two means of egress. This is best achieved using two fresh air intakes, as using an exhaust airway as the second egress could result in compromised escape or entrapment in the event of a fire.
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only as good, and as accurate and reliable, as the underlying data. It must be remembered that a mine is an extractive operation inside a ‘natural’ (ie not man-made) system (the geology of the mine) and therefore is both constantly changing (a dynamic process) and never perfectly understood in advance. Once a concept for the major airways has been established and the location and number of airways established, the optimum airway size can be established in a straightforward fashion as follows:
• Using a network analysis to simulate volumes, velocities, pressure drops and fan duty requirements for various airway sizes.
• Estimating capital and operating costs of airways and fans and solving for the lowest net present cost.
• For horizontal intake airways, setting a maximum velocity of 6 m/s, which is the practical limit before airborne dust becomes a serious problem. Horizontal exhaust airways, where dust is not an issue, can accept economic velocities up to 10 to 12 m/s.
• For vertical airways, a maximum velocity of 13 to 16 m/s is accepted as a good working limit, although with deep mines (and hence expensive shafts), velocities of up to 20 m/s are not uncommon. The ‘critical velocity’ range of 7 to 12 m/s should be avoided in wet shafts or where condensation could occur, as this is the point at which water droplets will remain suspended in the airstream, increasing shaft resistance and adding extra stress on the fan installation. Auto decompression must be taken into account in very deep mines.
• Towards the latter half of the mine’s life, major parts of the exhaust circuit may be closed off (for example, where stopes have been extracted and taken exhaust airways with them). Therefore the sufficiency of the ventilation system must be assessed in these latter years as well as in the early years.
FAN AND REFRIGERATION PLANT DESIGN The issues of fan and refrigeration design (sizing, type of installation, etc) are adequately covered in other sources (De La Harpe, 1989; Howes, 1983).
KEY ISSUES FOR VENTILATION AND REFRIGERATION DESIGN AT ENTERPRISE MINE Good ventilation practice is at least in part dependent on the particular mining operation. However, some of the issues to be considered at the Enterprise mine include the following.
Keeping heat sources out of the fresh (intake) airways to the working places • No Surface Intake (‘Push’) Fans: all electrical power input to these fans manifests as an increase in air temperature.
• Keep Circuit fans in intakes to the minimum. • Do not allow trucks to ‘idle’ in fresh air; set up park bays in return airways.
• Do not idle trucks or loaders under ventilation lines (all of which carry refrigerated air).
• Controlling Numbers of Diesel Equipment: Diesel Motors are about 30 per cent efficient. Hence One Elphinstone AD40 truck @ 367 kW at full loadØ generates 1 Ø’ Average’ engine loading is about 70 per cent of full load.
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MW in heat. Taking losses into account, the refrigeration system actually needs to generate more than 1 MWr of ‘coolth’ on the surface to service each AD40 at full load. Because these are such large heat loads, which will impact on environmental conditions underground, AND on capital requirements for refrigeration plant and operating costs, large heat loads must only be introduced after careful consideration of the knock-on effects on the local and overall ventilation and refrigeration plan, and on personnel.
Economical use of refrigerated air Because the total amount of air into the EM must be kept to the minimum practical (to minimise the heat load from autocompression and the hot surface temperatures), the air must be used to maximum advantage. Leakage and other losses must be kept to a minimum otherwise recirculation will occur and environmental conditions will deteriorate.
Maintaining ‘concentrated’ workings Mobile diesel plant also create significant humidity problems. Each litre of fuel burned produces about one litre of water vapour, but taking other factors into account, it is not uncommon for a diesel vehicle to produce seven litres of water vapour for each litre of fuel burned. It is therefore desirable to avoid wet scrubbers, due to the moisture they put into the air. Catalytic converters are the required standard at EM. A ‘Catch 22’ therefore potentially exists: if development rates fall behind, a tendency is to bring in more crews and equipment which results in more heat which results in lower productivity which results in development rates falling even further behind. One key issue for diesel equipment is how long it operates in a particular area. For example, a large truck that only enters a development end every 20 minutes has a much less profound impact on the airflow and refrigeration requirements in that development end than a truck working constantly. This is not only related to the time the truck is on location, but also due to the ‘thermal flywheel’ effect of the rock walls, which absorb a portion of heat when the truck is on-site and then expel it into the ventilation air when the truck leaves. This allows a less substantial refrigeration and, possibly, airflow requirement, than if the truck is continuously in the end.
• Use electrical equipment where possible: electric motors are much more energy efficient than diesel motors resulting in less heat and no noxious gasesÞ.
• Keep transformers and other heat sources out of the fresh air feeding the workings, wherever practicable. A modest 1 MVA transformer with an impedance loss of five per cent will generate 50 kW of heat constantly. Larger transformers will give off much more heat.
Keeping humidity out of the intake airways It is crucial that the intake airways be kept as dry as possible as wet or damp shaft or airway walls result in much higher heat flows from the strata into the air. Therefore nuisance water, particularly in the intakes, must be avoided. Wherever possible, it should be collected and pumped away.
Maximise air velocity on the job Use Air Movers to ensure velocity over the body is at least 1 m/s.
Direct feed from fresh air raises to the working place Where duct must be used to ventilate a workplace, air is fed directly from the fresh air raise (FAR) to each job. This avoids the situation where air blows out of the FAR and is then picked up somewhere else and ducted to the job. In this latter case, the air will be hotter than when ducted directly from the FAR. 20 m3/s of air in a 1.2 m diameter duct will reach the face of a 6 m x 6 m drive 400 m away in 23 seconds. The same airflow travelling through the drive itself would take 720 seconds. Clearly the heat picked up by air travelling in the duct will be inconsequential compared to that picked up in the drive.
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As the workings ‘spread out’, the heat load from the rock surfaces themselves can become the major internal heat load in the mine. It is vital that mining activities remain concentrated in the smallest practical geographical area.
Ventilation duct The weakest link in system is the ventilation duct. In the EM, air is ducted directly from each fresh air raise to the working placeð, so the EM is heavily dependent on ventilation duct. Installation standards and on-going maintenance of ventilation duct are critical. There should have no more than 30 per cent leakage for a well installed and well maintained duct.
Air conditioned cabins Maximise use of air-conditioned cabins: this productivity, safety and morale and reduces fatigue.
improves
Microclimate cooling When the circumstances are suitable, EM uses ‘cold vests’ for limited duration work in very hot conditions.
Judicious use of insulation Major chilled water lines (eg to bulk air coolers) are insulated. This is particularly important when these lines are located in return airways, as any ‘coolth’ lost from the pipes is totally wasted.
De-rating production and other schedules Production schedules (and all ancillary schedules) must be de-rated over summer for the EM, in accordance with the expected productivity loss on jobs that require high work rates. The fact that the burden of hot work will not fall evenly should be built into the schedules.
Effective maintenance and operation of the ventilation systems A properly planned and executed maintenance program for ventilation and refrigeration equipment is essential. The ventilation/refrigeration system is unique in that it doesn’t need the same output and performance all year round. However, it must achieve very high reliability and efficiency in summer; therefore as much maintenance as possible should be done in winter. Ventilation and refrigeration equipment must be operated so as to achieve maximum system effectiveness. It must not be
Þ However, even electric motors have their limits. At EM, motor failures on the Kiruna electric trolley assisted trucks have been traced to high dry bulb temperatures, which are now restricted to 35° DB. ð This is to ensure the air into the ventilation duct is as chilled as possible, thereby reducing temperatures at the working place, compared to other possible systems of auxiliary ventilation.
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operated to achieve the lowest running cost of the refrigeration plants themselves. The point is to optimise the efficiency of the system as a whole and in particular, to obtain optimum efficiency from the underground workforce.
Controlling changes to the ventilation/refrigeration systems There is no real difference between the ventilation and refrigeration systems and any other item of critical fixed ‘plant’. Any major changes to the design or operation of the system should be subject to the same standard as any other plant modification. These would include not only changes to ‘mine design’ but also to schedules (eg if ventilation raises are deferred)
Choice of clothing and personal protective equipment (PPE) Clothing has a significant impact on the ability of the body to cool itself via sweating as the following shows, based on a temperature of 28° C WB, 36° DB, air speed of 0.5 m/s and with safety hat, safety boots, socks and underwear: TWL for shorts only
201 W/m2
TWL for trousers and short-sleeved shirt
193 W/m2
TWL for trousers and long-sleeved shirt
184 W/m2
Note that PPE requirements are a part of the ventilation standard for each workplace. Therefore the choice of PPE impacts on airflow and refrigeration requirements, and should not be changed without examining the knock-on effects to productivity and ventilation costs. Key issues for PPE are: fabric vapour permeability and conductivity, clothing design (‘bagginess’) and amount of clothing and type of PPE. For example, leather boots are much cooler than rubber ‘gumboots’.
Dust explosions Dust explosions can be a potential risk in all coal mines and many operations mining fine-grained sulphide ores. Special precautions are required to manage the risks of dust explosions.
Condensation and Dew Point Mixing cold and hot humid air in a mine can create unexpected problems. These range from relatively minor matters such as the dripping of condensation from chilled water lines onto major road haulages creating road maintenance problems, to major issues such as the head ropes on the U62 hoisting shaft (a Koepe winder) slipping due to condensation on the ropes, resulting in a full skip of ore crashing to the bottom of the shaft. Wet metal surfaces due to condensation also increases the likelihood of corrosion due to rusting, and results in the build-up of more dust and grime than otherwise. In some areas, high humidity can result in deterioration of the rock strength (ingress of moisture into cracks) or shortened working lives of ground support due to corrosion.
Instrumentation Over the past 20 years, most mines have heavily instrumented their fixed plant and mobile equipment. This has improved performance, increased availability and reduced maintenance and operating costs. Sadly, few mines have properly instrumented the ventilation system [except perhaps the surface fans]. At the EM, the ventilation and refrigeration (VAR) system, being as crucial
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as it is to safety, health, productivity and costs, will be significantly instrumented. Some of these instruments are as follows:
• The top of important fresh air raises (FARs) will be monitored at Mine Control continuously for wet bulb (WB), dry bulb (DB) and CO (carbon monoxide), and at the bottom for WB and DB.
• All Circuit fans will be monitored continuously for motor amps. This provides a ‘surrogate’ for airflow and also indicates on/off.
• All underground Cooling Towers will be monitored at Mine Control continuously for WB, DB and CO. The Cooling Towers will also be monitored at Mine Control continuously for Chilled supply water flowrate and fault (on/off) indication. This will allow rapid fault-finding of any problems in the system, which will impact beneficially on productivity, costs, mine output and morale. It will also allow assessment of quality of air in the fresh air raises in the event of a fire, and provide a much better understanding in future years as to how the underground system responds to changes in surface conditions, refrigeration output, etc.
Fault-finding and training of personnel: supervisors, managers and workers in ventilation circuits This has already been mentioned. However, instrumentation alone will not result in the highest realistic standards of up-time from the ventilation system components or the best practical environmental conditions. Supervisors, especially, need to be trained as to the location of fans and other controls and how to make simple adjustment to underground cooling towers, etc. Plus controls on who and when plant is adjusted or maintained need to be established and enforced.
Systems to audit the ventilation system The importance of measuring the effectiveness of the ventilation system cannot be overemphasised. The elements of an effective audit process are:
• Identifying the customers and other stakeholders and understanding their needs: at the end of the day, the ventilation system has ‘customers’ who have views on what ‘fitness for purpose’ and ‘value for money’ mean in their context. The customers and their needs must be identified.
• KPIs: key performance indicators are the means of ‘keeping score’ and ensure the customers are getting a product that is both fit for purpose and value for money.
• Ventilation and Refrigeration standards: these are essential to ensure the customer gets what he is after. Many defects in the ventilation system are caused by the customer himself, so this requires careful but firm handling.
• PFDs: process flow diagrams are the essential pre-cursor to providing an adequate means to determine instrumentation requirements of the system, scope and regularity of audits, and an effective, timely, fault-finding response.
• Communication: it is vital that a process be set up and followed by all parties to ensure that defects are identified early and fixed promptly and that future mine design and schedules are compatible with future ventilation design (or vice-versa). With a well designed ventilation and implemented system, the key role of the ventilation staff becomes one of advising on necessary changes to the ventilation circuits and operating strategy as the mine develops, and auditing the effectiveness of
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individual components and the system as a whole. At EM, an audit process has been developed that encompasses, among other matters:
• daily checks of key ventilation components and conditions
ACKNOWLEDGEMENT The authors would like to thank Mount Isa Mines Limited for permission to present this paper.
and monitoring of KPIs;
• follow-up within 24 hours with written reports on all serious ventilation defects and failures, especially on ‘dust reports’ and ‘hot jobs’ or heat illnesses;
• issue of non-compliance reports for serious breaches of ventilation standards; and
• issue of bi-weekly and end of month KPI reports. ESCAPE (EGRESS) AND ENTRAPMENT When the ventilation system fails at EM during summer, temperatures will increase rapidly and the airflow throughout the mine will fall almost immediately. Natural ventilation pressure is at its lowest in summer, and the elaborate ventilation networks and non-return dampers installed on most underground fans will mean that airflows through the orebody will rapidly diminish. Experiments in the mine (Brake, 1997) have shown that the WB increased from 28° to 33° within about 90 minutes and to 33° to 36° WB (depending on the location) within two hours, with temperatures continuing to increase after two hours (albeit more slowly). These conditions have major ramifications on egress and entrapment, and consideration work has been undertaken to ensure mine workers can be kept in a safe condition, or escape safely, in the event of power failures or mine fires (Brake and Bates, 2000; Brake, 1999).
CONCLUSIONS The Enterprise Mine will be Australia’s deepest underground mining operation and will be among the hottest underground mines in the world. It will be the working environment for several hundred workers, technical and support staff and visitors for at least 15 years. During construction, a workforce of almost 700 persons needed to be ventilated and kept cool. With high virgin rock temperatures and the adverse Mount Isa surface summer conditions, the workforce will be exposed on a continual basis to adverse thermal conditions. Mount Isa Mines has developed a comprehensive system of engineering and other solutions to managing the risks associated with working in heat.
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REFERENCES Bates, G and Matthew, B, 1996. A new approach to measuring heat stress in the workplace, paper presented to The Australian Institute of Occupational Hygienists 15th Annual Conference, Perth, 30 Nov to 4 Dec. Brake, D J, 1997. Test of Survival Conditions – Enterprise Mine. Unpublished Internal MIM report. Brake, D J, Donoghue, M D and Bates, G P, 1998. A new generation of health and safety protocols for working in heat, in Proceedings 1998 Qld Mining Industry Health and Safety Conference, pp 91-100 (Queensland Mining Council). Brake, D J, 1999. An integrated strategy for emergency egress from an underground metal mine, in Proceedings 8th US Mine Ventilation Symposium (Ed: J C Tien) Pp 649-657 (Uni Missouri-Rolla). Brake, D J and Bates, G P, 1999. Criteria for the design of emergency refuge stations for an underground metal mine, The AusIMM Proceedings, 304(2):1-7. Brake, D J and Bates, G P, in press. Limiting metabolic rate (thermal work limit) as an index of thermal stress. De La Harpe, J H, 1989. Basic fan engineering, in Environmental Engineering in South African Mines, Chp 7, (The Mine Ventilation Society of South Africa). Donoghue, A M, Sinclair, M J and Bates, G P, 2000. Heat exhaustion in a deep, underground, metalliferous mine, Occupational and Environmental Medicine, 2000(57):165-174 Hemp, R, 1989. Sources of heat in mines, in Environmental Engineering in South African Mines, Chp 22, (The Mine Ventilation Society of South Africa). Howes, M J, 1983. Application of refrigeration in mines, Trans Instn Min Metall (Sect A: Min Industry), 92, April 1983. Pickering, A J and Tuck, M A, 1997. Heat: sources, evaluation, determination of heat stress and heat stress treatment, paper presented to Heat and Noise in Underground Mining Symposium, Nottingham, 17 April. Tranter, M and Abt, G A, 1998. The assessment of metabolic rate, core body temperature and hydration status during underground coal mining, in Proceedings of the 1998 Safety Institute of Australia Annual Conference, Gold Coast, 1998, pp 293-303 (Safety Institute of Australia). Van Rensburg, J P, Marx, H E, Van Der Walt, W H, Schutte, P C and Kielblock, A J, 1991. Estimated metabolic rates associated with underground mining tasks: conventional and mechanised mining operations, Ref 11/91: GE1B (Chamber of Mines Research Organisation).
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Bulk Low-Grade Mining at Mount Charlotte Mine P A Mikula1 and M F Lee2 ABSTRACT Although the Mt Charlotte gold resource was discovered in 1893, it was only in 1954 that geologists correctly interpreted drill data and defined a large low-grade orebody. The mine has been a large, low-cost and profitable producer since 1962 when bulk mining commenced using cut and fill. From 1968 a combination of long hole open stoping with delayed mass blasting was used. This proved to be an excellent mining method for the large, separate subvertical orebodies. The structured but otherwise good quality rock allowed very large stopes to stand unsupported. Exceptionally large equipment, for its time, was used to efficiently mine, muck and move the ore at low-cost. The mine only just survived severe cost pressures in 1974 - 76, due to very low gold prices. Better gold prices in 1979 and a need to access deeper ore meant a bold commitment to sink the Cassidy Shaft and refurbish the mine to sustain future profitability. In 1997 the Sam Pearce Decline connected to the underground workings, making it profitable to mine upper level remnant ore. Mining-induced seismicity has been a feature of mining at Mt Charlotte. The main cause has been shearing on structures. A mine-wide seismic system was installed in 1994. Large seismic events have occasionally made management question whether a safe working environment can be maintained. Thorough investigations led to an understanding of the problem and development of procedures to minimise seismicity.
INTRODUCTION The Mt Charlotte mine is located 5 km north of Kalgoorlie’s famous Golden Mile. It is presently operated by Kalgoorlie Consolidated Gold Mines (KCGM), which also operates the Fimiston Superpit centred over the old Golden Mile (Figure 1). Since 1962 the mine has produced 33.9 million tonnes at 3.4 g/t for 3.6 million ounces of gold. Production peaked in 1998/99 at 1 740 000 tonnes, making Mt Charlotte at the time Australia’s highest tonnage underground gold mine. This paper summarises Mt Charlotte’s history and mining methods. Key management strategies, several innovations and the application of a good understanding of the rock mass are also discussed. The latter have allowed the mine to be profitable and produce for over 37 years in the face of economic and seismic pressures.
BRIEF HISTORY Three Irish prospectors, Hannan, Flanagan and O’Shea, found the Kalgoorlie Goldfields in 1893. The discovery site was close to where Mt Charlotte’s Cassidy Shaft now stands. The resulting goldrush centered on the fabulously rich Golden Mile to the south. Underground mining at Mt Charlotte occurred intermittently and by 1945, small-scale workings had progressed to a depth of 215 m. Small open cuts were also worked. Exploratory drilling continued under various owners, until 1962 when Gold Mines of Kalgoorlie (GMK) acquired ownership (refer Table 1 chronology). By this time the size and grade of the orebody was correctly understood. Its ‘large dimensions and compact shape were a challenge to mining engineers to develop a method for its profitable extraction in toto’ (Brodie-Hall, 1988). 1.
Senior Rock Mechanics Engineer, Mt Charlotte Operations, Kalgoorlie Consolidated Gold Mines, PMB 27, Kalgoorlie WA 6433. E-mail: [email protected]
2.
Director, Australian Mining Consultants Pty Ltd, Level 19, 114 William Street, Melbourne, Victoria 3000.
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FIG 1 - Mt Charlotte mine location.
Bulk underground mining became possible when in 1962 the Department of Minerals and Energy allowed diesel engines underground for the first time in Western Australia. Mining commenced using mechanised cut and fill (C&F). The method was soon changed, however, to long hole open stoping with delayed mass blast (LHOS-MB). This was used to mine the Charlotte and Charlotte Deeps Orebodies (COB and CDOB, Table 2, Figure 2) until 1998. LHOS was used to mine the Reward and Maritana Orebodies (ROB and MOB). More recently, sublevel caving (SLC) has been used to mine remnant ore around old stopes. Economic pressures developed in 1973 when the gold price began falling significantly. Cost-cutting measures affected the Kalgoorlie mines and by November 1975 all the high cost Golden Mile operations had closed. The gold price dropped further, and in mid-1976 it was decided to close Mt Charlotte on 10 December 1976. Retrenchments began, but the gold price rallied and on 9 December the owners reversed the earlier decision. The next few years saw a dramatic reversal of Mt Charlotte’s fortunes. The gold price was so high that profit paid for a total refurbishment of the mine, including a new shaft, crusher and orepass system. A significant increase in production rate occurred from 1988 due to management pressures following changes in company ownership. Mining-induced seismicity has also threatened the mine’s survival. Production flexibility plus rock mechanics knowledge has, however, enabled the mine to minimise the effects of the occasional pillar failure, fault-controlled overbreak of stope backs and walls, and seismicity.
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Overview of performance
TABLE 1 Mt Charlotte key chronology. Date 1962
An important measure of the performance of a mine is the bottom line. Mt Charlotte costs have usually been low, and in fact production costs per ounce recovered have sometimes been cheaper than the Superpit mining at ten times the scale (eg 1992 93, $380/oz for Mt Charlotte, $398/oz for the Superpit). Peak performance was achieved in the early-1990s, when ore was mined, hoisted, crushed to 10 mm and delivered 5 km to the mill at $16.50 per tonne. The 2000 budget figure is $18 per tonne to the surface. Various performance data for 1962 to 2000 are described below:
Event GMK commenced bulk mechanised mining - cut and fill method
1968
Method change to open stoping
Early 1973
Financial pressure - Kalgoorlie Lake View (KLV) formed (partnership of GMK and Lake View and Star Ltd)
Dec 1975
Financial pressure - KMA formed (KLV and Homestake Gold partnership)
Aug 1976
Mine closure pending - retrenchments began
9 Dec 1976
Closure of the mine averted
Aug 1981
Cassidy shaft construction commenced
May 1985
Cassidy shaft commissioned
Apr 1988
Alan Bond’s Dallhold Resources acquired control of KLV, replaced WMC as manager of KMA, and significantly increased the mine production rate.
• Production rate (Figures 3 and 4). The increased scale of production from 1990 was accompanied by a fall in both grade and productivity. (Unfortunately pre-1991 productivity data could not be determined.) Since 1991, target tonnages (up to 1.8 million t/yr) were not reached, despite the extra hours worked, due to stope production constraints (refer the section entitled ‘Sustainable Production’). The doubling of production to 1.6 Mtpa came at the cost of 1.3 g/t lower head grade (a revenue loss of about $17/tonne). However productivity increased after retrenchments in 1998 and 1999 and the commencement of remnant mining on upper levels.
29 Mar 1989 KCGM formed (Joint Venture GMK and Homestake Gold of Australia) Aug 1989
Normandy-Poseidon acquired control of GMK; Alan Bond resigned from GMK
1993
400 000 manhours achieved without a Lost Time Injury
2 Sept 1996
Change from 8 hr shifts to 12 hr continuous roster
Dec 1997
Sam Pearce Decline connected to mine workings
Dec 1997
Launch of new safety initiative
Sept 1998
Retrenchments (210 down to 150 personnel)
Dec 1998
End of deep mining - shift to near surface remnant mining
July 1999
Retrenchments (150 down to 80 personnel)
• Development rate (Figure 5) was erratic, influenced by cost pressures in the 1970s, and the very large ROB5 stope coming on line in 1990. Remnant mining from 1999 returned high tonnages for very little development. The expanding production rate is reflected in vertical advance data in Figure 6.
• The cost data in Figure 7 have been adjusted to June 2000 dollars using the Manufacturing Production Price Index. Although costs were not available for all years, the trend of costs increasing with production is clear. The typical breakdown of key costs is shown in Figure 8.
GEOLOGY AND STRESS FIELD
10 Aug 1999 Relocation of site surface infrastructure from Cassidy Shaft to Sam Pearce Decline portal
The Mt Charlotte resource comprised several thick poddy orebodies (Figure 9) dipping 80° to the west, 40 to 90 m thick, up to 300 m along strike and 700 m down dip (directions relate to Mine Grid North = 38° west of True North). The ore is essentially a mineralised ‘stockwork’ that has preferentially developed in the central siliceous granophyre section (Unit 8) of the thick Golden Mile Dolerite (Bischoff and Morley, 1993). The stockwork comprises two sets of quartz veins with gold typically associated with alteration haloes (carbonation) adjacent to the veins. It occurs as free gold in fractured euhedral pyrite crystals.
The Sam Pearce Decline intersected Mt Charlotte’s workings in December 1997. This made it feasible to mine lower grade remnant ore in the upper mine when deeper reserves were exhausted in 1998. The last deep ore was hoisted in Cassidy Shaft in March 1999. The new decline also made it possible to profitably muck drawpoints and handle fired ore that had previously been abandoned, due to dilution and contamination with old red sand fill from the earlier C&F operations. It is currently planned to close the mine in 2001 when the remaining ore reserves have been extracted.
TABLE 2 The major Mt Charlotte stopes. Stope COB A
Mining method
Average depth (m)
Strike length (m)
Width (m)
Height (m)
Stope Reserve (tonnes)
C&F
100
190
70
140
1 800 000
COB B to H
LHOS-MB
500
235
45 to 90
590
19 750 000
CDOB
LHOS-MB
900
265
30 to 55
150
3 080 000
ROB2
LHOS
200
280
50 to 55
205
3 420 000
ROB3
LHOS
350
260
60
150
2 230 000
ROB4
LHOS
500
110
50
120
560 000
ROB5
LHOS
700
280
80
155
3 140 000
MOB4
LHOS
700
115
25
115
770 000
S2
LHOS
750
105
25
100
730 000
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Tonnes Ounces Productiv ity (t/hr)
Head grade Recov ery Grade
6 5 4 3 2 1 0 1964 1969 1974 1979 1984 1989 1994 1999
FIG 3 - Production tonnage and grade history.
FIG 2 - Long-section showing mine layout, stope blocks and major faults.
head Grade (g/t)
6 5 4 3 2
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0
500000
1000000
1500000
2000000
Annual Production (t)
FIG 4 - Life of mine production/grade performance.
5000
2000
4000
1600
3000
1200
2000
800
1000
400
0 1962
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Tonnes/Devt m
Development Metres Tonnes/Development metre
Development (m)
Two major reverse fault sets divide the rockmass and the ore into lozenge-shaped blocks. The older set dips moderately to the west (Neptune, Flanagan and Shea Faults). The other dips steeply to the northwest (Charlotte, Reward and Maritana Faults). The infill of both sets is typically thick healed foliated zones. Continuous structures (faults and shears) subparallel to the main faults are common. Structures are pervasively developed on the scale of large LHOS, and have been diligently mapped over the years, so that it is rare to come across an unknown fault in the upper mine in recent times. The Golden Mile Dolerite is stiff (E = 65 GPa) and strong (UCS = 175 MPa). The rock mass comprises well-interlocked blocks and has an RQD close to 100 per cent. Joints are rough/irregular, undulating and have thin chlorite infill. Typically, two plus a random joint orientation set are developed in any area. The worst structures for the stability of underground openings at Mt Charlotte dip moderately to the west, eg in strike drives and long stope backs. Further geomechanical data can be found in Mikula (1999a). Where it occurs, groundwater is very saline, but the mine is generally regarded as dry. As in other Western Australian underground mines, corrosion of exposed support (bolts, mesh) can occur rapidly. Pre-mining stresses have been measured on numerous levels (Lee, Pascoe and Mikula, 1999). The stress field is high and deviatoric. Typical magnitudes at 1 km depth are σ1 = 75 MPa, σ2 = 40 MPa, and σ3 = 25 MPa. Most of the mine’s structures are loaded close to their in situ shear strength in the virgin rock mass. Small reductions in normal, or increases in shear stress can therefore initiate shearing, some of which can be seismic.
0 1972
1982
1992
FIG 5 - Development history.
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COB stopes vertical advance, m/yr ROB stopes vertical advance, m/yr Production/Vertical advance, '000 tonnes/m/yr
50 m/yr, '000 t/m/yr
40 30 20 10 0 1962 1967 1972 1977 1982 1987 1992 FIG 6 - Vertical advance rate history.
500
36
400
30 24
300
18
200
12
100
6
0
FIG 9 - West-East cross-section showing orebody pods offset by west-dipping Neptune and Flanagan faults.
Cash cost ($/tonne)
Cash cost ($/oz)
Indexed Total Cash Cost ($/oz) Indexed Total Cash Cost ($/t)
0 500000 1000000 1500000 2000000 Production Rate (t/yr)
0
FIG 7 - Life of mine cost performance.
MINING METHODS Cut and fill (C&F) Bulk mining started in A Block of COB in 1962 using C&F (refer Figure 2 and Table 2). Stopes were 15 m wide between 9 m pillars. Fill was mostly mill residue, introduced via passes. At the time C&F was chosen over LHOS because it was intended to leave 20 per cent of the lower grade rock as pillars. This turned out to be impractical due to increasing production requirements. Pillars recovery, by SLC between fill walls, was also difficult.
Long hole open (LHOS-MB) 1998-99 0
1999-00
Percentage of operating cost 10 20 30 40
Dev elopment Stope ground support Drill & Blast Load & Haul Crush Mine Serv ices
FIG 8 - Key cost breakdown data.
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stoping,
with
mass
blasting
Method LHOS-MB was used to mine most of COB and CDOB. The method was adopted from the Algoma iron ore mine in Canada. It has allowed profitable production of about 95 per cent of the resource. A similar method was used at the Mt Lyell copper mine between 1972 and 1994 (Usher, 1992; Weston, 1998). The method involves the following steps (Figures 10 and 11):
• Mine a large stable primary LHOS (Figure 12), by sequentially firing 10 000 to 50 000 tonne rings, sometimes over several levels, while rib and crown pillars retain waste rockfill.
• Mass fire the crown and the adjacent rib pillar into the primary stope. Complex firings of up to 900 000 tonnes were common at Mt Charlotte in the late-1980s and 1990s. Blasts are designed to make the mass blasted rock just fit into the available void (primary stope + rib + crown pillar). Horidiam rings, from opposite corners of the stoping block, were often used to fire highly stressed and overbroken crowns.
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• Muck mass blasted ore from drawpoints (Figure 13) under
• maximum charge weight per delay = 250 kg; minimum delay
the caving rockfill, which in COB extended to the surface glory-hole. This rockfill improved regional ground stability around old stopes.
• dilution factors 120 to 140 per cent of the fired tonnes at 80
• Use well laid-out drawpoints for good draw. Stopes had longitudinal trough undercut (TUC) drives, with drawpoints often on both sides. If the ore width was excessive, a double undercut system was used.
• Use SLC retreat of drawpoints to the orebody limits, to scavenge any dead zones in the broken ore column above, and to recover trough ore. Until about 1994 production was satisfied by mining only one large stope at a time. Production flexibility relied on the number and availability of drawpoints. To ensure uninterrupted production, mine planning had to be well in advance of extraction. Ore was trucked to either the 15 Level crusher and hoisted up Reward Shaft, or tipped into the orepass system to the 36 Level crusher, and then hoisted up Cassidy Shaft. During the 1990s, the key LHOS-MB parameters were:
• maximum sublevel interval, 60 m between drill horizons, 80 m TUC to drill horizon;
• development drill drives 4 × 4 m, TUC 5 × 5 m, main access/ decline 6 × 5 m with arched backs;
• TUC blastholes 57 mm main rings, 2 m burden × 3 m toe spacing;
• production blastholes, main rings 140 mm, 4 m burden × 6 m toe spacing;
• slot 140 mm vertical holes, 3 m burden × 3 m spacing; • explosives included ANFO and bulk emulsion; powder factor
interval = 25 ms; and to 90 per cent of the estimated grade.
B Block Existing shafts serviced the B block stopes, which were 90 m high, with 30 m crown pillars. Stopes were drilled using vertical 51 mm diameter blast holes. Horidiam raises were used to drill 51 mm percussion holes in crowns. Mt Charlotte’s first large seismic event occurred when the 260 m deep B2 rib pillar failed by shearing on a structure (Bamford, 1971).
C, D and E Blocks Decline access was needed for equipment, so stope height was limited to provide earlier production tonnes. Short lead times were also desirable during the lean economic times. Several pillar failures and some extensive fault-related back overbreak occurred.
F and G Blocks When mine closure was averted in 1976, two factors worked in Mt Charlotte’s favour; the mine workforce was effectively new and much smaller, and capital expenditure became possible due to profits from the high gold price. Management used this opportunity to improve production by moving to taller stopes in F and G Blocks, using 165 mm diameter blastholes, and cable dowelling stope backs to reduce overbreak and improve crown pillar stability. To optimise the placement of cable dowels, special drives were mined on the top level of G Block.
= 0.20 – 0.35 kg/t;
FIG 10 - Typical open stope and crown pillar drilling pattern.
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FIG 11 - Typical arrangement of double-sided drawpoints, open stopes, trough undercuts and crown pillars.
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FIG 12 - View on 19 Level of the slot opened up for the commencement of the G4 open stope. Slot length is 47 m.
• regular array of drawpoints, to allow mucking to dilution
CDOB I1 and I2 These stopes were offset about 200 m down-dip of COB along Flanagan Fault. To mine them efficiently, a deeper vertical access was needed. Commenced in 1981, Cassidy Shaft was sunk to 1184 m depth. The internal decline access was completed to 32 Level (950 m) by 1990. Two very large stopes (I1 and I2) were planned, separated by the small I1 rib pillar. Flanagan Fault was parallel to, and in some places exposed in the stope backs. The main design issues were:
• Cable bolting for stability of the back. Although Flanagan Fault was undercut, excessive overbreak was delayed by arrays of bulbed cable bolts.
• The stability of the I1 rib pillar. Very high stresses were
predicted in this small squat pillar (33 m wide × 44 m thick × 94 m high). Favourably oriented structures were not available for the pillar to yield gradually. It was therefore intentionally preconditioned (weakened by blasting, see the section entitled: ‘Preconditioning and Destressing’) before the first I1 mass blast (Mikula, Lee and Guilfolye, 1995).
• Stope backfill was needed. During the I1 mass blast a 30 m × 25 m cross-section fill pass was fired through barren rock to link CDOB with the old COB workings above. Waste rockfill dumped at the surface was then able to report to the 32 Level drawpoints at no extra cost.
• Seismicity was anticipated. Given the depth of mining, and the preconditioning of the I1 rib, it was considered appropriate to install a mine-wide seismic monitoring system (Mikula, 1999b).
cut-off grades if wall or back overbreak occurred, but also because remote mucking was not practiced; and
• drawpoints were often SLC retreated to the orebody limits. ROB1-2 Bulk mining started in 1962, as sublevel benching between 3 and 5 Levels. From 1968 LHOS was chosen over mechanised C&F on a cost basis. In 1986 the stope was enlarged. Instead of conventional LHOS, and because access was difficult, one sublevel was fired using horidiam rings. The excellent rock quality and the limited stope width allowed ROB2 to be, at that time, one of ‘the largest free-standing open stopes in the world’ (Reed, 1982). Subsequent Mt Charlotte stopes were similarly large (Table 2). Some rock fill was added to ROB2 via a pass from the surface.
ROB 3 Wall dilution was large at 18 per cent (Bridges, 1982) mostly from an undercut section of the west stope wall.
ROB5 It was initially planned to mine ROB5 as series of clean primary stopes between permanent rib pillars. Pillar sizes and stability issues, however, meant a low extraction ratio. Pillar bursting, by shearing on intersecting structures, was considered likely and seismic damage could affect the nearby Cassidy Shaft. Instead, ROB5 was mined as a single large stope, accepting that wall and back overbreak was likely. This meant adopting the following operational strategies:
• Intentionally arching or cable dowelling stope backs. The
Long hole open stoping (LHOS)
onset of back overbreak was also managed by mining a narrow stope, then mass firing wall extensions.
Method Unlike the continuous Charlotte orebody, the other orebodies (ROB, MOB, CDOBS) were separate pods, each mined as large conventional LHOS comprising:
• transverse slots in the widest section of the pod; • widely spaced sub-levels, mostly 140 mm diameter down holes;
• upholes in undercut drives, and only a few rings fired ahead of full-face main rings;
• shaped and /or intensively cable dowelled crowns;
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FIG 13 - Typical drawpoint ore.
• Walls were naturally arched to the shape of the pod, but some loosened zones close to accesses had to be cable dowelled.
• Mucking overbreak to a break-even grade. Because overbreak tended to occur away from the advancing mining face, and it was also low grade material, fired ore was typically not diluted.
• The stoping sequence was varied to minimise potential seismicity due to shearing on structures. Shearing on a steep NE-SW striking structure near Cassidy Shaft did occur (2.6 on the Richter Scale) and the concrete lining of the shaft suffered minor cracking.
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• The deflection of Cassidy Shaft was periodically monitored as ROB5 was stoped past. Unacceptable shaft displacements did not occur. All of the ROB5 ore was mined, but dilution meant 120 per cent of the tonnes at 92 per cent of the estimated grade. Most of the overbreak occurred within direct mucking distance of the orepass system, it did not have to be trucked and tipped. At the time, the mine also had excess hoisting capacity.
SLC remnants When mining below 32 Level was found to be uneconomic due to lower ore grade and increased costs, attention focused back on the remnant ore in the upper level:
• COB B to E Blocks had been mined at a higher cut-off grade. The gradational nature of the ore meant there was a skin of payable ore, in places up to 50 m thick, surrounding the old stopes.
• Old drawpoints had been abandoned 20 to 30 years earlier when dilution by mill tailings (difficult to hoist) and waste rock (lower grade material) became excessive. However, lower mining costs and better treatment methods have since made this material payable, provided that the tailings could be recovered without using orepasses or hoisting. The Sam Pearce decline from the surface to 9 Level made this possible in 1997. Upper level remnant mining has meant an extra four years of mine life. Old drawpoints are systematically accessed and mucked to break-even grade. They are then retreated to the old stope limits, followed by transverse SLC of any remnant skins. Ore is trucked to the surface.
FIG 14 - The built-in-house rockbolter with hydraulic drill boom remotely controlled from the HIAB work basket.
STRATEGIES, TACTICS AND INNOVATIONS Orebody definition The Mt Charlotte ore resource is unusual, as it contains two main sets of quartz veins carrying gold. One set dips shallowly to the north, the other steeply north. Considerable exploratory drilling of the resource was done before 1950, but at that time it was assumed that the veins were all transverse to the orebody. Irregular high assay values were cut, but the results still indicated a very variable orebody, difficult to understand. Only in 1954 did geologists recognise the actual vein system and calculate a drilling direction that would minimise bias from any of the vein sets (Brodie-Hall, 1988). This strategy gave constant grades without arbitrarily cutting high assays, and a large 4.5 g/t orebody was finally defined. This drilling direction has been adhered to ever since, and has meant mining specially oriented diamond drill drives a couple of years in advance of stoping. Calculated stope grades match very well with production grades.
Ground support - development The evolution of ground support at Mt Charlotte was guided by rock mechanics input (Mikula, 1999a). Support levels have been high, and this is considered to be a major factor in the rarity of rock fall incidents:
• Spot bolting was initially used in the narrower upper level drives.
• Pattern bolting, using ten tonne mechanical shell anchor bolts and W-straps, spanned the period 1976 to 1990. Bolting off the dirt or from a loader bucket was abandoned, in favour of a far more efficient approach using a modified tractor with a high-reach work basket. The mine still has the benefit of both the original machine and the original operator 22 years later.
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FIG 15 - Example of flat back bolting pattern (0.6 to 0.9 m spacing) used 1987 to 1995.
• Shell bolts were subject to corrosion and became unreliable with time, and were replaced by 2.4 m long, 20-tonne capacity, 20 mm diameter, resin-anchored rebar bolts, in 28 mm diameter holes, from about 1987. For this bolt, a dedicated rock bolter was built in-house by attaching a hydraulic drill boom and a HIAB basket to an old loader frame (Figure 14). All functions were controlled remotely from the basket. This allowed quality assured bolting of development (close inspection of backs, manual scaling, operator always under supported ground, fast drilling of the correct hole size, machine tightening of nuts).
• For a time, drives were mined with flat backs and the bolt patterns were dense (Figure 15). Rock mechanics analyses led to the adoption of an arch back (except in drawpoints), the bolt pattern was widened, and later the W-straps were discarded (Figure 16).
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FIG 16 - Example of recent arch back drive with wide bolt pattern and no W-straps.
• If seismic potential is high, cone bolts are installed as secondary support in backs, and standard ungrouted Split Sets in walls. Cone bolts in walls tend to be damaged by equipment.
• If the rock mass is adversely structured, 4 m or 8 m long cable dowels are installed.
FIG 17 - COB stopes long-section showing significant extent of back overbreak in B to F stopes. Cable bolting controlled G Block overbreak.
• The good quality rock means that in most areas, scats can be
90
• Owing to cost, shotcrete has only been sparingly used,
80
mostly confined to key access drives that are intersected by major structures.
Ground support - stopes No ground support was used in stopes until 1979, when cabling was trialled unsuccessfully in E3, F1 and F2 stopes. Then under an AMIRA grant in 1982, F3 stope backs were successfully cable dowelled using plain strand cables. The G Block stopes and parts of the ROB5 stope backs were similarly cable dowelled. Later birdcage cable was trialled in ROB5, and bulb cable in CDOB. The main support problem in stope backs is overbreak under continuous west dipping structures striking parallel to the orebodies (Figure 17). After much experience, good and bad, the optimum solution for Mt Charlotte stope back support is considered to be to avoid the need for cables at all, by mining self-supporting arches. The optimum height of the arch is related to its width as shown in Figure 18. However if an arch shape is not feasible, stability is achieved by:
• Keeping primary stope widths narrow (under 30 m) with extension mass blasts to extract wider ore.
Height of Arch in Backs (m)
effectively controlled by check scaling, so meshing of backs is not normally required.
70 60 50 40 30 20 10 0 0
10 20 30 40 50 Stope Minor Span (m )
60
FIG 18 - Data relating minor stope span to height of arching of stope backs. Unsupported backs of sufficiently long stopes tend to overbreak to a natural arch of a height indicated by the black squares. Triangles indicate stope backs blasted into a very high arch profile. Circles indicate short stopes with end support to backs.
• Locating transverse cable dowel drives at about 15 m centres, directly under the proposed backs. This allows installation, and most importantly effective surface plating, exactly where the support is most needed. Fanning cables towards stopes, from some remote point, has not been very successful at Mt Charlotte.
• Installing very dense cable dowel patterns, to form ‘ribs’ of tightly reinforced ground, or artificial abutments, and strapping collars together to prevent unravelling of ground from between the cable collars. Overbreak of the rock mass between these ribs is controlled. It is understood that this ‘rib-reinforcement’ approach to supporting LHOS was first done both at Mt Charlotte and at Mt Isa in about 1982.
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Seismicity Mining induced seismicity and associated ground control problems have a long history in Kalgoorlie’s Golden Mile. About 80 events of Richter magnitude 2 or greater have been recorded at Mt Charlotte since bulk mining started; the largest was magnitude 3.5 on 27 June 1998. Mt Charlotte took a step forward in 1994 when a PSS mine-wide seismic monitoring system was commissioned, the first operational seismic system to be installed in an Australian underground mine (Mikula, 1999b). It led to a step change in Mt Charlotte’s knowledge of ground behaviour:
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events, and were able to make informed decisions on immediate responses.
• It was found that the seismic disposition of a new stope block was generally evident as soon as slot extraction commenced. A quiet slot tended to indicate that the stoping would also be comparatively quiet.
• Assessments of patterns and trends in seismicity led to a better understanding of event mechanisms. In time, this resulted in a 14-point Seismic Risk Reduction Strategy (Table 3) being developed for use at the stope design stage. Notwithstanding the above, predicting the chance and size of seismic events is a difficult rock mechanics issue at the mine (Lee, Beer and Windsor, 1990). Good conditions now, or in one area, do not mean that this will continue to be the case. Alternatively, known seismic ‘hot-spots’ may not always be prone to seismicity. Experience suggests that large seismic events can destress large areas, leaving the ore in a better condition to be mined (stressed, but loosened). Given the consequences of being wrong, however, a conservative approach is always prudent. Understandably, large events make managers, planners and engineers nervous, sometimes to the point of closing mines down. They rightly ask ‘why, and if so how, could mining still be TABLE 3 Mt Charlotte’s 14 point Seismic Risk Reduction Strategy. Note that some items overlap or differ in intent, to suit various scenarios and mining stages across the mine. Seismic strategy item
Summary
1. Backfill
Use to add confinement, control loosening
2. Stiffness
Keep local stope stiffness low, regional stiffness high. Avoid irregular lumps and mining fronts. Faults reduce stiffness.
3. Access
Do not mine accesses along faults. Avoid driving along any major structure.
4. Unlock Faults Quickly
Start stoping at faults so it can move early. Intersect structures early
5. West-Dippers
Avoid undercutting or overcutting west-dipping structures.
6. Stope End Abutment Access
Orient abutment development E-W
7. Stress Shadow The Faults
Unclamp faults and encourage gradual movement.
8. Abutment Stress
Use narrow E-W stope abutments. Avoid increased shear stress on clamped fault
9. Pillars
Avoid diminishing pillars. Avoid stress increases in and around pillars.
10. Blasts
Small blasts generally cause small stress changes, and are associated with small energy releases
11. Preconditioning
Intentionally weaken ground to reduce its ability to carry stress, and encourage movement.
12. Destressing
Redirect locally high stresses
13. Blast Timing
Die-down exclusion time of two hours.
14. Reoccurrence
New events are less likely in ground destressed by previous large events.
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done safely and profitably at Mt Charlotte?’ But knowledge gained has allowed more thorough review of designs, work practices and risk. The mine has survived, although sometimes at the cost of abandoning high seismic risk stopes.
Preconditioning and destressing Preconditioning is the intentional creation of a weakened zone, say by blasting, in a future pillar well ahead of mining. The concept arose from observations of previous rib and crown pillar behaviour in G Block. Some strong and unstructured pillars failed violently, while others with weakening structure yielded gradually. The I1 rib pillar was preconditioning in 1994 (Mikula, Lee and Guilfoyle, 1995). A 5 m wide zone across the future I1 rib was tight blasted. Stress and displacement monitoring showed that this locally achieved a reduction in the rock mass elastic modulus from 60 GPa to 12 GPa. However the zone began to lock up and again transmit significant load as the pillar was created when I2 stope was mined. During this time, seismicity within the rib was uncharacteristically rare, and was confined to the outer perimeter of the rib. Destressing, by sequentially mining narrow slots on a sub-level scale, is an option for managing perceived seismic risk. The slots are kept full of fired ore until the adjacent area is mined. Destressing slots may be mined in highly stressed abutments, eg a sublevel ahead of the upper levels SLC caving front, or between a stope and a fault.
Empirical performance data A wealth of very useful empirical design information has been compiled at Mt Charlotte by simply collecting and plotting mine experience. This data cannot be obtained using numerical models:
• Stope span versus height of arch in stope backs (Figure 18). The arch is formed by either overbreak to equilibrium, or by intentional blasting.
• Optimum drawpoint pillar width (Figure 19). • Matthews Stability chart (Figure 20). Rating 1 (no deterioration) Rating 2 (localised spalling/deterioration) Rating 3 (significant areas of wall slabbing) Rating 4 (slabbing affecting whole pillar) Rating 5 (complete pillar failure) ADOPTED DESIGN LINE
0 Depth (m BD)
• Management knew the location and size of all significant
200 400 600 800
1000 0.0
0.5 1.0 1.5 2.0 2.5 3.0 Width / height ratio of pillar
3.5
FIG 19 - Drawpoint pillar behaviour chart. The design line represents a conservative Rating 3, at which the pillars will be serviceable after a seismic hit in the area.
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Stable Transitio n
Stability Number, N
1000
Failure M ajo r Failure
P (fail) < 5 %
100
Dilution and overbreak
10
1 P (fail) 15% P (fail) 75 %
0.1 0
10 20 30 Shape Factor, S, (m )
FIG 20 - Matthews Stability Chart using Mt Charlotte data.
Monitoring Stope design and excavation may be done well at Mt Charlotte, but the feedback loop is only complete when the rock masses reaction to mining is known. A wide range of observations, records and instrumentation has been used at Mt Charlotte:
• • • • • • • •
visual observations are kept on sets of working plans; an extensive photographic record exists; stress change monitors (CSIRO HI cells);
Mt Charlotte’s large stopes are subject to dilution when ore is drawn under waste rock fill or from wall and back overbreak. Most dilution occurs as mucking is approaching completion, and is therefore of little consequence. Mucking ceases when the drawpoint grade drops to 2.5 g/t and depending on the original stope grade, this can allow for up to 50 per cent dilution. Large drawpoints and equipment allow large blocky overbreak to be removed. Very large boulders are, however, dealt with by either drilling and popping at the brow, bombing with bags of ANFO, or impacting with a projectile cannon.
Audits Mt Charlotte has always engaged external consultants to independently assess stope designs and to help management make the best possible decisions. In mining engineering, two or more heads have always been better than one. One outcome of external audits has been regular reporting and record keeping. They have been invaluable in the face of staff turnover, to prevent new engineers from repeating previous mistakes, and as records of mining and conditions in the old upper levels.
Sustainable production A recent benchmarking study by Australian Mining Consultants (McCarthy, 1999) into sustainable production levels, suggests that mines tend to follow the following cycle:
• A decision is made to increase production beyond what the
extensometers (rods, RWEs and URWEs);
ore reserves (tonnes per vertical metre) or the development rate can sustain. This decision may be triggered by a marginal analysis (eg hoisting or milling more ore because excess capacity is available, or gaining efficiencies of scale).
blast vibration monitoring; probe holes; seismic monitoring; and stope surveys using the cavity monitoring system (CMS).
Sometimes monitoring needs to look ahead to develop a knowledge base beyond the immediate mining requirements. Towards this objective, an array of stress change monitors was installed at the bottom of Cassidy Shaft, well away from any mining, to study possible changes in the regional stress field (Lee, Pascoe and Mikula, 1999).
Optimum stopes Definition drilling and block modelling on a scale of 10 m is appropriate for the Charlotte orebodies. This is due to two orebody attributes: the irregular poddy distribution of grade, and the gradual grade change towards a resource boundary (except where a fault forms the ore limit). As a result, practical stope boundaries are found to include significant amounts of higher and lower grade material. This means that ±5 m variations in actual stope wall positions, whether by blasthole deviation or by overbreak, involve mineralised tonnes, which can be mucked without much effect on the overall head grade. In the pre-computer age, planning ran two to three years ahead of mining. All mine plans were hand-drawn, and each block was fully designed before any development was mined. In about 1985
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a deliberate move was made towards applying off-the-shelf computer technology for core logging, ore reserves modelling and stope design. Mt Charlotte avoided the temptation to modify software to better suit perceived requirements. Optimum stope designs are found via a conventional iterative and evaluation process. VULCAN 3D visualisation software is used and includes input from geology, mine planning and rock mechanics. Both cut-off grade and block geometry are optimised. Geometry influences both mining efficiencies and rock stability.
• The grade drops and profitability decreases. • A trigger (metal price, exchange rate, costs) makes the mine unprofitable.
• This drives more economies of scale and/or operation, and worse profitability.
• Closure. Sustainable production level should always be within the limits of the physical orebody (size and grade), have flexibility for the lean times, and aim to be acceptably profitable even when the worst-case assumptions eventuate. Mt Charlotte has been subjected to the above pressures. Production rates up to 2.4 Mtpa have been considered. Figure 4 shows the loss in head grade accompanying increase in production.
Blast design and scheduling Blast designs were initially constrained by equipment. The 51 mm diameter holes used in 1962 were progressively replaced by 89, 115 and 165 mm blastholes for main rings. However, the 165 mm holes caused too much damage to the adjacent rock mass and vibrations in Kalgoorlie were considered excessive. A
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BULK LOW-GRADE MINING AT MOUNT CHARLOTTE MINE
reduction to 140 mm holes and 25 per cent lower powder factor were used from G Block onwards. The present LHOS blast parameters are:
• powder factor = 0.35 kg/t in the slot, 0.28 kg/t for main rings; • 35 t/drill m LHOS, 140 mm blastholes, 4 m burden and 6 m toe spacing; and
• 11 t/drill m for trough undercuts, 76 mm blastholes. The SLC blast parameters are:
• 8 t/drill m for remnants, and • 12 t/drill m for SLC. Blast holes have been lost due to the closure of highly stressed blastholes, ground movement or fall-off before firing. For pillars and stope extensions, blast holes were dumped slightly towards the stope, so that overbreak would only affect the ends of the blast holes. For crowns, the solution was to drill horizontally from two opposing positions using horidiam raises, then failure from the centre of stope backs did not matter, as the blasthole collars were still accessible. Because the mine had to warn the town of any large firings, the scheduling of mass blast drilling and charging needed to be very well planned. A first at Mt Charlotte was the bulk delivery of emulsion explosives underground. ICI developed a unit that could do this using only Cassidy shaft cage access. This made short work of the charge-up process for large mass blasts.
Horidiam blasting method Horidiam blasting is a way of drilling and firing crown pillars from raises. It uses standard components (stage, canopy, drilling rig - Figure 21), but requires safe procedures for getting in/out of the raise and shifting the rig to the next fan of holes. In essence, it is like an Alimack raise climber with a drill rig on top.
The original Mt Charlotte horidiam rig drilled 57 mm holes, but in about 1981 an in-hole-hammer horidiam rig was built drilling 115 mm holes. The only other Australian mines that have used the technique, as far as is known, have been Mt Isa (for ventilation shafts) and Mt Lyell (similar application to Mt Charlotte). The design and construction of the first ITH horidiam drilling rig, by Atlas Copco, for Mt Lyell was based on the Mt Charlotte rig (Usher, 1992; Weston, 1998). Horidiam drilling could easily be automated, with the driller operating the rig from a crib room. Rod handling, alignment and shifting to the next set-up would be automatic. Development of such a rig will, however, largely depend on informed mine planning engineers and an enthusiastic manufacturer.
Mucking and trucking A key success at Mt Charlotte has been the ability to move large quantities of material at low-cost. Initially the mine used LHD units for the cut and fill mining phase, but for COB B, conventional high profile open cut trucks and front end loaders were used. These proved more than adequate for the job, and delivered low-cost per tonne. Trends towards larger equipment continued over the years:
• 32 tonne 769B trucks in 1972, • 10 tonne CAT 988B loader in 1978, and • 45 tonne Kiruna 501 trucks in 1981. The 45 tonne truck was a rare sight underground even five years ago, let alone in 1981. The large equipment necessitated larger access drive dimensions, which increased as the mine deepened. Several simulation studies were done to optimise the fleet. Preventive maintenance was a priority and it was done underground in the 14 Level workshop. This enabled the mine to maintain old equipment in serviceable condition, ie capital cost written off, but still delivering tonnes or drill metres at very low cost.
Orepasses and hoisting shaft A bold decision in 1978 approved the construction of the new Cassidy Shaft, at $43 million in 1981 dollars, despite doubts about the gold price and the reserves at depth. The decision was vindicated when by January 1980, the gold price was US$850/oz. This generated enough profit to pay for the shaft and the general refurbishment of the mine. Construction commenced in August 1981, the shaft was commissioned in May 1985, and it reached full production in November 1986. Key data:
• 1184 m deep, circular at 6.5 m diameter (to handle the high stress field and blast vibrations);
• • • •
FIG 21 - The horidiam rig.
For some orebodies it can mean greater sublevel intervals, less sublevel development and lower costs per tonne. It is an ideal method for remotely accessing, drilling and wrecking highly stressed crown pillars, which is its application at Mt Charlotte. Its main disadvantages are the cramped working space, noise, and the perceived higher risk of working in confined spaces and over vertical openings. The main problems with general industry acceptance are ignorance, and the ease of designing conventional blast holes for conventional equipment.
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Koepe winder, ground mounted, 12.5 m/s hoist speed; 13.5 t skips, 500 t/hr capacity; Design production capacity 1.3 Mtpa; and A hoisting cost of $0.73 per tonne achieved.
Along with the new Cassidy Shaft, a large underground crusher and orepass system was installed. The crusher delivered 150 mm ore for hoisting, and a secondary crusher on surface produced a 10 mm size ready for milling. The crusher on 36 Level, at a depth of 1100 m, was in a high stress field. Backs and walls fractured to a depth of 2 m. Support was a combination of bolts, mesh and shotcrete. To allow reinforcement placement concurrent with development and before full spans were exposed, excavation was by a multi-lift benching method. Two raises were mined to the top of the chamber, which was then gradually silled out and secured. Blastholes tended to deform in the stress field.
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Average Risk (1 to 25 = highest to low est) 0 5 10 15 20 25
1 5
20
9 13
10
17 21
0 25 Jan-91 Jan-93 Jan-95 Jan-97 Jan-99 Jan-01
Top 10 Average Risk Rating
LTI and LTI Freq Rate
30
LTI Frequency Rate LTIs per MT production Top 10 Average Risk
FIG 22 - Risk and Lost Time Injury (LTI) data. The ‘Top 10 risk trend’ is the average risk rating of the ten worst accidents reported each month.
Ground control & seismicity Run or rush of material Working at heights Explosives and blasting Falling objects (not rockfall) Ventilation, fires, fuming Operating fixed plant Operating other mobile equipment Operating loader Housekeeping M anual handling Operating truck Trip and fall M aintenance of equipment Running services
The main ore and mullock passes were constructed as a series of offset raisebored shafts, linked by short transfer rises on the levels. Under impact of free-falling ore the transfers quickly eroded and overbroke, and pass walls suffered high stress breakout, tensile wall slabbing and abrasion. Over the years, parts of the orepass enlarged to more than 40 m by 25 m in section. There were fears that the enlarging pass could affect the shaft. Hang-ups, caused by large tensile slabs, were common until the pass was kept full. In hindsight, the pass and crusher system was located too close to the Cassidy Shaft; very little room was left to excavate replacement passes.
Decline access The original internal decline was a product of the mining system. The dream of a connection to the surface led to construction in 1997 of the 3 km Sam Pearce Decline. Like Cassidy Shaft 12 years earlier, the Sam Pearce resulted in a significant extension of mine life. It became feasible to bypass the orepass system and rework old COB drawpoints.
Waste water Waste water in the mine was efficiently managed using a series of dams and pump stations on 39, 32, 28, 20, and 9 Levels. This kept lifts shorter, and provided water storage capacity during pump maintenance.
As the mine is located in the suburbs of Kalgoorlie, the workforce tended to be local, more stable and settled. There was always a core of older workers, some of whom had worked at Mt Charlotte many years. With training and experience, much was achieved; the maintenance people were able to change all four winder ropes in only 18 hours, or get impossibly large equipment underground via the shaft. Few contractors were used. The mine operated on three shifts × six days until September 1996, when a 24-hour roster was introduced. Workforce numbers were:
• In excess of 500 when mining commenced in 1962. • 360 from December 1976 after the near-closure. • 200 to 220 from 1982 to 1997. During this period production increased from 0.8 to 1.6 Mtpa.
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Ergonomics, noise, hygiene Total All Categories
FIG 23 - Accident risk profile of the mine, based on 1003 accident reports since December 1997.
Safety initiatives Safety initiatives at the mine have included defined operating procedures for major tasks, and a Permit system requiring Manager authorisation before critical tasks can be commenced. In 1993, 400 000 manhours were achieved without a Lost Time Injury (LTI). LTIs normalised to tonnes production have shown a downward trend (Figure 22). In a watershed change in 1997, the accountability for mine safety was transferred from the Safety and Training Department to line management. A new accident reporting system was implemented, and three groups comprising engineers and staff were formed and trained to run the system (Mikula, 1999c):
• Accident Investigation Review Group, to monitor accident reports, do investigations, and management and other groups.
provide
feedback
to
• Safe Work Procedures Group, to use operators to develop Job Hazard Analyses and Safe Work Procedures, and to update the Procedures in line with Accident feedback.
Workforce
• 150 after cost pressures in 1998. • 80 after further cost pressures in 1999.
Communications and signage
• Monthly Inspection Review Group, to oversee workplace inspections with emphasis directed by trends in accident reports. Line Management took the lead in reporting accidents, and a no-blame policy is used. A Probability/Consequence Matrix is used to numerically rate the risk level of an accident from 1 to 25. Accident reports jumped 2000 per cent from five per month to over 100 per month in two years. The benefits of this system are:
• a database of the risk profile of the mine is now available (Figure 23);
• trends in accidents are immediately apparent, and are rapidly addressed, which reduces the chance of a major accident occurring; and
• the risk level of the mine has been measurably decreasing over time as high-risk situations have been addressed (Figure 22).
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BULK LOW-GRADE MINING AT MOUNT CHARLOTTE MINE
CONCLUSIONS Mt Charlotte was originally the poor relation of the southern Golden Mile mines, but it eventually emerged as a large producer, and the only one to survive the 1976 cost crisis. It has been an efficient, low-cost, large tonnage, underground hard rock mine, with a history of the application of new technology, innovation and continual improvement of mining methods. All have contributed to its profitability, long life and efficient use of finite ore reserves. Key factors in Mt Charlotte’s success have been appropriately timed and well-considered management decisions, while keeping costs under control. Perhaps the best summary of the Mt Charlotte philosophy is: make it simple, make it big and be sure!
ACKNOWLEDGEMENTS The authors wish to express their appreciation to ALL those who ‘sailed in her’. Special thanks to the following for their assistance in the preparation of this paper: Dion Fotakis, Ian Grljusich, Kevin Malaxos, John Sztermula, and of course to KCGM for allowing us to tell a small bit of the Mt Charlotte story.
REFERENCES Bamford, W E, 1971. Stresses Induced by Mining Operations at Mount Charlotte, in 1st Aust NZ Conf Geomechanics, pp 61-66. Bridges, M C, 1982. Review of Open Stope Mining, AMIRA Sponsored Project Report #81/P138, January. Bischoff, K and Morley, C, 1993. Geology, resources definition and reserve estimation at Mt Charlotte, Kalgoorlie, WA, in Proceedings International Mining Geology Conference, pp 1-17 (The Australasian Institute of Mining and Metallurgy: Melbourne).
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Brodie-Hall, L, 1988. The Mt Charlotte Story, in Research & Development for the Minerals Industry. WA School of Mines, pp 153-157. Lee, M F, Beer, G and Windsor, C R, 1990. Interaction of stopes, stresses and geologic structure at the Mt Charlotte mine, WA, in Rockbursts and Seismicity in Mines (Ed: Fairhurst) pp 337-343 (Balkema). Lee, M F, Pascoe, M J and Mikula, P A, 1999. Stress Fields in Western Australia, A Kambalda - Kalgoorlie Case Study, In Workshop on Mining in High Stress and Seismically Active Conditions, Perth, 15-16 April. Australian Centre for Geomechanics McCarthy, P, 1999. The Temptation of Tonnage, Australian Mining Consultants Pty Ltd letter to the Mining Industry, October. Mikula, P A, Lee, M F and Guilfolye, K, 1995. Preconditioning a large pillar at Mt Charlotte mine, in Proceedings Underground Operators Conference, pp 265-272 (The Australasian Institute of Mining and Metallurgy: Melbourne). Mikula, P A, 1999a. Choices in ground reinforcement at Mt Charlotte mine, in Rock Support and Reinforcement Practice in Mining, Kalgoorlie, pp 285-293. (Balkema). Mikula, P A, 1999b. Four years of seismic monitoring at Mt Charlotte gold mine, in Mine Seismicity and Rockburst Risk Management in Underground Mines, Australian Centre for Geomechanics, Perth. Mikula, P A, 1999c. Development of a geotechnical risk control system at Mt Charlotte mine, in Mining in High Stress and Seismically Active Conditions, Australian Centre for Geomechanics, Perth. Reed, M, 1982. Mining at Mt Charlotte – Current Practice, in W.M.C. Underground Operator’s Conference. Usher, R E and Kennewell, G J, 1992. Evolution of Mining and Current Practices in the Prince Lyell Orebody, Mount Lyell Mining and Railway Company Limited, in Proceedings Fifth Underground Operators Conference, pp 37-46 (The Australasian Institute of Mining and Metallurgy: Melbourne). Weston, A J, 1998. Re-opening of Mining at Mt Lyell, in Proceedings Seventh Underground Operators Conference, pp 7-14 (The Australasian Institute of Mining and Metallurgy: Melbourne).
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The Mine Management System at Olympic Dam Mine S Youds1 ABSTRACT The operations management system at Olympic Dam has evolved over a two-year period to the state described in this paper. This evolution has been driven largely by the requirements of the recent expansion of the Olympic Dam operation from three to nine million tonnes of ore per year. Such a change process demands an holistic approach to mine management and several techniques have been employed including XPAC, Mine Works Planner as well as a basic philosophy of ‘just-in-time’ completion. The philosophies, software and tracking mechanisms for managing a large tonnage operation are described, outlining a fully integrated scheduling system from shift to shift detailed activity planning to a 20 year production plan. Moreover, the vision of production exceeding the current limit of nine million tonnes means that there are to be further developments to this state-of-the-art system in use at the Olympic Dam mine.
INTRODUCTION Scheduling is a mystery to some, dreaded by others and a time consuming but necessary task to others. It is generally the responsibility of a senior mining engineer. Usually because of the length of time necessary and the shortness of time available to effect a reliable schedule the task is often done in a superficial way. The methods and tools used will depend on the size, complexity and tradition of the mine. All agree it is a vital part of any process … without time related goals there is no achievement and no reward. Everyone involved feels the need to improve the process used but doesn’t know where to start or what tools are available. The traditional boundaries of short-, medium- and long-term schedules are counter-productive naming conventions for describing the process of scheduling when assessing the function as a whole. There is a requirement for a seamless transition between the three traditional components for an effective system. Defining the scheduling boundaries to ease the task of managing the distribution of the work and the positioning of personnel on hierarchical structures sets up barriers to communication. The use of the traditional long-, medium- and short-term definitions is not used in this article however the various parts of the system do loosely fit into these categories. It is the interaction between the roles that is important to keep the whole function on an informal and mutually trusting basis. The longer term less detailed schedules should be seen as help or a starting point to the shorter term plans not a hindrance. The system at Olympic Dam is still developing, however the shorter term schedules have taken shape. The basic starting point is the use of a five-year schedule for the production of ore from stopes. This details the tonnage and grade of the multi-metal ore from individual stopes. Various constraints of ventilation and ore handling are used to reflect the practical restrictions to production at the mine. At present this function is performed on a spreadsheet. It is anticipated that this five-year schedule and a 20-year plan will utilise XPAC as a tool for producing the longer term plans. The main gains to be had with this software tool are the auto scheduling functionality combined with the speed with which various scenarios can be produced. The site-specific tailor-made constraints and priorities are altered to give different 1.
Olympic Dam Scheduling Engineer 1997-98, WMC Olympic Dam, PO Box 150, Roxby Downs SA 5725. Email: [email protected]
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results based on the various what-if scenarios that senior management deem as likely alternatives. The reliability of the results is dependent on the ability of site personnel to model the actual scheduling mechanism in the software. This is currently in the development stage. If it is successful it will give the management team at OD a powerful long-term planning tool which will enable them to make the strategic decisions necessary to maximise the value of this orebody to the company and to the shareholders. The five-year production schedule is used as the building block for developing the associated five-year development and drilling schedules. Mine Works Planner (MWP), a project management type package is used for this purpose. The output from Mine Works Planner is used to develop the shorter term three month, monthly rolling, schedules (the Forecast). These shorter term functions are carried out by the operations personnel as opposed to the longer term work, which is a function of the Technical Services Dept at OD. The three month schedules are further broken down into 12-day plans, which are updated every two days … a 12-day, two-day rolling, schedule. The mine at Olympic Dam (OD) that the management system described here, is required to run has the following details and variables. The mining method used is Long Hole Open Stoping (LHOS) in granite hematite breccia ore, hosted in granite country rock. The stopes are back-filled with materials from cemented aggregate fill (CAF) of varying strengths to rockfill from both underground and the surface. The various mining areas are divided into discrete ventilation districts exhausted by one pass airflow using centrifugal primary fans on the surface. The ventilation district production rates are directly proportional to the amount of ventilation available. The material extracted from a stope in production ranges from 30 000 t to 75 000 t per month depending on the use of truck/loader or loader/train ore haulage respectively. The geotechnical restraints on stope size dictate small plan areas, typically of 35 by 45 m and no vertical height limit, creating stopes with skyscraper-like dimensions. The quality constraints on production include smoothing the grade for copper, gold, silver and uranium. The smelting through put rates are dictated by the level of the copper:sulphur ratio in the ore. The effectiveness of the autogenous mill is determined by the iron silica ratio and the size grading of the material hoisted from the mine.
FEATURES OF THE OLYMPIC DAM SYSTEM The scheduling function at Olympic Dam evolved over two years from a monthly forecast of relevant activities in the 3 Mt/a mine to the scheduling system now being used for the expansion and operation of a 9 - 10.5 Mt/a mine. The scope of this paper is to detail the system and software used and to explain the philosophies developed which form the core of the system. Figure 1 demonstrates the flow of information from the generation of the 20-year schedule through to the shift by shift management schedules. While adjustments are made to the system, the basic philosophies have remained unchanged. The use of rolling schedules is a feature of the scheduling system at OD. By rolling, it is meant that a schedule, for example covers a period of three months, each month a new schedule is produced with another month added. This maintains the three-month look-ahead schedule. The period that is planned for looks further ahead than the immediate time frame required.
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OLYMPIC DAM MINE SCHEDULE DATA FLOW
ORE RESERVE 20 YEAR PRODUCTION SCHEDULE PHYSCIAL AND ECONOMIC LIMITATIONS
(X-PAC)
5 YEAR ACTIVITY SCHEDULE (Mine Works Planner)
THREE MONTH ROLLING SCHEDULE Physicals by day
12 Day, 2 daily Rolling Schedule Detail Physicals by Sub shift divisions i.e (crib breaks) FIG 1 - Flowchart of scheduling hierarchy.
Therefore any serious problems identified in the next three months are known, at worst with a minimum of two months notice. This extends the reaction time for planning around the issues, leading to less pressured decisions and better solutions. Two months has been found to be sufficient time to react to issues which aren’t in the scope of the five year schedules, but nonetheless have significant potential impact on production. A further extension of this ‘reaction’ period would give further reduction in the reliability of predictions with questionable benefit to operations. The ratio of period of coverage to period of update would depend on: 1.
the degree of volatility in the change environment that the mine operates under;
2.
the actual time units, ie days, months or years of the individual schedule; and
3.
composition of the schedule, ie what activities are defined in the schedule.
The various time frames would be adjusted based on the effectiveness of the individual schedules in assisting in the coordination and planning of operations. One of the important tasks in finding a suitable scheduling system is defining the balance between detail and scope or time frame. For example the shortest system at OD has a 12-day look ahead short-term schedule, this is updated every two days, hence a 12-day, two day rolling schedule. The detail used in this schedule is to the level required for on shift decisions. For
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example fan and hoist outages or portal access restrictions would be included. The effect of stope firings and preparation work would be included to allow for any affected activities to be re-scheduled. The stope firings required for the month have been included in the three month, monthly rolling schedule as required by the production bogging plan. However the exact timings will not have been detailed, as this level of detail at the three monthly level is superfluous and likely to be inaccurate. The nature of the scheduling role is one that requires an intimate knowledge of the whole operation. This presents continuity problems when individuals move on as part of their career progression. A system of over lapping or rotating the various engineers involved at the operation through the function, on for example a three-monthly cycle would be a way of overcoming this issue. This has not been resolved as yet at OD. One part of the system at OD that has resulted in significant gains in the structure and organisation of the schedules has been the use of a project type approach to preparing the individual stopes for production. This means that all work that is defined in the schedules is assigned to the stope that it is required for (Figure 2). One result of this improvement has been the gains in communication to all levels in the hierarchy. The basis of the activities has a structure and a purpose for everyone to appreciate. This makes the policing of the schedules a largely self-driven exercise. By assigning activities to stopes, in the cases when the connection is tenuous it gives the individual supervisors some time frame and context to work by. Another
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THE MINE MANAGEMENT SYSTEM AT OLYMPIC DAM MINE
Stope : Purple 12 11 April 2000
Capital Dev’t Metres Operating Dev’t Capital Dev’t Access Dev’t
FIG 2 - MWP Single stope, ALAP constraint and inter activity links. (Refer to CD-Rom for colour explanation.)
gain is the use of the stope naming and numbering system in the code hierarchy in the scheduling package itself. This results in the filtering and sorting of data being in a format which is understood at all levels. This is not the advice of project schedulers who avoid logic to the numbering system for flexibility and simplicity. Owing to the number of people using the outputs and adjusting the schedules, in an on-going mining schedule, the positives are greater using the stope-based (logic driven) naming system. In traditional project scheduling it is generally only system aware individuals who adjust and use the schedules so a non-user friendly system is appropriate. In addition to this the variety of activities without logical grouping make the use of a logic driven activity numbering system unattractive. The use of a Just-in-Time (JIT) philosophy is a progression from the previous point. Once all the activities have a time driven basis, (ie they have a connection to a stope in the production schedule) they can be organised for completion in a timely manner for production from that stope. Using the JIT thinking as an underlying theme of the scheduling function has been one of the building blocks of the scheduling system at OD. It will be clear to most professionals working at mines that the time value of money and the economic sense of spending development dollars as late as possible without impacting on production are areas for potential improvement. At the same time mining individuals get very nervous if development is left to the last minute, as the maxim ‘If it can go wrong, it will go wrong’ applies very aptly to underground mining. All activities are defined by their relationship to stopes due to come into production. Using this approach it is possible, once the development and drilling required for a stope has been defined
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to, using comfortable advance or drilling rates, time the activities to occur at the latest time before production starts. The importance of making all members of the team from the development teams to the maintenance personnel aware of the impact of not achieving targets is vital. If anything it is this JIT philosophy that has resulted in the ability of the mine to be comfortably able to achieve the ambitious targets of the expansion. However as with the chicken and the egg which came first, the culture change to a belief in schedules or the awareness of the need to achieve targets. In practice the two developed simultaneously, each assisting the other to change the culture of the whole team at OD mine, one without the other would not have worked.
SOFTWARE USED XPAC The use of this scheduling package is still in the development phase. The key aspect of this software is the auto-scheduler. One of the most time consuming tasks at OD during the expansion has been the necessity for a number of five-year production schedules to be developed due to the expansion related change environment. The existing Excel spreadsheet driven system had swelled to a ten mega-byte plus monster over the years. As a result these variations took some time to complete and as scheduling personnel will have experienced, the senior personnel awaiting the answers generally want them yesterday. The need for a system to run a number of what-if scenarios in an expanding mine is a significant assistance to the strategic decisions required (Figure 3).
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Geological Reserve Database
XPAC In-Situ Reserve Database
Data Manipulation
Procedures Based On User-Defined Logic
XPAC Mining Reserve Database
Additional Data
Product targets & specifications. Sequencing rules & constraints.
20 YEAR SCHEDULE DATA FLOW
Second Adjacency Macro
AutoScheduler Module
20 YEAR Production Schedules
FIG 3 - XPAC diagram of information flows.
A predictive tool that runs variations with preset achievement hurdles in a short time adds another dimension to the planning and management processes at a large mine. It is envisaged that this package will be used to choose the stopes to mine and back-fill to achieve the required grades of copper, uranium, copper:sulphur ratios, gold and silver, given the target hoist tonnes. The constraints of ventilation and ore handling can be factored in. This customer definition of the outcomes will be used to generate a number of variations by changing the priorities and restrictions. It is this aspect of generating ‘blue sky what-if’ scenarios which allows true optimisation of the mining and treatment of an orebody. In itself XPAC is not an optimisation tool, however by changing the various priorities the best NPV or profit over mine life or profit at x production rate, etc can be investigated. Up until this point, the use of general commonsense rules have been used to produce an ‘optimal’ schedule. For example stopes are arranged in mining sequence to maximise rockfill as opposed to cemented aggregate fill (CAF), therefore the largest stope in a series is taken last. However although rockfill is obviously cheaper than CAF, it has not been substantially proven that this is the best economic course of action in all cases. There are other economic considerations which have not been factored into the evaluation. For example it may be economically beneficial to mine the higher grade stopes first in order to have an overall NPV benefit from the reserve due to the time value of money and long-term commodity price trends. The current system is not flexible to changes in strategy, XPAC on the other hand can be adjusted and an answer to a ‘what if’ given in a few minutes. The move to the use of the auto-scheduling tool is analogous to changing from an empirical calculation system to a quantitative calculation system. Any factors, which are deemed to have an effect on the bottom line, can be included.
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The central function of XPAC, the auto scheduling system is a complex mathematical system, which selects stopes to mine on the basis of the rules, priorities and restrictions that have been set up at the introduction of the software to an operation. These are adjusted as the operators of the mine gain a greater understanding of both, how the software operates and of the implications it has for their orebody. For the first time the personnel involved in scheduling will have the time to assess the impact of various courses of action. Usually they are at a disadvantage with regard to the production of schedules, so the first version is accepted. The auto scheduling functionality allows ways of mining the orebody, which are novel and innovative to be assessed in a quantitative and hence comparative fashion. The value of vertically integrating XPAC by extending the detail to give short-term schedules like those produced with Mine Works Planner (MWP) has yet to be assessed. The complexity of the rules and restrictions required for XPAC to duplicate MWP’s role makes this option currently unworkable. The difficulty in defining the rules of mining order for stoping and back-filling have proven difficult enough without increasing the complexity by an order of magnitude with the addition of individual stope activities and the associated resources (power, services, machinery, ventilation). From the output point of view, this program starts from the traditional finishing point and works backwards. The tonnes of metal that are required for the project’s rate of return or downstream customers can be used as a starting point ie it is given as one of the requirements that the auto scheduler has to fulfill. This package will give an answer as close to this point given the other constraints under which the mine operates. Obviously the less restrictive and prescriptive and therefore more
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period depends on the size of the stope, as drill metres are directly proportional to tonnage. As shown in Figure 2, the development is then assigned ‘As Late As Possible’ (ALAP) constraint and linked to the front of the raise boring activity in this way it occurs only as needed for production. In this way the development is completed, as it is needed rather than months in advance. It is the identifying of individual physical rate driven activities, which allow a resource driven schedule to work. It is this detailed aspect which makes it a powerful planning and budgetary tool. The monitoring of progress according to the schedule is then a relatively straightforward process as each activity has a time-frame in which it should occur. In addition the coordination of other activities and smoothing of the use of scarce resources such as power, ventilation and water can be attempted which is a substantial advance in managing the underground environment. For this system to work effectively, each development drive (slot, drill drive or access) has to be defined and the order that they will be mined (based on logical access and fresh air) has to be defined. The links between activities that are shown on Figure 2 as lines between activities, are used for this purpose and are generic with this Gantt chart-type program. All individual drilling activities have to be defined, up-holes, down-holes and raise-boring. Defining an activity includes inputting the physicals and the resource the activity utilises eg drill metres, ventilation exhaust fan, etc. Each type of activity (development, down-hole drilling, etc) has a predefined resource rate. The links firstly between activities of the same stope allows the smooth non interactive preparation of a stope for production. In this way it is possible to ensure no unsafe interaction of activities on different drill levels. Secondly the links between activities of the same resource allow the use of the various underground machinery to be scheduled (Figure 4). It is the ‘same stope’ links that allow the
generic the rules, the greater the span of possible answers and the greater use this tool can be put to as a system for reviewing the basic concepts and paradigms used in the mining of individual orebodies.
Mine works planner This mine specific software package was developed in South Africa. It is a Gantt chart type project management tool. In the system at OD it has fitted neatly into the medium-term scheduling time-frame. Although it could do a number of functions at OD it is used to identify the activities and resources required to meet the production requirements. These production requirements are defined in the five-year production schedule (XPAC output). Five-year development, raise boring and production drilling schedules defining the timing of the activities and the level of resources are the output that MWP supplies to operations personnel. They are constructed using the stope start times as defined in the production schedule as starting points. The various activities required for each stope to come on line are defined as to the physicals (development metres and drill metres) required at each sublevel (ie separate work area). The physical-based rate functionality of the software is then employed to identify the period needed to prepare the stope, ie working back from the stope start time to allow sufficient time to start the stope as planned. As mentioned before, all activities are related back to the start time of the stope, which requires the activity be complete for production. The code nomenclature used in the data tables uses this fact to filter and sort activities. In practice, the JIT organisation of work is achieved by fixing the raise boring function at a point in time, which will allow the production drilling to occur comfortably. The length of this
Downhole Drilling 11 April 2000
Rate
Desc.
Name
ODpul3052F/ 52 PURPLE 130 Downholes 300pdm/d ODye02429B/ 29 YELLOW 24 Downholes 300pdm/d ODye02426B/26 YELLOW 24 Downholes 300pdm/d ODor03742D 42 ORANGE Upholes
300pdm/d
ODsc00926D 26 SCARLET Downholes
300pdm/d
ODam00446F46 AMBER 4 Upholes
300pdm/d
ODpu4546F 46 PURPLE 45 d/holes
300pdm/d
ODin00000C JUMBO 18 REBUILD ODin00000C JUMBO 19 REBUILD
Total Dev Metres
1141
1017
877
695
102mm Drill Metres Uphole Drill Metres
14106
37211
48845
58425
10228
18120
24509
9456
OD R/Bore Metres
213
300
396
497
RESOURCES Task Critical Summary
Summary
Milestone
Percent Complete
Baseline
Critical Task Links
-
C/BLT
PDDRL
DMDRL
PBOG
T4DRL
BFIL
RBORE
DEVEL
CBDRL
PUDRL
PC/UP
RBOG
B/C
RFIL
PASS
CONST
FIG 4 - Same resource MWP schedule for scheduling of machinery. (Refer to the CD-Rom for colour explanation.)
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scheduler the freedom to forget about the generic rules of individual stope activity precedence in order to concentrate on smoothing the resource usage. MWP, as do most of the comparable project management packages, has mouse operated manipulation of the links on the Gantt chart part of the screen allowing easy manipulation of the schedule. Either target lateral development advance per day or drill metres per rig defines the resources available. If necessary the calendars used can include regular non-working time to allow an effective reflection of the rates required to achieve the necessary targets. At OD, the 24-hour continuous roster didn’t require a complicated calendar to reflect the inactive time. A set drill rate per rig to allow for rig maintenance, site moves and acceptable downtime is used currently although planned maintenance periods can be included if required. Development is approached differently. A total metre rate per day was defined, based on the capacity of the available personnel and equipment. A base heading advance rate was used in the individual stopes. The rate chosen was comfortably achievable. The JIT system based on ‘As Late As Possible’ development will indicate when a heading should be started to complete the development at the rate defined to allow sufficient time for drilling, etc before the stope starts production. This keeps the pressure off development crews to get high advance rates. Apart from the safety gains, an additional benefit is the greater number of active headings giving additional operational flexibility. One feature of using the Gantt chart function is that the scheduler can see graphically what time individual activities should be done to achieve the desired production start dates. The skill in setting up the schedule is to make all the activities adjust automatically to changes. If activities are locked in place, time wise, the flexibility of such tools is lost. Moving activities is the most time consuming part of adjusting the common spreadsheet based schedules and hence limits their ability as a management/planning tool in running ‘what-if’ scenarios. The output from MWP is transmitted to the personnel who control the various departments’ activities. They use this output as the basis for the three month, monthly rolling schedules (or monthly forecasts), which are developed towards the end of each month in conjunction with ventilation, production, services, ground control and electrical departments. The fact that the basis for the short-term schedules are already linked to avoid clashes means that co-ordinating all the activities, traditionally the onerous scheduling duty is already incorporated. The major task at this stage is to factor in the changes that have occurred since the last re-schedule. Those tasks not scheduled for in MWP and those one-off short-term events such as portal closures or VIP visits form the agenda of the ‘forecast meeting’. This meeting occurs as a communication mechanism once a month to resolve any contradictions in or changes to the five-year activity schedules (MWP) in the next three months. The output of this meeting is the three-month, monthly rolling schedule, each individual activity type being the responsibility of the relevant section engineers in operations. The one further schedule that is used is the short-term 12-day, two daily rolling schedules. This uses the output from the three-monthly schedule with localised adjustments as necessary. The content of this schedule reflects those activities that have firstly a significant impact on production and secondly highly changeable activities. These have been distilled into the following:
• production tonnes from individual stopes to reach required monthly targets;
• stope firings and preparations for firings;and • ventilation maintenance shutdowns, power outages and shaft
Each individual shift requirement is detailed for the next 12 days. Every two days a meeting to identify the next 12 days’ events and restrictions is held internal to operations personnel. MWP has no real use in this time period schedule. It could be adapted to be of use but the heavy workload required for the input would not be realised in terms of increased management control. The task is relatively straightforward from a scheduling point of view and the time saving benefits of a linked schedule would not be realised in this situation. The balance of reducing the level of maintenance required to operate a schedule into perpetuity and having a scheduling tool, which reflects current activities, is a contentious issue. The overriding factor is the need to have a schedule that works and therefore has the confidence of those who use it. As MWP is developed to include cablebolt installation, electrical resourcing, water resourcing as a part of the resource leveling process the once onerous three - four hour ‘forecast’ meeting taking up key personnel from all departments becomes just checking over the details and factoring in any localised changes.
MONITORING PROGRESS In any scheduling system, the schedules are only as good as they are achievable. Unless there is a system of monitoring the processes that are described in the schedules, and more importantly, a system which allows management of deviations in a timely fashion, there is no point in spending time developing effective reflective schedules. The systems that have been introduced were developed under the guidance and insistence of the current operations manager at OD mine, Ian Smith. It consists of two main components. The first is a weekly chart which monitors the development and drilling rates required to reach the required targets by the dates defined in the five-year schedules in MWP (Figure 5). As this rate climbs it shows to those in the chain of command where attention needs to be given to individual activities. This method has been used with success in managing the work associated with the current OD expansion. Specifically the development and construction that has been done in preparing stoping areas in a timely fashion for the 300 per cent increase in production rates. The second is a series of pie charts showing the planned and actual progress in development, drilling, production and filling of individual stopes (Figure 6). Each activity type has a segment of the pie chart that represents it, with development and then drilling around to filling. Two arrows, actual and planned, move clockwise around the ‘face’ of the chart comparing actual progress with planned progress. For ease of reading the various charts are grouped into the respective activity types in which they are planned to be. This chart is updated every month from the reconciled monthly actuals and using the scheduled activities from the five-year schedule. The major benefit of this graphical representation is that the reader does not require intimate knowledge of the operation to get a feel for how the operation stands in regard to achieving its targets. The weekly critical activity charts are the driver that the managers and supervisors of individual areas use to assess their own progress and take appropriate action. This is the main tool used to allow the operators to manage the situation themselves. The higher the rates escalate the higher the level of supervision becomes involved. The pie charts give a broader picture from which senior managers can have their attention drawn to the key areas which are being performed out of time. Activities being early or late have a far reaching detrimental effect on the scheduling function. For choice most schedulers would prefer early to late.
maintenance.
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THE MINE MANAGEMENT SYSTEM AT OLYMPIC DAM MINE
FIG 5 - Weekly critical task data showing rates needed.
CONCLUSION
Name of Stope
end
Filling
start (Date)
Development
Planned Actual
Drilling Production Production start date Time from start of development to end of filling
Total production tonnes
The scheduling system described here is a complete system from the long-term 20-year time frame to a shift by shift forecasting function. As it is still under development there is scope for further improvement and refinement. The major changes in the scheduling function have been the culture changes that have occurred. Without these none of the scheduling changes would have given the results that have been achieved. This has resulted in an effective planning system which has enabled the mine to plan and implement the smooth transition from a three million tonne a year operation to over three times that production rate in a year, an impressive achievement in an underground mine. The key components of a scheduling system are: 1.
a method of generating the activities that need to be done to achieve required target metal tonnes and a time frame by which those activities need to be done;
2.
a system by which the resources available can be assigned to those activities to accurately reflect the actual ability of the mine to achieve the targets required;
3.
a monitoring process which in a timely manner, graphically or numerically compares actual achievement against the planned scheduled activities; and
4.
a culture which realises both the importance of achieving the targets, the impact of non achievement and a belief in the use of schedules at all levels of the hierarchy.
FIG 6 - Pie chart for an individual stope showing actual against plan.
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Materials Handling at Olympic Dam Mine — Olympic Dam Aiming for ‘Gold’ P Bowman1 ABSTRACT WMC Limited, owner/operator of Olympic Dam, the world’s largest underground uranium mine and the fourth largest underground copper mine has now completed an expansion that has been classed as a benchmark in the mining industry. To achieve a production increase from 3.4 to 9.25 Mtpa from an underground mining operation in two years was a major challenge, it was essential to review the overall materials handling system from ‘the face to the stockpile’. The challenge was two-fold: a.
How to deliver 30 000 tonne of ore per day to the processing plant, from a single underground mine?
b.
To provide the processing plant with a consistent metallurgical feed. This paper traces the transformation from the old to the new materials handling system at Olympic Dam.
INTRODUCTION Olympic Dam Corporation is located 560 kilometres north of Adelaide. The operation is named after a nearby livestock watering dam which was built by a local pastoralist to commemorate the 1956 Melbourne Olympic Games. The purpose built town of Roxby Downs situated 16 kilometres to the south of the operation, retains the name of the pastoral lease on which it was established. In 1975, a WMC exploratory drill hole intersected 38 metres of one per cent copper approximately 353 metres below the surface. A further eight holes drilled in a wide grid pattern were barren. The tenth hole, drilled north of Olympic Dam intersected 178 metres of copper, uranium, gold and silver indicating a major discovery. Subsequent drilling has positioned Olympic Dam as the fifth largest copper deposit, the largest single source of uranium and the tenth largest gold deposit in the world. A joint venture was formed in 1979 between the BP Group (49 per cent) and WMC (51 per cent) to progress development of the discovery. In 1982, the Whenan Shaft was sunk into the orebody for further exploratory drilling and for extraction of bulk samples for metallurgical testwork. This work was carried out in a pilot plant built in 1983 and which demonstrated that separation of copper and uranium could be achieved by producing a flotation concentrate for smelting and refining of copper and leaching of the flotation tailings for recovery of uranium. In 1986 the joint venturers committed to Stage 1 for the development of a mine, concentrator, hydrometallurgical plant, smelter, refinery and associated infrastructure to produce approximately 45 000 t of copper, 1200 t of uranium oxide, 20 000 ozs of gold and 50 000 ozs of silver per year. These facilities were commissioned in 1988. The smelting process is a single stage direct to blister flash furnace made possible by high concentrate grades (55 per cent) and copper/sulphur ratios from mining the higher grade chalcocite/bornite parts of the orebody. 1.
Executive General Manager, WMC Copper Uranium, GPO Box 860K, Melbourne Vic 3001.
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Two ‘optimisation’ expansions of these original facilities were undertaken in 1991 and 1993 to lift production capacities to 85 000 t copper, 17 000 t uranium oxide, 30 000 ozs gold and 350 000 ozs silver per year. In 1993, WMC purchased the BP Group’s share of the Operation giving WMC 100 per cent ownership. In 1990, work started on studies into a major expansion of Olympic Dam and following conceptual, pre-feasibility, feasibility and optimisation studies, WMC announced in July 1996 that an expansion to a nominal capacity of 200 000 tpa copper would proceed. Ore mining and treatment would be lifted to approximately 9.25 Mt/a ore providing approximately 4300 tpa uranium oxide, 80 000 ozs gold and 850 000 ozs/a silver. In January 1999, WMC announced that the new copper smelter at Olympic Dam was in full operation two months ahead of schedule. And two months later, on 26 March 1999, the Prime Minister of Australia, John Howard officially opened the Olympic Dam expansion project. The Olympic Dam expansion enabled an increase in production from 85 000 t of refined copper and associated products per year, to a nominal capacity of 200 000 t per annum. By late-1999, more than 160 kilometres of underground roadway had been completed, along with the new underground rail haulage system, which is only the second of its type in operation throughout the world. Access to the mine is via a surface decline and the Whenan shaft. Whenan shaft was originally sunk as an exploration access and later upgraded for hoisting. The 4 km long service decline connecting to Whenan shaft was constructed later to accommodate the increase in service demand. The 600 metre deep Robinson shaft was developed in 1995 to cater for the increase in development tonnage, and as a part of the recent expansion a new shaft named the Sir Lindesay Clark Shaft was sunk to a depth of 860 metres. This shaft was fitted with the largest mine winder in Australia in terms of power, rated at 6.5 MW and with a hoisting capacity of 1375 tph.
GEOLOGY AND RESERVES The orebody consists of a large number of semi-discrete mineralised zones scattered throughout an area 7 km long, 4 km wide and over 1 km deep. The ore bearing zone has a 350 m thick overburden of barren sediments. Proved and probable ore reserves as of December 1999 are 605 million tonnes averaging 1.8 per cent Cu. This implies that a very long life can be expected from the Olympic Dam Operations, even at the greatly expanded production rate. Initial exploration was based on a conceptual model for the formation of sediment hosted copper deposits and modelled gravity and magnetic anomalies. Subsequent drilling has positioned Olympic Dam as the fifth largest copper deposit in terms of contained metal and the deposit as the largest single occurrence of uranium in the world today. Based on drill hole data, point estimates of grades are generated for the centres of 5 m × 5 m × 10 m blocks throughout the orebody. Estimates are generated for Cu, U3O8, Au, Ag, S and density. Once this process is completed, contours can be generated for individual metal concentrations and for total combined in situ dollar values. Based on long-term predictions of
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metal prices and exchange rates, individual metal concentrations of each block are factored and summed to give a combined dollar value. Currently the in situ dollar value contour is taken as the mining design cut-off. The average dollar value of each stope has to be greater than $60 for probable reserves and $70 to be included in the Proved ore reserves.
TABLE 2 Mine development, production and backfill statistics. Statistic
Value
Unit
770 000
Tonnes per month
1100
Metres per month
Average stope production rate
30 000
tonnes/month
Backfill
158 000
m3/month
300 000
tonnes
Total tonnes of ore hoisted Total development
TABLE 1 Olympic Dam ore reserve and resource. Tonnes (× 106)
Cu (%)
U3O8 (kg/t)
Au (g/t)
Ag (g/t)
Average stope size Stopes in production (online )
24
Stopes per month
Proved
121
2.4
0.6
0.6
4.2
10
months per stope
Probable
485
1.6
0.5
0.5
3.4
Average stope production duration
Total Reserve
605
1.8
0.5
0.5
3.6
Average backfilling duration
2
months per stope
Average curing time
3
months per stope
Category
Measured
500
1.8
0.5
0.5
3.6
Indicated
1150
1.3
0.4
0.5
2.9
Inferred
670
1.1
0.4
0.4
2.4
Total Resource
2320
1.3
0.4
0.5
2.9
ORE HANDLING THE OLYMPIC DAM EXPANSION PROJECT The nature of the orebody, being a large volume of relatively low-grade mineralisation at depth, lends itself to a strategy of optimisation and economic success through achieving economies of scale and the ability to mine and process ore at a high production rate. The completion of the expansion has enabled the Olympic Dam mine to produce approximately 9.25 Mtpa of ore yielding 200 000 tpa of cathode, 4300 tpa of uranium concentrate, 80 000 ounces of gold, and 850 000 ounces of silver. The achievement of these targets has necessitated the construction, purchasing, and implementation of a significant quantity of infrastructure including; a new ore haulage shaft, a new automated rail haulage system, together with associated ore pass, haulage level, underground crusher, and fine ore bin system. This is, of course independent and exclusive of the processing plant expansion. The total cost of the expansion project was approximately $A1940 M. The production targets for the expansion have been met by retaining the same mining method Long Hole Open Stoping (LHOS) incorporating Cemented Aggregate Fill (CAF), however the sublevel intervals have been increased from 40 to 60 metres, and a general trend to larger stopes (500 000 to 1 000 000 tonnes at the north end) is in place. This increase in production means an inevitable increase in the requirements for men and machinery to break the rock required by the plant. Downhole drilling requirements have tripled, while uphole (both for production and cable bolting) drilling requirements doubled in terms of both machinery and crews. Stope charging requirements also rose drastically with trained chargers’ numbers swelling from four to 16 operators. In all, approximately half a million metres of production drilling is required each year. In order to maintain consistent plant feed grades of copper, uranium, and copper/sulphur ratio, a large number of stopes are required online at any given time. This requires a high degree of flexibility in terms of mine planning, ventilation and development scheduling, the efficient operation of the mine is an intricately interlocked system of planning and the execution of each stage forms an interdependent part of the whole. The achievement of the production targets in order to maintain plant feed and provide critical cash flow to the project, is a fundamental success factor for the mine. The important key to trouble free scheduling of production from the large number of stopes (up to 24) online at any one time, is the materials handling system.
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Design of underground ore pass and rail haulage system The ore pass system is a major component of the expanded operation. It comprises 12 ore passes and 50 finger passes. Tramming distance, the amount of ore to be fed through each of the ore pass and the stoping configuration are the three main criteria considered in designing the ore pass system. Ore from each of the stopes is tipped by a diesel loader through a ‘grizzly’ consisting of four panels of 1.2 m × 1.2 m size into a 3.0 m diameter ore pass. A portable rock breaker is used to break any over sized material. Each ore pass is connected to a 4.5 m diameter surge bin below, (or 4.0 metre diameter bin at the south end), which has the capacity of around 1000 tonnes. An automated rail haulage system has been installed on the 64L (-740 mRL) for ore transportation from surge bins to the crusher. This is based on the LKAB Kiruna Mine which is the largest underground iron ore mine in the world. The rail haulage system is totally computer controlled with no personnel travelling on the train. Once a surge bin exceeds a set minimum level of ore the train is directed to the respective ore chute. Owing to the nature of the orebody, ore passes have been designed at varied angles. The minimum dip angle has been set at 65°. No ground support requirement for the ore passes has been anticipated at this stage. Owing to radiation and ventilation issues all ore passes have been designed on the fresh air side of the mine and operate under negative pressure. Normally an ore pass can have three to five tipping points. Only one tipping point can be active at any one time. A top exhaust ventilation system is used to control reverse flow resulting from tipping into an ore pass. Initially there are two trains in operation, each containing 14 cars, running at a 70 per cent capacity of 300 tonnes per train. Each mine car has a capacity of 14 cubic metres. The rail system has a 1670 tonnes per hour haulage capacity. Based on the current design parameters a single train load of ore can be unloaded into the dump station at a rate of 3000 tonnes/hour. This ore is fed from the Gyratory crusher at the base of the dump station by a plate feeder and conveyor system into two crushed ore storage bins each of 3000 t capacity. The Gyratory crusher has a design capacity of 2000 tonnes/hour, crushing to a maximum size of 200 mm. Crushed ore from the ore bins is discharged onto two conveyors via vibrating feeders. Ore is then hoisted to the surface through the newly constructed Sir Lindesay Clark shaft which is
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equipped with a Koepe double compartment four hoist rope hoisting system. The skips are designed for bottom discharge with a payload of 36.5 tonnes (21.5 m3) the skips discharge into a 120 t capacity Run of Mine (ROM) bin. Ore is withdrawn by a vibrating feeder from the ROM bin, onto a feeder which transfers the ore to a sacrificial, 1800 mm wide, belt conveyor fitted with a tramp iron magnet. Ore is then conveyed by a 1500 mm wide overland belt conveyor to a 1000 tonne surge bin. Ore from the surge bin is extracted by two vibratory feeders onto a 1500 mm wide conveyor to the luffing and slewing stacker. Ore is stockpiled onto two longitudinal stockpiles which has a 120 000 tonne blended total combined capacity. Ore is reclaimed from the stockpiles by operating two adjacent apron feeders which have been designed to maximise ‘rathole’ live capacity. Ore from the feeders is conveyed by 1500 mm and 1800 mm wide belt conveyors from an underground tunnel to AG mill 2 and 3 respectively.
Blending points Owing to the complexity of the materials handling system at Olympic Dam blending is achieved at five locations within the system:
• selective LHD feed from multiple stopes to a common ore pass;
• selective rail truck loading from multiple ore passes; • twin trains tipping into crusher; • ability to hoist ore from both shafts that report to a common overland conveyor; and
• reclaim conveyor loaded from multiple feeders. WASTE ROCK (MULLOCK) Waste rock generated from the underground development is generally tipped direct into a suitable secondary or tertiary stoping void. When such a void is not available, the waste rock is hoisted to the surface by campaigning in place of ore, and is fed through a 500 tonne surface mullock bin adjacent to the 1000 tonne ore bin. The waste rock is stockpiled and then used in the manufacturing of CAF.
BACKFILL In the Olympic Dam Operation stope back filling plays a critical role in stope sequencing and the design process. Following the extraction, depending on the individual circumstances, stopes are filled with CAF, underground mullock or a combination of both of these materials. A typical CAF mixture has 57 per cent of crushed rock by wet weight, 26.5 per cent of deslimed mill tailings and sands, 2.5 per cent of cement, five per cent of pulverised fly ash and nine per cent neutralised tailings liquor producing a nominal fill strength of 3 Mpa. Specific CAF mix designs exist to allow strengths to be delivered from 1.0 Mpa to 5.0 Mpa. The CAF plant has been upgraded and will continuously mix CAF in a pug mill at 6000 m3/day. Crushed rock for the CAF plant is supplied from a purpose excavated limestone quarry with the current annual requirement being approximately two million tonne. Filling cost is a significant portion of the mining cost. The possibility of reducing this cost while maintaining the minimum fill strength for the given stope geometry is currently being investigated. Some of the areas under investigation are CAF mix, alternative binders, stope geometry, CAF strength and continuous CAF delivery systems. An overall review is being conducted to identify areas where the non-cemented fill proportion can be increased.
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The CAF plant has a maximum capacity of 350 m3/hr and operates at 300 m3/hr. This is achieved by continuous mixing in a Pug Mill with discharge in to CAF surge bins for loading in to CAF delivery trucks of 15 m3 capacity. CAF is delivered to a rock dropper bore hole which delivers from the surface to the top of the stope. Underground pipe work reticulation is currently being tested with a view to establishing a continuous delivery from the CAF plant in the longer term.
TECHNOLOGY INNOVATIONS Ore from a stope is currently mucked using conventional loader units. Remote mucking is used only to allow the retrieval of the final broken ore from the open stope at the end of the production phase. WMC and Lateral Dynamics (Australia) have successfully tested a Loader Automation System at the WMC Leinster underground operation in Western Australia. This system comprises laser sensors and on board computers allowing guidance from a designated point as well as tele-remote control. It has been successfully tested travelling with a minimum clearance of 50 cm against the wall. Direct production trial comparisons of the Autotram technology performed at the Leinster operation has shown a 20 per cent increase in productivity against the manual remote mucking. Considering the reduced travel time and the ability to work through blast times it has been estimated that the system can achieve up to a 40 per cent increase in productivity. The project is now being tested at Olympic Dam using an Elphinstone R2900 loader mucking from stopes to ore passes. Commercial availability of this unit is expected with in year 2000. Automation and remote control of mobile equipment such as LHDs is more challenging than the track bound train system. Such a system will add flexibility to the existing ventilation requirements whilst improving safety and system efficiency.
Mine control room The complete materials handling system from the ore passes to the surface stockpile is monitored and controlled in the mine control room. This purpose built dust free room houses numerous video screens and computer monitors, sensors installed at strategic locations update screen mimics with information such as ore pass levels, productivity rates, activity status and process availability. Video cameras positioned at the train loading points transmit real time video coverage to the control room assisting the mine controller with his function of loading the train. Video coverage and control is also available for:
• checking for spillage on the rails beneath the loading chutes; • viewing the dump point and the presence of any large rocks in the crusher throat;
• viewing the feeders delivering crushed ore onto the loading station belt;
• viewing the skip at the tipping point in the sky shaft; • monitoring the overland conveyor; and • monitoring the stacker and stockpile formation. CONCLUSIONS Olympic Dam Corporation has in place a materials handling system that can provide consistent quality mill feed at a rate of 30 000 tonnes per day and we are on track to achieve our nameplate production target of 200 000 tonnes of copper and associated products in our first full year of production. Olympic Dam has accepted the challenge, endured the pain, and is now out of the blocks and leading the race for GOLD … and SILVER and COPPER and URANIUM.
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The Automation of Western Mining Corporation’s Olympic Dam Underground Rail Haulage System R Doubleday1 and D Mee2 ABSTRACT The Western Mining site at Olympic Dam, near the community of Roxby Downs in South Australia is one of the largest orebodies of its type in the world. It contains known reserves of over 10.6 M tonne of copper, with economic secondary metals including uranium, silver and gold. In 1996, an expansion project was commenced to increase mine production approximately threefold, to produce 200 000 tonnes pa of copper, 4600 kg pa of uranium oxide, 3600 kg pa of gold and 50 000 kg pa of silver. This expansion included both underground mining and surface processing investment. Because of the scale of operations, it was decided that ore should be transported underground by a railway system. In July 1997, ALSTOM Automation and Control Limited (then Cegelec Australia Limited), was awarded the contract to deliver a control system to manage and automate the rail haulage system for Western Mining’s Olympic Dam Mine, as part of this major upgrade and extension project. This paper explains the overall design and the approach taken to implement route selection, route locking, train interlocking and track topology using normal industrial control equipment such as Programmable Logic Controllers. Included is a discussion of how this automation has improved the productivity of the haulage system by more than 15 per cent.
INTRODUCTION The Western Mining site at Olympic Dam, near the community of Roxby Downs in South Australia is one of the largest ore bodies of its type in the world. It contains known reserves of over 10.6 M tonne of copper, with economic secondary metals including uranium, silver and gold. In 1996, an expansion project was commenced to increase mine production approximately threefold, to produce 200 000 tonnes pa of copper, 4600 kg pa of uranium oxide, 3600 kg pa of gold and 50 000 kg pa of silver. This expansion included both underground mining and surface processing investment. Because of the scale of operations, it was decided that ore should be transported underground by a railway system. In July 1997, ALSTOM Automation and Control Limited (then Cegelec Australia Limited), was awarded the contract to deliver a control system to manage and automate the rail haulage system for Western Mining’s Olympic Dam Mine, as part of this major upgrade and extension project.
UNDERGROUND RAILWAY OPERATIONS The underground rail haulage system includes a rail track network at the 640 level, below the main orebody. Up to three trains (two at the initial stage) are to run on this track simultaneously, each train having a front and rear loco and from 12 to 18 wagons, 16 being typical. In normal operation, there are no crew on the train. A typical journey for a train consists of a return trip from the central dumping station to one of over 12 loading chutes. A complete cycle averages 30 minutes. At the loading chute, the train and load chute are remotely controlled from the mine operation room (on the surface) to load ore into
1.
Chief Engineer – ALSTOM, Automation Systems, ALSTOM Automation and Control Limited, 373 Horsley Road, Milperra NSW 2214.
2.
Manager – ALSTOM, Intelligent Systems, ALSTOM Automation and Control Limited, 373 Horsley Road, Milperra NSW 2214.
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each of its wagons. At the dumping station, the train travels at low speed across the dump where the wagons are tipped to empty the ore into a large rock crusher. The crushed ore is hauled to the surface by a mine hoist. ALSTOM also supplied the electrical and control system for the hoist, which included a 6.5 MW ac integral drum winder powered by a cyclo-converter. The train, now empty, is free to begin another outward journey to load up with more ore. The topology of tracks and chutes is modified as the mine is developed.
LOCOMOTIVES AND ROLLING STOCK Each train consists of two locomotives, one at each end of a rack of 12 to 18 ore wagons. The UK supplied locomotives take their power from an overhead catenary at 600 V dc. A pulse width modulated power transistor switching system regulates power to the dc traction motors. The motor control, pneumatic braking and loco auxiliaries are all controlled by a LMC (Locomotive Controller), supplied by the loco builder. This subsystem takes direct inputs from the controls on a driver’s panel, for maintenance or manual operation, a locomotive, or pair of locomotives may be operated by on board crew using a joystick and control switches on the driver’s panel. Remote control of each locomotive is achieved through the use a serial port to the LMC. The status of the LMC is gathered continuously. When a change is detected it is passed to the higher levels of the control system. Typical auxiliary commands include raising and lowering the pantograph, controlling the horn, etc. When two locos are coupled into a complete train, there is coupling between the LMCs as well. During commissioning, it was found that a hard wired serial connection proved more reliable than the originally proposed radio link. This connection allows the two locos to work in tandem, with one a designated master and the other a slave.
RAILWAY NETWORK AND CONTROLS The track layout is one of a set of spurs that branch from a central ‘yard’, consisting of a dumping station, by-pass loop, and routes into and out of the maintenance workshop loop. When a loaded train enters the dump station, the locos are suspended with their wheels off the track, while the carriages are forced to follow a non-horizontal path that causes the ore to spill out of the hinged side door and onto the top of the crusher stockpile. Having emptied, the train proceeds past the dump station, and then may reverse and enter the loop, to go back in the direction it came from. Alternatively it may continue in the same direction and move to a loading point at the other end of the mine. The ends are referred to as North and South ends. While a loco is moving through the dump station, it receives no power from the catenary, and can provide no traction, as its wheels are not in contact with the track. Thus it relies on the push or pull from the loco at the other end of the train. The arrangement of track and catenary is such that at least one loco is receiving power or exerting traction at any one time. A power cable between the locos ensures that as long as one pantograph is in contact with the catenary, power is fed to both locos.
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Automatic Loco Controllers (ALCs)
LOADING ORE Each load chute is essentially a large storage bin constructed from the rock, which at its lower end has a hydraulically controlled steel gate. Opening or closing the gate regulates the volume of ore that pours from the throat of this device. In the ore pass above each loading bin, the ore is mined and tipped into the tops of the bins through large screens (grizzlies). An ultrasonic device measures the level of the ore in the ore pass. Load chutes may be operated either remotely, from the mine control room on the surface, or locally using a control pedestal at each chute. Manual operation is only used to clear abnormal material conditions that prevent operation from the control room. During a manual loading operation from the chute pedestal, the train is first placed in its initial position beyond the chute, and is then remotely controlled to begin moving in a direction towards the dump station side of the chute. As the carriages pass under the chute, the operator controls the throat opening and the train speed so as to fill each carriage as it moves past. The chute is closed as the final carriage is filled, but before the trailing loco moves under the chute. In remote loading from the mine control room, essentially the same procedure is carried out using two closed circuit television (CCTV) channels to provide vision of the loading operation from above and the side of the train. At each of the load chutes, there is a break in the overhead catenary, to allow the ore to flow from the chute into ore wagons without damaging the catenary. To prevent the live pantograph from connecting with, and being damaged by, the steelwork of the loading chute, it must be isolated from the 600 V dc bus on the locomotive and lowered as the train passes under each chute. Furthermore, since the spacing of the load chutes along some of the loading spurs is only slightly longer than the maximum length train, accurate speed and position control are required to ensure that both pantographs are not lowered concurrently.
CONTROL ROOMS AND STAFFING The Rail Haulage system can be controlled from either or both of two controls room. The system is normally controlled from the Mine Control Room (MCR) located on the surface. The MCR has two operator workstations. Each workstation consists of a control system HMI screen, two CCTV screens and a specially constructed operator’s chair, with joystick controls for the load chute and train operation during loading. There is another complete operator’s workstation underground, at the 640 level Locomotive Maintenance Room (LMR). This workstation is not normally staffed. One person normally operates the Automatic Rail Haulage system. This person usually loads one train while the other is in transit to the dump station or to a load chute.
CONTROL SYSTEM STRUCTURE Following an extensive analysis of the operational, maintenance and environmental requirements ALSTOM chose to implement the control system into the following sections: 1.
Automatic Loco Controllers (ALC),
2.
Train Control System (TCS),
3.
Load Chute Controllers (North and South),
4.
Human Machine Interface (HMI), and
5.
Train Scheduler.
The Automatic Loco Controllers (ALCs) are programmable logic controllers (PLCs) located on-board each locomotive and form the mobile part of the overall system. The other parts of the system are land based and connected together with an extensive, duplicated fibre optic network.
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The ALCs form the mobile part of the control system. It is the primary role of the ALC to maintain an accurate location of the locomotive and report this location to the TCS. In return, the TCS passes track information about the state of the rail network, route requests, and start/stop commands to the ALC for execution. Using the track status information, the required route request and the stored information about the network topology, the ALC commands the LMC to move or stop appropriately. As the train progresses along its route, the ALC operates lights, horn and pantograph as required. These operations are stored in tables inside the ALC and are triggered by the location of the train. The automatic operation of the pantograph has greatly reduced the cost associated with damage to electrical equipment and lost production caused by incorrect manual operation of the pantograph. Determination of position is crucial for safe and efficient operation of the whole automatic train system. The ALC determines its position using two methods and compares the results to determine if the results are valid. Continuous position is determined using a pulse encoder connected to the locomotive’s wheels. This position is compared against radio tags that are embedded in the track approximately every 100 m. Any significant error in position is removed at each tag. Position and Tag numbers are stored in tables in the ALC. Tags missed, read early, late or out of sequence cause the automatic operation of the train to stop and alarms raised to the operator. The ALC communicates with the rest of the control system over a serial link across a dedicated leaky feeder radio network installed in the 640 level of the mine. During the course of commissioning, significant improvements to the reliability of this important link was achieved by good radio engineering, correct installation and placement of the leaky feeder antenna cable and the correct placement and tuning of in-line distribution amplifiers. To improve the availability even further, a high speed serial link was established between the ALCs to provide an additional path from the TCS to the ALC via the other ALC.
Train Control System (TCS) The TCS is the central, land based train control PLC. Its primary function is to evaluate the condition of each section of track based on the position of each locomotive and the condition of land based equipment such as track points, access gates, ventilation and traction power. This information is gathered directly from I/O connected to the PLC. The TCS takes the position of each locomotive, and associates it with its corresponding train. Sections of track occupied by each train are marked as occupied. Using the position of the trains, equipment such as points are prevented from moving. Point locking prevents points from moving while a train is near or over a set of points. Operator or automatic requests for a train movement are evaluated by the TCS. If all the sections of track involved in the route of the train are available, each section of track is allocated to that train, otherwise the request is ignored. Code associated with the track points evaluates the allocation of trains across the sections of track that make up the points. If the points are not otherwise interlocked, (locked, in manual, out of service, etc) the points are moved to the required position. If all the points are not in the required position before a certain timeout, the route allocation fails and the allocation is removed. If the allocation is successful, the route number is sent to the ALC. From this point the ALC continuously evaluates the track status of the selected route to ensure it is safe for it to proceed along it. Following successful allocation of a route, the TSC automatically signals the CCTV switch to sequentially display the cameras along the route to the operator so that they may determine if any spillage or any other obstruction exists. If there is no obstruction, the operator sets the train in motion.
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THE AUTOMATION OF WESTERN MINING CORPORATION’S OLYMPIC DAM UNDERGROUND RAIL HAULAGE SYSTEM
There are two Load Chute Controllers, one for the northern load chutes and the other for the southern chutes. The main function of these PLCs is the operation of the hydraulic controls associated with the load chutes and associated spur ventilation. When a route associated with a loading chute is selected, the corresponding spur and chute ventilation is started and the hydraulic power pack for the chute is activated. The commands, from the mine control room or the local control panel are used to operate the train and the throat of the load chute.
the competing requirements of safety, cost, simplicity, flexibility, maintainability and completeness. Industrial ore haulage systems provide no small challenge. From ALSTOM’s main line rail transport experience, the safety factors associated with the operation of trains are well understood and the implementation of signalling systems provided an excellent framework for our automation design. The real challenge is to translate these requirements into a industrial environment with generally available, cost-effective, flexible hardware/software systems rather than the relatively costly domain specific hardware and software systems associated with main stream railway signalling. What follows are some of the data structures implemented after extensive analysis of the operational requirements.
Human Machine Interface (HMI)
Sections (of track) Table
The human machine interface (HMI) for the control system was implemented as an integrated part of the site wide SCADA system used by Western Mining at Olympic Dam. The HMI provided two modes of control for each train. In ‘Auto’ mode, the Train Scheduler (TS) dispatches the train to a load chute according to a set of rules in the TS. The TS automatically requests the route for the train and if successful, the operator must acknowledge the CCTV scan of the route after which the train automatically proceeds to the loading chute and is postioned for loading without further operator intervention. Average trip time is about five minutes. The operator is prompted to load the train using the joystick controls. When the operator takes control of the loading, the CCTV views are automatically adjusted to show the overhead and side views of the rail loading operation. When the loading is completed the route through the dump station is selected and the train travels to the dump and is emptied without further operator action. When the train is in ‘Manual’ mode, the operator selects train and route to travel rather than having the TS select them. The movement of the train along the route is still automatic. This allows the operator to position trains for non-standard operations such as trips into and out of the maintenance workshops. The HMI supplies a comprehensive alarm and event logging facility along with a dynamic rail system overview showing the position and condition of all elements of the system.
A section represents the smallest length of track that is managed in terms of its availability, allocation and occupancy. Each end of a section has a location relative to a known datum. Many operational and safety requirements are embedded in the selection of section boundaries. For example, the section lengths associated with points are chosen to ensure the approach distances of approaching trains on the fork of the points do not interfere. As the topology of the track network grows, so do the number of entries in the Section Table. The Section Table is a static table and is common to the TS, TCS and the ALCs.
The ALC then proceeds along the route, following a speed profile stored in tables in the ALC for each route. The ALC then stops the train at the required stopping position for the route.
Load Chute Controllers (North and South)
Train Scheduler (TS) The Train Scheduler (TS) is a redundant pair of Personal Computers running the IP.21 real time database. ALSTOM has used this database in many SCADA and industrial control disciplines including metals processing line, road traffic management systems, port terminal materials handling systems and power station history monitoring systems. Both TSs are running concurrently. Only one is selected as Master by the TCS. The main function of the TS is to analyse the current state and position of the trains, the current condition of each of the routes and the priority, condition and level of each of the loading chutes. Every 20 seconds, using a relatively simple set of rules, the TS performs the analysis. If any of the trains are in ‘auto’ mode, and the analysis indicates the train should move, the corresponding train and route request is passed to the TCS and the functions described above are set into action. An auxiliary function of the TS is to check the various configuration tables stored in the TCS for validity to ensure inadvertent changes are not made to critical system data.
DATA MODEL OF THE RAIL HAULAGE SYSTEM One of the more challenging aspects of creating any system of automation is to find a model of the physical and operational characteristics of the system to be automated that provides for
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Section Route Table A route is a collection of sections required for a journey. Information associated with a route includes the direction of travel and the type of routes (for example a loading route under a load chute). By specifying a single route number, the elements of the system (TS, TCS, ALC, etc) can specify or determine all of the track network to be used during the operation of the train across the particular route. The Section Route Table is a static table and is common to the TS, TCS and ALCs.
Speed Profile Table A table of speed zones for each route is stored in each ALC. This static table is used by the ALC to control the speed of the train as it traverses the track.
Train Equipment Tables Static tables exist in the ALC to determine where, along each route, the ALC should sound its horn and operate pantographs and lights.
Tag and Tag Route Tables These static tables exist in the ALC to describe the location of each tag embedded in the track and which tags should be detected by the ALC as it traverses the route.
Section Status Table This is a dynamic table whose contents are continually re-calculated by the TCS. Each section has availability based on any hardware or software interlock, a service state (in or out of service), allocation status to each train (only one train at a time, of course) and an occupancy status (again only one train at a time). This table is transferred, on a change of state basis, to the ALCs. The ALC uses this to ensure that the track and associated equipment ahead is clear.
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Route Status Table This dynamic table is calculated continually by the TCS from data in the Section Status table. The table provides the dynamic state of every route including its allocation and availability. The TS uses the Route Status tables to determine which routes are available for any idle trains. These tables are managed as spreadsheets that are down loaded to the processors in the system.
SOFTWARE DEVELOPMENT, TESTING AND COMMISSIONING The supply side of the project was performed in ALSTOM’s NSW engineering offices. The requirement for as much off-site testing as possible lead to the construction of a model train network to simulate the real system to a scale of 1 to 500. N-scale locos were used with a digital controller and controlled switch points. The model development allowed the safety and table based control algorithms to be very well tested prior to delivery to site. As a result, most of the on-site commissioning activity involved proving and tuning the static data tables described above rather than making changes to the system software. The software was developed following ALSTOM’s accredited ISO9001 quality system, and extensively tested using the model train system, resulting in an exhaustive witnessed factory test. It was then delivered to site (June 1998) and commissioning began in the middle of an extensive construction period on the mine site. The commissioning had to take place in the middle of this busy construction program with all the difficulties of access, timing and availability that that entails. In particular, much of the track work and many of the load chutes were not constructed, and the mine environment placed a heavy demand on the robustness of the locos and track equipment. There were operational and financial pressures to begin the haulage of ore immediately. As a result, the trains were driven manually for at least one of the two 12-hour shifts per day. However, as a result of the fundamental design and the robustness of the implementation, the commissioning was able to proceed along whichever path the available equipment took it. One of the most challenging commissioning activities proved to be the interfacing of our rail haulage system to the radio system used in the mine. Although a dedicated leaky feeder antenna system was installed to support our system, it initially
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proved unreliable due to high RF noise generated by the locomotive dc chopper drive system, the physical installation of the antenna cable and the location and tuning of distribution amplifiers. All aspects of the rail haulage system have now proved themselves in automatic operation. All of the originally planned track construction has now also been finished and an additional load chute has been added. All aspects of the control system are now in daily service. The mine operators are now enjoying maximum haulage rates, 24 hours per day, at low cost, using the ALSTOM rail haulage automation system.
BENEFITS The Olympic Dam Mine operates on two 12-hour shifts per day. Between shifts, lower levels of the mine, including the rail haulage level, are evacuated to allow for production and development blasting. The Automatic Rail Haulage System does not require crew underground, and can continue to haul ore during the shift change. This manifests itself in a 15 per cent productivity improvement in the utilisation of the rail haulage system, and hence contributes significantly to the overall mining operation. Other improvements in equipment utilisation are realised during periods of adverse radiological or other environmental conditions when parts of the rail haulage level become unsuitable for human access. The automatic train can continue to operate through these conditions. Manual operation of the Locomotives has been characterised by much higher levels of wear and tear and equipment damage than automatic operation. If the increased levels of breakdown and maintenance time are taken in to account, the over all productivity of the automatic system is significantly greater than crewed shifts.
CONCLUSION Increased productivity and better utilisation of expensive capital works can be obtained by the appropriate use of automation. In the case of WMC’s Olympic Dam mine the 15 per cent increased equipment utilisation, reduced break down and maintenance costs and reduction in underground staff provide compelling arguments for, and rapid return on investment in, the ALSTOM automatic rail haulage system.
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Open Stope Design at Normandy Golden Grove Operations T M Calvert1, J B Simpson2 and M P Sandy3 ABSTRACT The empirical stope design techniques developed by Mathews et al (1981) and refined by Potvin (1988) and others, have gained rapid acceptance in Australia in the 1990s as a simple, ‘first-pass’ means of designing primary stopes. However, it is recognised that these ‘stability graph’ methods can only provide broad design guidelines. Recently published stability graphs (Hutchinson and Diederichs, 1996) encompass a wide range of geological environments, which are not discussed, and the influence of time and blasting practices cannot be easily reviewed. The ‘scatter’ in the graphs must reflect, to some degree, different operating practices as well as subtle but important differences in the pre-mining conditions that are not adequately accounted for in the ‘modified stability number’. Development of a ‘local’, site-specific stability graph may overcome some of these problems. At Normandy’s Golden Grove Operations, Western Australia, detailed reviews of stope stability performance, including Cavity Monitoring System (CMS) surveys, are undertaken for every stope, with the objective of developing a series of local stability graphs, customised to site conditions and calibrated in terms of ‘Average Falloff Thickness’ (AFT). 1.
MAusIMM, 148 Maida Vale Road, High Wycombe WA 6057.
2.
MAusIMM, Mount Isa Mines Ltd, Mount Isa Qld 4825.
3.
MAusIMM, Australian Mining Consultants Pty Ltd, 9 Havelock Street, West Perth WA 6005. E-mail: [email protected]
Cablebolting of stope hangingwalls is used to develop reinforced ‘beams’ on each sublevel, effectively dividing the full span into smaller, more stable ‘subspans’. By analysing stope stability performance in terms of fall-off, cost-benefit studies can be undertaken of the effect of cablebolting and other options for modifying stope spans. The approach used to develop the local stability graphs, and their integration into routine stope design procedures at Golden Grove, is described. The experience gained during stope extraction in the Scuddles orebodies has been successfully applied to the Gossan Hill ‘C-Panel’ orebody. This methodology should be applicable to other underground mining operations.
INTRODUCTION Golden Grove Operations is located 230 km east of Geraldton and 50 km southeast of Yalgoo in Western Australia (Figure 1). The operation comprises the Scuddles mine and treatment plant, the adjacent Gossan Hill mine, and 120 km2 of exploration tenements covering the 35 km long mineralised belt. Normandy Golden Grove Operations, a wholly owned subsidiary of Normandy Mining Limited, owns and manages the operation. Prospecting of the Golden Grove belt for base metal deposits commenced in 1971. By 1973, a resource of 15 million tonnes grading 3.4 per cent copper had been estimated at Gossan Hill. The non-outcropping Scuddles zinc-copper deposit was discovered in 1979 and development commenced in 1988, with the mine entering production in 1990. Gossan Hill was brought into production in mid-1998, concurrent with an expansion of the treatment plant to 1.2 million tonnes per annum. In 1998 - 99, Golden Grove produced 110 355 tonnes of zinc and 3130 tonnes of copper in concentrate.
FIG 1 - Golden Grove location map.
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The Scuddles and Gossan Hill deposits are hosted by Archean greenstones of the Warriedar Fold Belt, and occur in a sequence of felsic lavas, tuffs, volcanoclastics and sediments intruded by numerous mafic and felsic dykes. At mine scale, a steeply west-dipping horizon of thinly bedded cherts and tuffaceous sediments hosts the zinc mineralisation at both deposits and is underlain by coarse pyroclastics and epiclastics. The Scuddles deposit is over 600 m long and up to 50 m wide. It occurs as three distinct stratabound massive sulphide lenses Main, Deeps and Central - which are immediately underlain by a 40 m to 50 m wide zone of stockwork mineralisation. The massive sulphides are principally sphalerite, chalcopyrite and pyrite. In the stockwork zones, pyrite and chalcopyrite are the common sulphides. The Scuddles orebody is accessed by shaft and decline. The 4.0 m diameter circular shaft, used for ore haulage extends to a depth of 630 m. A 4.5 m × 5.0 m decline from surface, at a gradient of 1:7, provides access for personnel and materials. Highly mechanised sublevel open stoping methods are employed. The 30 m – 55 m high stopes are backfilled with a mix of tailings, development mullock and screened waste. The permanent workforce (approximately 270 including contractors), which operates on a ‘two weeks on one week off’ basis, resides in Perth and Geraldton.
STOPE DESIGN The empirical stope design techniques developed by Mathews et al (1981) and refined by Potvin (1988) and others, have gained rapid acceptance in Australia in the 1990s as a simple, ‘first-pass’ means of designing primary stopes. However, it is recognised that these ‘stability graph’ methods can only provide broad design guidelines. Recently published stability graphs (Hutchinson and Diederichs, 1996) encompass a wide range of geological environments, which are not discussed, and the influence of time and blasting practices cannot be easily reviewed. The ‘scatter’ in the graphs must reflect, to some degree, different operating practices as well as subtle but important differences in the pre-mining conditions that are not adequately accounted for in the ‘modified stability number’. Development of a ‘local’, site-specific stability graph may overcome some of these problems. A number of geotechnical and mining reviews of Scuddles and Gossan Hill have been undertaken since the project commenced. Feasibility studies were prepared as part of the project approval process for the Deeps (at Scuddles) and Gossan Hill for the A, B and C Panel zinc orebodies (Simpson, 1996; Yesil, 1997; Sandy, 1997). As part of the geotechnical studies the effects of stress distribution on the stability of previous, current and likely future stopes at Scuddles Mine were assessed using 3D stress analyses. The analyses indicated that various types of failure could develop at Scuddles as a result of stress re-distribution around the stopes, including rock spalling in the corners due to high stress concentration, and slabbing and block failure in the backs and hangingwalls due to stress relaxation. Observations underground have supported the modelling results. With the relatively strong ore types at Scuddles and Gossan Hill, deep stress-related failures have generally not been experienced, except in late stage, isolated pillars.
STABILITY GRAPH METHOD The stability of stopes is not only controlled by stresses but by geological structures, particularly in stope hangingwalls and footwalls which are often ‘relaxed’, ie they experience a loss of confinement during stope extraction. Stope wall inclination also has an important influence on stability. To assess these effects, the stability graph method, (Mathews et al, 1981; Potvin, 1988 and Hutchinson and Diederichs, 1996) was applied. This
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considers the effect of stresses and structures as well as stope size and shape. Stoping conditions are described by a stability number, N’, which is defined as: N’ = Q’ * A * B * C where Q’ = modified tunnelling quality index (after Mathews et al, 1981) A = the rock strength factor which is defined as the ratio of uniaxial compressive strength of intact rock, σc, to the induced compressive stress, σi, on the stope boundary under assessment. ‘A’ value ranges are as follows: A = 0.1 for σ c/σi < 2 A = 0.1125*(σ c/σi) - 0.125 for 2 < σ c/σi < 10 A = 1.0 for σ c/σi > 10 B = the joint orientation adjustment factor which depends on the difference between the orientation of the critical joint and the stope face under assessment. C = the gravity adjustment factor which is based on the potential for rockfalls from the roof or slabbing or sliding from the stope walls due to gravity. Prior to the current study, stope hangingwall spans were designed based on general ground conditions; typically a Q’ value of 30 was assumed for all stopes in the Deeps and 20 for central stopes, and this was adjusted using the ‘A’, ‘B’ and ‘C’ parameters to obtain N’. This in turn gave an indication of ‘allowable’ strike length. Typically, for Scuddles, ‘B’ was set to 0.2 and ‘C’ was set to 6.5. This value was then plotted on the appropriate stability graph, to determine a design hydraulic radius. Given that the sublevel interval is usually fixed, the main design variable was stope strike length.
GEOTECHNICAL DATABASE The Rock Tunnelling Quality Index, Q, and the Modified Rock Tunnelling Quality Index, Q’, are calculated from borehole core samples and compiled in a database. Geotechnical logging of all drill core at Golden Grove has been undertaken since 1990. This logging involves the evaluation of the core for the following parameters (after Barton et al, 1974):
• RQD = (length of core > 100mm)/(length of run) × 100 per cent
• Joint Number Factor, Jn • Joint Roughness Factor, Jr • Joint Alteration Factor, Ja The Stress Reduction Factor (SRF) is set to 1 or 2, and Jw (the Joint Water Factor) commonly has a value of 1, as conditions are dry and impermeable. These factors are entered into a database in Micromine, and the Tunnelling Quality Index, Q, and Modified Tunnelling Index, Q’, are calculated. Q’ values are linked to the Vulcan mine planning software for use in mine design. Typical Q’ values for the different rock types within Scuddles are given in Table 1. The stability problems at Scuddles can be summarised as:
• stope stability in terms of hangingwall and back stability; • pillar (secondary stope) stability; and
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TABLE 1 Typical Q’ values for Scuddles rock types. Unit
Typical Q' Range
Hangingwall Sediments
10 - 20
Hangingwall
20 - 70
Orebody
40 - 100
Footwall
40 - 100
• stability of other mine openings such as declines, drives, cross-cuts, etc. Numerical and empirical analyses were carried out in 1996, to establish the effects of stress re-distribution and geological structures on the stability of mine openings. Numerical modelling was carried out using MAP3D-SV (Wiles, 1996). The program was used to simulate previous and current stopes in Main lens by modelling their actual mining and filling sequences. Various dimensions of primary and secondary stopes were also modelled to simulate future mining geometry in the Central and Deeps lenses. The models provided estimates of mining induced stresses, strains and displacements around the openings. To calibrate the model, two overcoring in situ measurements of pillar stresses were performed in the 460 level hangingwall. The test results were in close agreement with the stress values predicted by the model, improving confidence in the use of modelling to back analyse past stoping, and as a predictive tool. To assist in analysing the performance of stopes, records were maintained within a stope folder. A geological reconciliation was undertaken for each stope. In some stopes a Cavity Monitoring System (CMS) survey was undertaken by contractors and a wireframe model created for use in VULCAN three-dimensional mine modelling software. The Cavity Monitoring System gathers comprehensive three-dimensional survey data from underground stopes and voids which are inaccessible to traditional surveying methods. CMS surveys have been widely used in Australia since the mid-1990s (Gilbertson, 1995; Rauert, 1995). The equipment comprises a laser rangefinder integrated into a motorised scanning head, which is attached to a horizontal carbon fibre boom and inserted into the stope. The rangefinder rotates around the boom to survey the cavity and data is sent via cable to a data logger. Volume calculating software then provides an XYZ or DXF file in the mine co-ordinate system which can be imported into CAD systems or mine modelling software such as VULCAN, DATAMINE or SURPAC. With the stope void imported into a three-dimensional wireframe model, it becomes a powerful tool for analysing stope performance. The CMS thus provides data to reconcile the actual stope shape against that of the design, and determine fall-off and dilution within the stope using the mine modelling software. Examples are presented in Figures 2 and 3. CMS surveys can also be undertaken at different stages of stoping to monitor dilution as stoping progresses. In some circumstances, this can assist the rapid development of a local stope stability database. Normandy Golden Grove Operations purchased their own CMS for stope performance monitoring in 1998.
STOPE PERFORMANCE MONITORING AND RECONCILIATION A decision was made in early-1998 to quantify stope performance at Golden Grove.
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FIG 2 - Cross-section through a single lift bench stope at Scuddles.
Geotechnical reconciliations have been undertaken since April 1998 on completed stopes. The aim of the reconciliations was to develop a local stability graph and dilution estimation chart from which to better analyse stope performance and to understand the behaviour of the hangingwall. This in turn could be used to improve stability and minimise dilution in future stope design. The reconciliations take into account the installed rock support and reinforcement and the ground conditions. The geometry of the stope is noted.
Methodology Using triangulations of the designed stope shape and the final stope shape (based on the CMS survey) the performance of the stope can be analysed. This allows the Average Falloff Thickness (AFT) of the hangingwall to be calculated, and an estimate of dilution tonnes to be made. This is essentially the same as the ‘ELOS’ concept (‘Equivalent Linear Overbreak/Slough’) described by Pakalnis et al, 1996. The attraction of this approach is that AFT is insensitive to orebody thickness, whereas expressing dilution in percentage terms requires information about the orebody thickness. Results from orebodies of different thicknesses cannot be compared directly. The stability graph method is used in the analysis to assess how the designed stope would be expected to perform, and if the amount of falloff from the hangingwall could have been predicted and avoided. It has been found, thus far, that the performance of stopes generally correlates well with the published stability graphs (Potvin, 1988; Nickson, 1992).
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Results The results are summarised in the Scuddles local stability graphs in Figures 5 and 6. Individual stope case histories relating to the stopes in the database are presented in these figures:
STOPE CASE HISTORIES 600z39 Dilution in this stope was negligible even though the stope plotted in the transition zone on the Potvin and Nickson stability graph. The lack of dilution was probably due to the ‘skin’ of ore left on both the hangingwall and footwall after firing. Approximately 33 800 tonnes of ore was recovered from the stope, with the final CMS stope volume 8050 m3.
370z53 This stope had extensive dilution with an average fall-off thickness of 4.8 m (approximately 15 000 tonnes). It plotted in the unstable/caving zone on Potvin and Nickson’s stability graph. Several factors contributed to the dilution in this stope. The development on the 370 level had undercut the stope hangingwall. The hangingwall on the 370 level was cablebolted prior to the mining of the stope below (440z53) however this was not able to contain the hangingwall failure in 400z53. This continued up through the hangingwall of 370z53. During the filling of 400z53 there was a failure of approximately 2 m of ore from the backs (ie the bottom of 370z53). The hangingwall dip was shallower than average for Scuddles; this was considered a significant factor in the mining of the stope. There were no unusual features in the hangingwall geology which would indicate potential failure. 63 000 tonnes of ore were recovered, with the final CMS stope volume 18 942 m3.
370z48 FIG 3 - Cross-section through a two sublevel panel stope at Scuddles.
In compiling a geotechnical reconciliation the first step is to compile the physical information relating to the stope, such as reserve tonnes, stope dimensions, hangingwall dip and geology, and the blasting practice. The hydraulic radius is then calculated and Q’ values estimated for the hangingwall using Vulcan and the NGI drilling database. Figure 4 is an example of a stope reconciliation data sheet. Two values of Q’ are calculated - for the first 5 m into the hangingwall and the next 5 m, ie Q’0-5 m and Q’5-10 m. This approach can assist interpretation, for example where fall-off is arrested by a competent zone in the hangingwall. In reconciliations the median value of Q’ is used. When designing a stope the ‘lower quartile’ value of Q’ may be more appropriate, particularly in new operations or areas where limited stoping experience is available. The rock strength factor A is calculated and the gravity adjustment factor C is determined from the charts based on the failure mode (Hutchinson and Diederichs, 1996). From these values N’ is calculated and the stope is checked on the appropriate stability graph. The average fall-off thickness (AFT) can be calculated by comparing the designed stope outline to that of the CMS survey after stoping is complete. In a multiple lift stope, where the hangingwall has been reinforced on the intermediate sublevels, the span may be effectively divided into smaller ‘subspans’. These results should be analysed separately from the single span database.
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The average fall-off thickness in this stope was 0.9 m from the hangingwall (approximately 4700 tonnes). On the Potvin and Nickson’s stability graph it plotted in the stable zone. The hangingwall fall-off was quite uniform across the stope. 32 400 tonnes of ore were recovered, and the final CMS volume of the stope was 6808 m3.
350z49 Dilution averaged 0.9m from this stope’s hangingwall (approximately 4300 tonnes). It plotted in the stable zone of Potvin and Nickson’s stability graph. 37 200 tonnes of ore were recovered, and the final CMS volume of the stope was 11 483 m3.
235z16 There was a 1.65 m average hangingwall fall-off thickness in this stope (approximately 6500 tonnes). It plotted in the stable zone of Potvin and Nickson’s stability graph. Although the hangingwall geology was ‘typical’ with no obvious dykes or intrusions, the hangingwall had been undercut from the previous stope. 29 200 tonnes of ore were recovered, and the final CMS volume of the stope was 8407 m3.
235z46 This stope had quite extensive dilution (approximately 18 000 tonnes) with an average fall-off thickness of 3.2 m. The stope plotted in the stable zone of Potvin and Nickson’s stability graph, however it was influenced by the hangingwall geology.
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FIG 4 - Example of stope reconciliation sheet.
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700z28/725z27/750z26
FIG 5 - Scuddles stopes plotted on a modified stability graph.
This stope was a multiple lift stope, with a hangingwall drive in the middle of the overall span, which was supported with cables. A separate local stability graph should be developed for this situation. The 25 m sublevels were cablebolted, with three cables per ring. AFT was less than a metre (approximately 4000 tonnes) with most of the fall-off occurring in the south end of the stope (up to 3 m). Generally, experience at Golden Grove suggests that where an intermediate sublevel has been effectively cable bolted, ‘subspan’ design can be used. A crude relationship relating overall (supported) span to unsupported subspan has been developed by Nickson, 1992 (Hutchinson and Diederichs, 1996). This is normally only applicable to relatively dense cable bolt patterns with typically four - six cables per ring. As this stope had only three cables per ring, leaving it on the ‘borderline’ for subspan design, the modified stability number (N’) was also calculated for the designed combined stope hangingwall and plotted on the stability graph in Figure 5. Using this method the stope plotted in the caving/unstable zone.
435z57/460z55 This was also a multiple lift stope, which had minimal dilution with an average of 0.84 m hangingwall fall-off thickness (approximately 8800 tonnes). On the Potvin and Nickson stability graph it plotted in the transition zone. 130 160 tonnes of ore was recovered, and the final CMS volume of the stope was 30 680 m3.
APPLICATION TO OTHER OPERATIONS CMS surveys are now widely used to assess stope performance in Australia (eg Gilbertson, 1995; Rauert, 1995). At operations where stope surveys are routinely undertaken, the development and application of a site specific stability graph involves the following steps:
• Compare stope void survey results with design. This is usually done by ‘slicing’ the CMS model and presenting the survey outline on the original ring sections.
• Calculate the Average Fall-off Thickness (AFT). • From the geotechnical database, calculate median Q’ values FIG 6 - Scuddles stopes plotted on ‘ELOS’ stability graph.
A dacite dyke, 2 - 6 m thick, ran almost parallel to the hangingwall, and the stope hangingwall failed to the contact of, or through to the centre of the dyke. The fall-off was mainly from the centre to the south end of the stope, consistent with the dyke, where as much as 5 m fell off the hangingwall. During stoping, a large mass of about 2000 tonnes fell from the hangingwall close to the brow. 63 200 tonnes of ore were recovered, and the final CMS volume of the stope was 15 936 m3.
555z39 Major hangingwall failure occurred in this stope, with up to 3.5 m falling from the hangingwall (approximately 46 400 tonnes). It plotted in the caving/unstable zone of Potvin and Nickson’s stability graph, and failure could have been predicted using this method. 129 300 tonnes of ore was recovered, and the final CMS volume of the stope was 39 714 m3.
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for appropriate regions of the stope hangingwall, eg Q’0-5 m and Q’5-10 m. These are then used to calculate N’ values for each stope hangingwall.
• Plot N’ values against Hydraulic Radius on a ‘Stability Graph’. Different symbols should be used for each point to represent various categories of AFT. Once a significant number of stopes have been assessed the resulting plot can be subdivided into regions based on various levels of AFT. This then forms the design chart against which future stopes can be assessed. The potential benefits of alternative designs or support can be evaluated in terms of increased or reduced dilution. The chart can also be used to develop initial stope designs in new stoping areas in nearby projects. Obviously, caution must be applied where there are differences in the geology or other factors such as the amount of structure. However, if conditions are broadly similar, a locally developed chart may provide better guidance than the more general ‘published’ databases. Stoping in the recently developed ‘C Panel’ at Gossan Hill has been extremely successful. The designs were assessed against the experience gained in the Scuddles orebodies.
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The database developed for Scuddles is still relatively small and the stability/AFT regions are currently rather poorly defined. An on-going effort will be necessary to develop a chart that can be used confidently for design.
CONCLUSIONS • The Stability Graph Method is applicable to Scuddles. Although the database is still very small, stopes not affected by ‘unusual’ factors such as undercutting sublevel development, or adverse geological features, tend to plot in well-defined regions.
• Stopes that would be classified as ‘transitional’ using the published Stability Graph (Hutchinson and Diederichs, 1996) experienced minimal dilution ( 10 (Figure 1). In contrast, 50 per cent of the operations had a hanging rock mass quality of fair or worse, with a modified NGI classification Q’ < 4 (Figure 2). The relatively weak wall rocks necessitate relatively small open stopes. For many mines, the typical open stope varies in size from 20 000 tonnes to 100 000 tonnes. Moss et al, (1992) makes the same conclusion in a summary of the ground conditions in ten mines in Northern Manitoba.
FIG 1 - Orebody rock mass quality from a survey of 34 Canadian open stope mines (Potvin and Hudyma, 1989). More than 85 per cent of the mines had an orebody rock mass quality of good or better.
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FIG 2 - Hanging wall rock mass quality from a survey of 34 Canadian open stope mines (Potvin and Hudyma, 1989). More than 50 per cent of the mines have fair, poor or very poor hanging wall rock mass conditions.
Typically, the orebodies consist of massive sulphides encased in thin schist wall rocks… This juxtaposition of strong competent ore rocks with highly anisotropic poor quality hanging wall and footwall rocks has led to a number of specific ground control problems….
Stoping direction The implications of stoping direction can have a large impact on the development needed to exploit an orebody. Transverse mining dictates that extensive waste development is needed to access open stopes (Figure 3). If longitudinal mining can be successfully employed, the majority of development needed can potentially be driven in ore (Figure 4). The stoping direction is largely dictated by width of the orebody, however, rock mass quality also plays a role in determining the stoping direction. Figure 5 shows the longitudinal mining width compared against orebody rock mass quality (Potvin and Hudyma, 1989). If the rock mass quality of the ore is fair (Q’ < 10) longitudinal mining has been employed in mining widths up to 15 metres, however, 75 per cent of the operations had a mining width of ten metres or less. If the rock mass quality is good (Q’ > 10 and Q’ < 40), longitudinal mining was employed in mining widths of up to about 20 metres. It is worthy of note that some operations reported back instability and occasionally failures at mining widths greater than 15 metres. It is interpreted that longitudinal mining in orebody widths larger than 15 m has an increased risk of instability and rock falls. There were very few operations that employed longitudinal mining in orebodies wider than 20 metres. Transverse open stoping is utilised in orebodies with orebody rock mass qualities varying from poor to very good (Potvin and Hudyma, 1989). However, there were no instances in which transverse mining was used in mining widths less than 15 metres (Figure 6).
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Stoping parameters With the size of open stopes limited by weak wall rocks, the opportunity for large multiple lift open stopes is uncommon. Traditional open stoping, in which there are multiple lifts, and one specifically designed drawpoint horizon on the bottom have been employed in the past in large operations such as Kidd Creek (Bedford, 1981) and Geco (Schwartz, 1978). Figure 7 shows a traditional multiple lift open stoping operation. However, in the 1980s and 1990s the trend in Canadian open stope mines is ‘short’ stopes (single lift or double lift) with fast turnaround times eg Williams mine (Bronkhorst et al, 1993), Kidd mine (Tannant et al, 1998), and Golden Giant (Bawden and Reipas, 1989). In the smaller orebodies, it could be argued that the capital cost for the development required for multiple lift stopes cannot be justified if similar production rates can be achieved from smaller stopes. Many larger mines, which have the infrastructure development in place for multiple lift stopes, have nevertheless made a transition to ‘short’ stopes. The large stopes generally have higher risk of uncontrolled failure than small stopes, and the larger the stope, the more difficult it is to arrest the failure. Scheduling large stopes can be a challenge, with all aspects of the mining cycle taking longer. A longer primary stope cycle time means that mining of adjacent ore reserves are delayed until the primary stope is mined, filled and the fill has set. The trend to ‘short’ stopes has some profound operational consequences. One of these consequences is that after the first lift, stopes have a flat-bottom. There is little opportunity to ‘shape’ the draw area to increase the amount of conventional mucking. Remote mucking in flat-bottomed stopes can exceed 50 per cent of the total stope production. Small pillars could be left to remedy the flat-bottom situation, but these pillars would be unrecoverable and significant ore losses in high-grade orebodies are not desirable. Figure 8 (redrawn from Bronkhorst et al, 1993) illustrates a typical cross-section of a single lift stoping operation, with the resultant flat-bottom stopes.
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FIG 3 - Transverse blasthole open stoping requires significant waste development (after Potvin and Hudyma, 1989). The mining method allows for complex stope sequencing.
FIG 4 - Longitudinal open stope mining, or sublevel retreat (after Potvin and Hudyma, 1989). Much of the development necessary for this mining method can be kept in the orebody. This is a relatively uncommon open stope mining method in Canada.
Use of cablebolting Another consequence of ‘short’ lift stopes is the need to support stope backs at each lift to prevent significant failure. Systematic cablebolt support of stope backs is a very common practice, with most stope overcut development supported on about a two metre by two metre pattern of six to eight metre long, twin strand cablebolts. While there have been notable failures through cablebolted stope backs, for the most part, stope back cablebolting ensures the desired stability. Cable bolting of stope hanging walls is also a common procedure. Most hanging wall cablebolts are installed at the top drilling sublevel of the open stope. The purpose of these cablebolt is not to provide stability of the hanging wall, but rather
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to prevent any instability from propagating to the up dip stope (Figure 9). If a hanging wall failure occurs in a stope, normally it is very difficult to prevent this failure from propagating to the next lift. Pattern cablebolting of the drilling horizon can choke substantial failures and prevent the failure from propagating to other stopes. Figure 10 shows hanging wall cablebolt patterns used at Kidd mine to control stope hanging wall instability. Uniform cablebolting from a hanging wall drift, to provide pattern reinforcement of a hanging wall surface is not a common practice, but numerous operations have occasionally used this approach (Kidd mine, Tannant et al, 1998; Golden Giant, Anderson and Grebenc, 1995; Brunswick Mining, Garland, 1994). An example of uniform hanging wall cablebolting is shown in Figure 11. While this is an expensive practice because
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FIG 5 - If rock mass quality is fair (Q’ < 10), most longitudinal open stope mining has been limited to a mining width of ten metres (after Potvin and Hudyma, 1989). For good rock mass quality or better, longitudinal mining has been employed in mining widths of up to 20 metres. However, several operations reported back instability problems and back failures at widths greater than about 15 metres. There have been few operations that have successfully used open stoping in mining widths greater than 20 metres.
FIG 6 - Transverse open stoping is utilised in orebodies with rock mass qualities varying from poor to very good (after Potvin and Hudyma, 1989). However, there were no instances of transverse mining used in mining widths less than 15 metres.
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FIG 7 - Multiple lift open stoping (after Potvin and Hudyma, 1989). Longhole and blasthole versions of multiple lift open stoping were widely employed in Canada from the 1960s to the 1980s.
FIG 8 - A typical cross-section of a single lift open stoping mine (redrawn after Bronkhorst et al, 1993). The flat-bottom on the mucking level results in significant amounts of remote mucking, while the flat back at the top of each stope must be kept relatively stable.
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FIG 10 - Two hanging wall cablebolt pattern used at Kidd mine to prevent hanging wall failure. In this particular case, there is a weak fault several metres in the beyond the hanging wall (Tannant et al, 1998).
FIG 9 - Cablebolts are choking the hanging wall failures in the 56-735 and 54-735 stopes (redrawn from Tannant et al, 1998). Cavity surveys show that overbreak stops at the hanging wall cablebolts at the top of each open stope.
of the cost of the drift, it has been a successful approach because most of the hanging wall can be pre-reinforced and the cablebolt orientation is near optimum. An interesting variation on the full hanging wall drift is to use short stub drifts (Sands, 1990). The Winston Lake mine had a dilution problem from a weak, chert hanging wall. One of the
FIG 11 - Pattern cablebolting to control hanging wall instability at Golden Giant (Anderson and Grebenc, 1995).
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FIG 12 - Pattern cablebolting of the hanging wall at Winston Lake mine (Sands, 1990). Access to the orebody is from the hanging wall. Mining of these accesses was leading to instability and dilution in the weak, chert hanging wall. Short cablebolt stub drifts were developed from the access cross-cuts, several metres into the hanging wall of the orebody. The cablebolt reinforcement of the hanging wall was very successful in reducing dilution.
factors causing the hanging wall dilution was that the access to the open stopes was from the hanging wall. The mining of the cross-cut accesses was preconditioning the hanging wall. However, the hanging wall accesses were also an opportunity. Rather than drive a full-length cablebolt drift between access cross-cuts, two 12 m long airleg stubs were driven from cross-cuts parallel to the orebody at a distance of several metres into the hanging wall (Figure 12). Cablebolt fans were drilled from the stub drifts to achieve a two metre toe spacing between cables, and a two metre spacing between rings. This cablebolting was very effective in controlling hanging wall dilution at Winston Lake. While extensive hanging wall cablebolting seems cost prohibitive, it needs to be justified in terms of the real cost of waste dilution. Anderson and Grebenc (1995) documented the cost of dilution at the Golden Giant mine. The direct mining, milling and administration cost of handling the mine’s 13.6 per cent dilution was $5.4 million, or $29 per tonne of dilution (nine per cent of this dilution came from the hanging wall). The typical direct cost to mine, haul and process one tonne of rock is $20 to $40. Nickson (1992) found that the installed cost of cablebolting in Canadian mines varied from approximately $19/metre to $36/metre, with an average of $27/metre. Effectively, the cost per metre of cablebolt is roughly the same as the direct cost of one tonne of dilution. In other words, each metre of hanging wall cablebolting pays for itself if it eliminates one tonne of waste
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dilution. In 20 000 to 100 000 tonne stopes, 20 per cent hanging wall dilution is 4000 to 20 000 tonnes of dilution. There is plenty of opportunity for well-designed cablebolts to pay for themselves.
Dilution control Dilution control has become one of the most important ground control issue in Canadian mines. In a survey of Canadian mines in 1988, 40 per cent of the mines reported that dilution exceeded 20 per cent (Pakalnis, Poulin and Hadjigeorgiou, 1995). Moss et al (1992) highlights dilution control as one of the main ground control concern in all ten northern Manitoba mines. Moss’ conclusion is supported by Yao et al (1999) who performed a dilution study on 39 stopes in a northern Manitoba mine, and found an average of 23 per cent hanging wall overbreak. When you can measure what you are speaking about and express it in numbers, you know something about it; but when you cannot measure it, when you cannot express it in numbers, your knowledge is of a meagre and unsatisfactory kind. Lord Kelvin. The development of a laser stope survey system or ‘Cavity Monitoring System’ (CMS) was a major step in the quantification of open stope performance and identifying the location of dilution (Miller et al, 1992). Until that time, dilution estimates had been based on the tonnes drawn from a stope.
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TABLE 1 Stope extraction summary for one years’ production at Golden Giant (reproduced from Anderson and Grebenc, 1995). Stope
Planned Tonnes
Mined Tonnes
H/W Waste
F/W Waste
Backfill Mined
Over Break
450-Q5
75 025
67 786
873
1059
450-Q6
76 922
77 740
1715
450-Q7/8
154 114
153 119
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Ore Bench
% Dilution
% Recovery
% Overbre ak
2952
944
2408
9171
2.58
87.78
0.00
7201
7.29
90.64
1145
2320
3.13
45
9079
6.78
94.11
0.03
450-Q9
70 618
65 693
4926
267
1569
11 641
7.35
83.52
2.22
450-Q10
66 877
71 223
4105
2646
1441
4036
6314
12.25
90.56
6.03
450-Q11
62 636
68 525
6717
1791
12 991
1433
4052
34.32
93.53
2.29
440-H2
47 598
54 147
3888
2748
798
322
1206
15.62
97.47
0.68
440-H4
26 793
25 324
505
568
1753
473
10.55
100.00
1.77
453-2W
80 923
77 008
1483
440
2746
2.38
100.00
3.39
460-Q9/Q10
60 514
103 765
41 005
1790
1991
3366
70.72
94.44
3.29
450-Q13
56 934
61 088
6969
1407
570
1132
5057
15.71
91.12
1.99
450-Q12
58 176
65 944
11 390
654
16
4291
20.70
92.62
0.03
466-3/4W
116 584
146 203
17 988
4629
8905
3393
20.70
97.09
7.64
1521
466-0
68 694
60 532
1632
326
2753
2841
2.85
95.86
4.01
446-D1
166 698
201 454
2633
6475
27 157
1480
5.46
99.11
16.29
456-5W
91 785
115 192
6326
294
5417
450-8W
55 320
54 713
755
1837
985
460-Q11
21 327
26 745
1380
2689
582
1 357 538
1 496 201
121 272
29 955
8.93
2.21
Total % of planned
18 092
6760
13.11
92.63
19.71
4459
6.47
91.94
0.00
1041
960
21.81
95.50
4.88
33 084
74 119
81 271
13.58
94.01
5.46
2.44
5.46
5.99
While this is an indication of stope performance, the location of the overbreak could not be determined, and quantifying underbreak was at best an estimate. Through the use of a cavity monitoring system, Anderson and Grebenc (1995) were able to quantify stope performance at Golden Giant. Over a one-year period, there was 8.9 per cent hanging wall dilution, 2.2 per cent footwall dilution, 5.5 per cent overbreak into adjacent ore pillars, 2.4 per cent cemented fill dilution, and 6.0 per cent ore loss (Table 1). Measuring the problem is the first step towards mitigating it. Clark and Pakalnis (1997) find that the factors related to dilution include:
• irregular wall geometry; • undercutting walls (with mine development, or as a result of overbreak from an adjacent stope);
• blast hole deviation; • blast hole layout (fan drilling versus drilling parallel to stope walls);
• blast hole offset or standoff from the final stope wall; and • the number of blasts used to extract the stope. Stope dilution impacts a mining operation in several ways. There is a direct cost for mining the waste rock. There is the indirect cost of mining and processing valueless rock in lieu of ore, which in some cases also reduces concentrator recoveries. There is also an indirect cost associated with handling of the dilution, which is often in the form of oversize. Oversize and the additional material lower stope productivity and draw out the stope cycle time. Anderson and Grebenc (1995) quantify the effect of dilution on the planned mucking time of the stope (Figure 13). Increased mucking time means that other stopes must be brought online early in order to meet production targets.
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FIG 13 - Variance to the planned mucking time as a result of dilution (redrawn from Anderson and Grebenc, 1995). The graph demonstrates that with increased dilution, the planned mucking time increases at rate of 2.5 times. This increased rate is a result of the slow mining rate for oversize dilution.
MINE SEQUENCING Bottom-up mining There are a number of operational, scheduling and economic reasons for mining an orebody from bottom to top. From a
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geotechnical point of view, bottom up mining is an effective means of stress management. When mining in high stress conditions, ideally extraction progresses from the bottom of a mining block towards its top. As the total extraction increases and the stress concentrates, the extraction horizon moves towards the shallower levels of the mine and towards the areas of lower premining stresses. As a result, excessive induced stress and deteriorating ground conditions are better managed. Since many of the Canadian operating mines are now deep (by Australian standards) and have a high stress environment, stress management is often a key geotechnical design consideration. Deep mining usually make shaft access more productive and more economically attractive than declines, especially if extraction is to commence at the lower levels of the mine. Consequently, Canadian mine infrastructure typically includes a hoisting and man and supply shaft, and an underground crusher, connected to a network of ore passes and waste passes, through a haulage level. This significantly differs from most Western Australian top-down mining operations with decline access. Bottom-up shaft mining requires an intensive capital program and may delay initial production by well over one year. Ultimately, it will minimise stress related problems as extraction progresses.
In the case of very large Canadian open stope mines (Kidd, Creighton, Brunswick), the mines have incrementally deepened over time. This gives an overall top down mining direction, but individual stoping blocks are extracted in a bottom up sequence (Figure 14).
Stoping sequence The stoping sequence is driven primarily by the ore grade (high-grade first), but also by operational issues (access, ventilation, filling, etc) and by rock mechanics considerations (stress management). Producing the best possible grade early in the project, along with consistent tonnage output throughout the project life are fundamentals of a successful underground mining operation. Consistent tonnage is only achieved if ground conditions are controlled and geotechnical risks are well managed. The overall stope sequencing strategy followed by many Canadian open stope mines is influenced by the high stress conditions of deep mining. The general retreat sequence is from the central portion of the orebody towards the extremities. This gradually pushes induced stresses into the abutments. It also avoids concentrating stresses in remnant pillars.
FIG 14 - Incremental deepening of a large Canadian mine (Hudyma et al, 1995). The levels from 575 metres to surface were mined first. As mining extracted the upper levels new main levels were established. In this case, a new, deeper main level was established every 125 to 150 metres deeper, every several years. New shafts were sunk and then deepened to access the lower levels. There is no decline from surface for this mine.
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FIG 15 - Pyramid retreat as practiced at Golden Giant (redrawn from Bawden and Reipas, 1989). The mining sequence starts in the middle of the block and expands laterally and vertically. Virtually no rib pillars are developed in this mining sequence. Operationally, this is a very development intensive open stoping sequence.
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FIG 16 - The Williams mine uses a version of the pyramid retreat. Rib pillars are created, but they are mined relatively quickly after primary stopes. A rule of thumb is never mine more than two sublevels ahead of a pillar before recovering it. This mining sequence was producing 1.0 million tonnes per year, from each stoping block (Bronkhorst et al, 1993).
FIG 17 - Longitudinal blasthole open stoping showing part of a 1-4-7 mining sequence (after Potvin and Hudyma, 1989).
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FIG 18 - The 1-5-9 stoping sequence at various stages (after Potvin and Hudyma, 1989). Stopes 1, 5, and 9 are lead stopes and kept one or two lifts ahead of stopes 3-7-11, which are also primary stopes. The pillars (even numbered stopes) are extracted one or two lifts behind the 3-7-11 stopes.
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The ideal stress management stoping sequence is a systematic retreat from the centre-out, without pillars. This pillarless mining sequence was applied in Block 3, at the Golden Giant mine and named pyramid retreat, as mining progresses in a triangular shape. The Golden Giant pyramid retreat sequence is illustrated in Figure 15. Although conceptually sound, the capacity of this mining system to produce ore is severely constrained by the stope cycle time. Stopes must be mined, filled and cured before the adjacent stope can be mined. With active mining on a large number of sublevels, there are substantial development, scheduling and logistic challenges. For example, to mine Stope #6 in the sequence (very early in the mine life), seven sub-levels must be developed and operational. Finally, all stopes must contain cemented fill, adding to operating cost. A compromise is to keep the general triangular retreat shape but using a primary and secondary stope arrangement. This system has been used at the Williams mine and is illustrated in Figure 16. This allows for a number of primary stopes to be mined simultaneously, increasing the productivity of the mining system. Also, since secondary stopes fill masses will never be exposed, unconsolidated fill can be used in almost half of the total volume to be filled. One of the underlying principles of this system is to recover pillars (secondary stopes) as early as possible in the sequence. Some basic sequencing rules have been learned with experience:
• never mine more than two sublevels ahead of a pillar before recovering it; and
• never mine on both side of a pillar at the same time in high stress. There are variations in primary/secondary sequencing. In low stress, mining every second stope can be done successfully (see Figure 3). In high stress, the second rule can be satisfied by mining primary stopes in a 1-4-7 sequence. Stopes 2-5-8 become secondary stopes with one fill wall exposure; while stopes 3-6-9 are tertiary stopes with fill walls on both sides (Figure 17 shows a part of 1-4-7 sequence). Another variation involves mining stopes 1-5-9 as primaries, then stopes 3-7-11 as primaries (pillar on both sides) and even numbered stopes are as tertiaries (fill on both sides). Figure 18 shows six steps to a 1-5-9 mining sequence. Some mines have designed small primary stopes and large secondary stopes in order to reduce the use of cement in the fill. When stress levels approaches rock mass strength, mining will be continuously done in failing ground conditions. The effect of time on the stability of excavations will become an important consideration in the design, planning and scheduling of activities. ‘Just in time’ mining will become a critical success factor. The overall process of mining, including development mining, stope preparation, mucking and filling cycles must be accurately planed and executed.
The final selection must be justified in terms of net present cost being offset by the value of additional ore recovery. Cemented rock fill systems were used extensively in Canadian open stope mines throughout the 1980s and 1990s. Most commonly, these systems involved underground cement batch mixing stations under which, trucks full of waste rock or quarried rock fill would drive, and a cement slurry sprayed over the rock fill load. The truck would then tip the cemented fill from the edge of empty open stopes. In free dumping, it should be noted that the cement slurry component typically segregates towards the stope wall under the dumping point. The backfill wall furthest from the dumping point often has a low cement content, which may result in stability problems when exposed. Ideally cemented rock fill is tipped from all of the sides of a stope that are going to be exposed. In Canada, it is not common to find a cemented rock fill system that tips down boreholes, as shown in the Williams mine sketch in Figure 8. Cement rock fill systems have high operating cost, with cement content often around five per cent and a dedicated fleet of mobile equipment. However, they offer fast curing, competent fill (commonly 2 to 4 MPa, albeit there may be a segregation problem), flexibility and low capital cost. However, the speed of filling is limited by the mobile equipment dedicated to this operation. The filling schedule can easily become a production bottleneck and the capacity to catch-up with this type of filling system is very limited. Many Canadian mines have rapidly taken advantage of new developments in paste fill technology. The Golden Giant, Westmin, Chimo, Macassa and Lupin mines are examples of operations that have changed filling systems in the mid-1990s from cemented rock fill and cemented hydraulic fill to paste fill. This trend is still continuing with Brunswick Mining, Kidd, INCO and Cambior operations. Other mines such as Louvicourt have been designed to operate with paste fill from the start of their mining operation. The attraction towards paste fill can be explained to some extent, by the environmental pressure existing in North America. There is a significant incentive to dispose of mines tailings underground. Furthermore, mines will commonly achieve reductions in operating costs due to:
• • • • •
potentially lower cement content; fewer operators; no mobile equipment maintenance; faster setting time than cemented hydraulic fill; use of the complete tailing fraction (rather than just the coarse fraction in cemented hydraulic fill); and
• almost no free water draining into the mine working. The capital cost of paste fill plant, boreholes and distribution system is usually high.
SUMMARY
Backfill Most open stope mines in Canada use backfill as a mean to achieve total recovery of their ore deposits. Backfill is also seen as essential to provide regional stability and local support in deep mining conditions. Backfill is required as a floor to work on in ‘short’ lift, bottom-up mining sequences. The selection of backfilling systems is driven by operational demands such as:
• • • •
fast curing; speed, reliability and flexibility of fill delivery; availability of low-cost material; and the capital cost of building the plant.
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The mining horizons in Canada are gradually moving deeper and into higher pre-mining stress fields. As a result, open stope mining has evolved during the last 30 years from large scale multiple lift stopes, to small and fast turn around single lift stopes. The smaller stopes have reduced the risk of rock mass instability at the cost of increased remote mucking and more intensive ground support systems. To combat high stresses, a number of mining sequences have been used to better ‘manage’ the induced stresses around stopes, extract pillars earlier and maximise ore recovery. The filling systems have also evolved to become more flexible, with faster delivery. It has also become a requirement for the fill to cure within a month. Consequently, cemented rock fill and more recently paste fill systems are now almost exclusively used in Canadian mines.
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REFERENCES Anderson, B and Grebenc, B, 1995. Controlling Dilution at the Golden Giant Mine, 1995 CIM Mine Operators Conference. Bawden, W F and Reipas, K, 1989. Stope Sequencing at the Golden Giant Mine, 9th Underground Operator’s Conference, Sudbury, Ontario, Feb 1989. Bedford, J E, 1981. Sublevel Stoping at Kidd Creek Mines, in Design and Operation of Caving and Sublevel Stoping Mines, (Ed: D Stewart), (American Institute of Mining, Metallurgical, and Petroleum Engineers, Inc). Bronkhorst, D, Rheault, J and Ley, G M M, 1993. Geotechnical principles governing mine design at Williams Mine, in Innovative Mine Design for the 21st Century, (Eds: Bawden and Archibald), (Balkema: Rotterdam). Clark, L and Pakalnis, R C, 1997. An Empirical Design Approach for Estimating Unplanned Dilution from Open Stope Hanging Walls and Footwalls, 99th CIM Annual General Meeting, Vancouver. Garland, K, 1994. Personal communication. Hudyma, M R, Milne, D and Grant, D, 1995. Geomechanics of sill pillar mining in rockburst prone conditions – Phase One Final Report: Sill pillar monitoring using conventional methods, Noranda Technology Centre, January 1995. Miller, F, Potvin, Y and Jacob, D, 1991. Laser Measurement of Open Stope Dilution, 93rd CIM Annual General Meeting, Vancouver.
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Moss, A, Reschke, T and Greer, G, 1992. Observations on Ground Behaviour in Manitoba Mines, 94th CIM Annual General Meeting, Montreal. Nickson, S D, 1992. Cable Support Guidelines for Underground Hard Rock Mine Operations. Unpublished MASc Thesis, University of British Columbia, 223 p. Pakalnis, R, Poulin, R and Hadjigeorgiou, J, 1995. Quantifying the cost of dilution in underground mines, Mining Engineering. Potvin, Y and Hudyma, M R, 1989. Open Stope Mining Practices in Canada. Presented at the 91st CIM Annual General Meeting, Quebec City, May. Sands, D, 1990. Cable Bolt Support at Winston Lake Mine. 59th Annual Meeting and Technical Sessions, Mines Accident Prevention Association Ontario. Schwartz, A, 1978. Pillar Recoveries Using Consolidated Fill at Noranda Mines Ltd, Manitouwadge, Ontario, in Mining with Backfill, 12th Canadian Rock Mechanics Symposium, (Canadian Institute of Mining and Metallurgy). Tannant, D D, Diederichs, M and Seldon, S, 1998. Hanging Wall Relaxation and Cablebolt Support for Deep Stopes at Kidd Mine, 100th CIM Annual General Meeting, Montreal. Yao, X, Allen, G and Willett, M, 1999. Dilution EvalutionUsing Cavity Monitoring System at HBMS – Trout Lake Mine, 101st CIM Annual General Meeting, Calgary.
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Open Stope Mining Strategies at Brunswick Mine B Simser1 and P Andrieux2 ABSTRACT By 1996 ground control-related problems at Brunswick Mine had raised serious concerns regarding the safe and economical extraction of large portions of the reserves. These difficulties were mainly caused by high stress levels in pillars, instabilities resulting from wide spans supported by ageing support elements, and chronic difficulties supplying rock fill from the surface quarry to the underground workings, which resulted in many large voids remaining open for long periods of time. Significant changes have been implemented at the site since 1996, amongst which were a change of mining method, the introduction of paste backfill, the establishment of stress management measures, and the introduction of shotcrete on a large-scale. This paper focuses on the regional ground conditions associated with both the previously used primary-secondary mining method and the newly implemented pyramidal pillarless technique, as well as on some of the stress management measures implemented at the mine.
INTRODUCTION The Noranda Brunswick Mine, previously referred to as the Brunswick Mining #12 mine, is a 3.5 million tonne per year underground base metals producer located in northeastern New Brunswick, Canada, that has been in continuous operation since 1964. Lead-zinc-silver-copper ore is mined from up to ten subparallel en-echelon massive ore-bearing sulphide lenses striking roughly North-South and dipping about 75° to the West. The deposit has an overall strike length of close to 1.2 km, and an overall width, all lenses included, of up to 200 metres. The maximum mining width in the widest ore lenses is close to 70 metres. The orebody is a vent-proximal massive sulphide deposit that formed above a feeder pipe by the venting of metal-laden hydrothermal fluids at the bottom of an ancient sea (Luff et al, 1992). Five successive phases of deformation have shaped the deposit to its current complex form. An intrusive porphyry dyke runs along the orebody, located mainly on its footwall side but cutting across it towards the south end of the mine and running along its hanging wall from there southwards on. Over 15 distinct geological horizons are encountered throughout the mine (Luff, 1977). The massive sulphides are hosted by metamorphosed volcaniclastic sediments and tuffs that overlie a sequence of felsic volcanic rocks. These units show a wide range of mechanical properties. The heavy, stiff and strong massive sulphides in which the ore is typically found have unconfined compressive strengths of about 210 MPa, and a density of 4300 kg/m3. The encasing metasediments, which are much lighter, softer and weaker, have unconfined compressive strengths ranging from 30 to 60 MPa, and densities from 2600 to 2900 kg/m3. Further, the older chloritic and sericitic footwall meta-sediments tend to be more competent than the younger hanging wall sequence of chloritic sedimentary units, which are typically very schistose with phyllitic partings, and often exhibit slickensided surfaces. The rock mass properties of the dyke are also quite variable, ranging from strong in massive horizons, to weak in narrower regions. 1.
Noranda Inc, Brunswick Mine, PO Box 3000, Bathurst, New Brunswick E2A 3Z8, Canada.
2.
Itasca Consulting Canada Inc, 1375–14 Regent Street, Sudbury, Ontario P3E 6K4, Canada.
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Geostructural characteristics throughout the mine are complex due to the presence numerous sets of structures. Of particular interest are the four distinct sets within the stiff sulphides (Godin, 1987). One of those sets is subhorizontal and tends to affect the stability of the stopes’ backs in some areas, while others are subvertical, some of which playing major roles with regard to seismic activity. Small and thin bands of weak metasediments are also locally sandwiched within the hard sulphides, further contributing to difficult ground conditions. Local rock mass strengths are quite variable depending on the frequency and orientation of structural features. The far-field major principal stress is subhorizontal and oriented East-West, perpendicular to the strike of the orebody. Its virgin gradient is 0.052 MPa per metre of depth. The far-field intermediate principal stress is also subhorizontal, but oriented North-South, with a virgin gradient of 0.044 MPa per metre of depth. The far-field minor principal stress is subvertical, with a virgin gradient of 0.028 MPa per metre of depth. At a depth of 1200 metres, the pre-mining major principal stress is thus in the vicinity of 60 MPa, while the mining-induced major principal stress, in the abutment of certain excavations, can locally reach as much as 150 MPa. Stress relief in the form of fracturing around the skin of excavations is commonly noted throughout the mine. A wide array of severe ground instability difficulties started to be encountered at the mine in the early-1990s, which culminated in 1995 and 1996 with the loss of entire active mining zones, seriously threatening the very viability of the operation. In 1998 alone, 1.2 million tonnes of scheduled ore were affected to some degree, from delayed to written-off, due to ground condition-related difficulties. This represented close to 35 per cent of the yearly production and created havoc on the operation as planning had to be quickly and significantly re-adjusted, and alternate horizons quickly brought into production. Numerous strategic and tactical alternate mining approaches had to be used to overcome these challenging conditions, and allow the mine to meet its commitments towards its clients and shareholders. This paper reviews the history of mining at Brunswick Mine, and presents selected case studies discussing some of the variations in stope design and mine sequencing strategies used over the past years.
HISTORY OF MINING AT BRUNSWICK MINE The mine was initially exploited by primary-secondary open stoping in the upper part of the orebody. Mining was switched in the early-1970s to mechanised cut-and-fill, starting at the bottom of the 575, 725 then 850 levels, at depths of 580, 715 and 850 metres respectively. As cut-and-fill mining progressed upwards and the sill pillars above the mining horizons started to get smaller, effectively pinching more stress in less ground, mining conditions became more difficult. Compounded with ever increasing stress levels were the very wide back spans inside the cut-and-fill stopes. The high exposure of personnel, constantly working under these wide and highly stressed backs due to the very entry nature of the cut-and-fill mining method, warranted a switch back towards less exposing primary-secondary open stoping towards the end of the 1980s. The 1000 Level was started with this method, while the remainders of the 850 and 725 levels were switched to it. At that time, very large stopes were designed, averaging 75 000 tonnes.
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Primary mining initially went well, but problems started to be encountered as extraction ratios increased and secondary pillars started to take significant loads. As mentioned, increasingly severe difficulties started to be encountered in the early-1990s. On one hand, natural factors — such as poor local geostructural conditions, sharp mechanical property contrasts between adjacent rock types and high in situ ground stresses — resulted in adverse general conditions. On the other hand, excessively large excavations, inadequate ground support systems, the cutting of numerous pillars in which ground stresses concentrated, and a number of mining sequences based upon sound production considerations but poorly adapted to the emerging geomechanics-related difficulties, resulted in an inadequate match between the mining approach and the prevailing rock mass conditions. Problems were further aggravated by chronic difficulties supplying rock fill to the mined-out primary stopes, which resulted in numerous large voids remaining open for extended periods of time (Andrieux and Simser, 2000). A number of specific problems were observed in the large open stopes. In the first place, the exposure of large spans of weak hanging walls metasediments caused gravity-driven instabilities that resulted in hanging wall overbreaks, which, in turn, caused added dilution, difficult mucking conditions and an increased risk of trapping remotely-controlled mucking equipment. Such failures typically undercutted the hanging wall of the next lift above, making it even less likely to remain stable when exposed. Unless tightly backfilled, such failure also had a strong potential to propagate sideways, thus affecting the recovery of the neighbouring secondary stopes. This type of difficulty was largely resolved by reducing the size of the stopes, ensuring development does not undercut weak walls (by keeping access drifts away from the weak contacts), and prompt backfilling. The 1998 switch from rockfill to paste backfill made a significant difference in that it resulted in a much more reliable delivery of fill to the underground workings, a much tighter placement that provided quicker support, and allowed to backfill otherwise difficult to access caved areas. Details of the conversion to paste backfill at Brunswick Mine can be found in Moerman et al, 1999 and Ouellette et al, 1999. Another stability problem encountered in the large open stopes was the result of high stress conditions and the associated seismic activity. This type of instability steadily increased as mining evolved, ie as the extraction ratio increased and deeper mining took place. Shake damage is the dominant mode of seismically-induced failure at Brunswick Mine, even though rock bursting, or damage occurring in the form of violent expulsion of large volumes of rock, does also occur, but to a much lesser extent (Simser, 2000). Seismicity almost always locates within the hard and brittle sulphide horizons — many different mechanisms are at play, the principal being high stress concentration in pillars, high stiffness contrast between adjacent rock types, and slippage along discrete structural features (Simser and Andrieux, 1999). With regard to the later mechanism, a particular set of discontinuities called the ‘A-shears’, which are subvertical and striking roughly NW-SE, is often associated with seismic activity. Seismicity became a serious concern in the mid-1980s, and a first seismic monitoring system was installed in 1986. Successive system upgrades took place as seismicity got worse, and in 1995 a state-of-the-art ISS microseismic monitoring network was commissioned (Hudyma, 1995). The current network configuration has 90 channels and an overall sensitivity of about local Richter Magnitude –2.8. Current seismicity at Brunswick Mine remains high, 24596 seismic events were recorded in 1999 alone, of which 283 were above a Local Richter magnitude -0.2. Every single seismic event recorded is manually processed to both fine-tune location and provide reliable quantitative source parameter data (Simser, 1996), which, in turn, can be used to help maintaining safe mining conditions, understanding the reaction of the rock mass to
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mining, and assist in the design of future mining (Simser et al, 1998). Stress-related instability is however not necessarily associated with seismicity. High ground stresses in weak and soft material usually lead to aseismic squeezing conditions, whereby the rock progressively deforms and fails — this type of failure can be very difficult to control, and the various ground support systems tried to date at the Brunswick Mine under such circumstances have had limited success. Decreasing stress levels, which usually occur as part of a stress cycle (when a region is first subjected to high stress levels while it sits in the abutment of active mining, followed by a stress decrease as mining has progressed past it and put it in a stress shadow, effectively taking away the clamping forces that held the strata together) can also lead to very unstable ground conditions. As the amplitude of the stress cycle increases and the high stress component further breaks up the rock mass prior to it being declamped by the subsequent stress decrease, the likelihood of instability increases too. Highly laminated rock masses are also prone to this kind of failure, even with limited stress drops. At the Brunswick Mine, this type of instability can usually be successfully controlled with adequate ground support, such as long bars, cablebolts and/or shotcrete, depending on the spans involved and the severity of the situation. By mid-1996 ground control-related difficulties had reached the point where the operation was having difficulties maintaining its scheduled production. A number of measures have been implemented at Brunswick Mine since mid-1996 in an attempt to turn the situation around. Firstly, a drastic change of extraction method from primary-secondary to pillarless pyramidal mining has been completed. This has required a significant adjustment at all levels of the operation due to the much more complex sequencing and much reduced operational flexibility associated with it. Secondly, the average stope size has been significantly reduced from 75 000 to about 40 000 tonnes. Stopes are also individually designed — in terms of final size, geometry, orientation, access and ground support — to suit the local conditions. Thirdly, stress management measures have been established, mainly in the form of adapted mining sequences, and, in one instance so far, of a large panel destressing blast (Brummer et al, 2000). Fourthly and as previously mentioned, paste backfill has completely replaced rockfill, which has had a significant impact. Finally, shotcrete has been introduced on a very large-scale at the operation as a remarkably effective ground support system in an array of situations. By the end of 1999, a total of 60 570 tonnes of shotcrete had been applied underground for ground control purposes.
FIRST CASE STUDY — PENDANT PILLAR IN THE SOUTH END OF THE 850 LEVEL This first case study helps illustrate why pillarless pyramidal mining was contemplated as a viable alternative to primary-secondary extraction. There are many examples at Brunswick Mine of highly stressed remnant/pendant pillars that have become very difficult, if not impossible, to safely and economically recover. One of the main problems with those is seismic activity, which can easily deny access to working areas, either by damaging openings or by preventing access due to concerns of possible sudden ground failures. Relatively squat ore pillars at Brunswick Mine generally do not fail in a timely fashion and can remain seismically active for extended periods of time. The transition from ‘squat’ to ‘slender’ pillars has not been comprehensively studied at the mine, but generally occurs between width-to-height ratios of 0.5 to 1.0. A ‘squat’ pillar’s ability to maintain load over long time periods creates a seismic hazard and the uncertainty of production that comes with it. Whereas a pillar that fractures sufficiently to effectively eliminate the seismicity can have more predictable behavior, albeit with the increased risk of gravity type failures.
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Figure 1 shows such a ‘squat’ pendant pillar, with a width-to-height ratio of 2:1, that has remained seismically active over several years due to mining-induced regional stress redistribution. This pillar, located between the 69 and 70 stopes from 850-3 sub to 850-2 sub, was cut in 1991 with approximate dimensions of 20 m along strike (North–South), 10 m East–West, and 33 m vertically. Figure 2 shows the south end of the mine in a longitudinal section looking west. As previously mentioned, the major principal stress is subhorizontal and oriented East–West, perpendicular to the strike of the orebody. The shaded areas correspond to open stopes mined-out between 850-3 sub and 850-2 sub (above the elevation of the plan view shown) — extraction was also complete below the 850-2 sub, but in the form of cut-and-fill stopes. The plan is shown three times side-by-side, once to display the seismicity recorded between 1 March 1995 and 31 December 1996, once to show the seismicity recorded between 1 January 1997 and 31 December 1998, and once to display the seismicity recorded between 1 January 1999 and 22 August 1999. Mining activity in the south end of the mine, both above (on the 725 Level and as far away as the 575 Level) and below (on the 1000 Level) this region, still triggers occasional seismic activity in this pendant pillar. The pillar clearly has not yet failed despite its relatively small size and the cumulative damage it has sustained over the past nine years. Such pillars are obviously very difficult and costly to extract, and prevent any reliable production from being planned out of them. The pillar shown in this example has been abandoned and its reserves written off.
the resulting secondary pillars between them were created from the early-1990s up to 1996, when mining in the region had to be suspended due to massive bouts of seismicity and the associated large-scale ground failures. The slender 348-8 pillar between the mined-out 347-8 and 349-8 stopes has obviously failed as no seismicity is occurring in it — this pillar had a planned width-to-height ratio of 0.43:1 (30 m N–S and 70 m E–W). The #8 Lens in this region narrows from South to North, which is why the other secondary pillars (the 350-8, 352-8 and 354-8 going north) had increasing planned width-to-height ratios of 0.67:1, 0.75:1, and 1.2:1, respectively. As can be seen on Figure 3, as the width-to-height ratio of the secondary pillars increases so does the strength of the pillars, and the levels and duration of the seismic activity in them. Extraction has now resumed in these secondary pillars, but at a relatively slow mining rate, and following the extensive and costly ground rehabilitation work rendered necessary by the seismically-induced ground failures of July 1996. It is difficult to reliably plan consistent production from such regions, as seismically-induced ground failures can put mining on hold at any time, and prevent further mining until the required rehabilitation work is completed. Mining activity above this particular region, in the north end of the 850 Level, also often produces seismicity in it as a result of stress redistribution. Figure 4 shows a longitudinal section of the same area with events plotted from 19 August to mid-day on 21 August 2000. Blasting in the 850 block triggered the seismic activity shown the sill pillar between the 349-8 and 351-8 stopes.
SECOND CASE STUDY — NORTH END OF THE 1000 LEVEL
THIRD CASE STUDY — PILLARLESS PYRAMIDAL MINING ON 850 NORTH
The second case study also illustrates why pillarless pyramidal mining was contemplated, as secondary pillars at Brunswick Mine have shown similar trends as the pendant pillar discussed above. Figure 3 shows a plan view of the north side of the 1000-3 sub, at a depth of about 900 m below surface, with the seismic activity recorded locally in January 2000, which was typical of the event rate for the area. The circled areas correspond to primary stopes mined-out between 1000-3 sub and 1000-2 sub ¾
The 850 North block sits above the 1000 North end discussed in the previous case study (See Figure 4). It extends from approximately 700 to 850 metres below surface; with the bottom 40 to 50 m mined-out using cut-and-fill extraction. The block was subdivided in four sublevels with one sill pillar at the bottom, below the lower cut-and-fill mining, and one at the top, above the 850-4 subelevation. The zone has two principal ore
70 stope 69 stope
N sigma1
FIG 1 - Seismicity associated with the pendant pillar between the 69 and 70 stopes on the 850 2 sublevel. Grid spacing is 100 m. Events shown in the left figure are from the period March 1995 to end of 1996, middle figure January 1997 to end of 1998, right figure January 1999 to August 1999. Depth below surface is 810 m.
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69/70 pillar N FIG 2 - Longitudinal view looking west of the 850 south ore zone. Seismic events plotted as circles scaled to magnitude for July 1999. A total of 2377 events were recorded in the area shown, 74 events occurred in the pendent pillar. Elevations shown are in mine coordinates, surface elevation is 2652, co-ordinates increasing upwards (metres).
FIG 3 - Plan view of the north side of the 1000-3 sub showing the seismic activity recorded in January 2000. Grid spacing is 50 m. Orebody steeply dipping to the west, note hangingwall development is part of ventilation exhaust system, not for stope access.
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FIG 4 - Longitudinal view looking west of the 850 and 1000 North Main Ore Zones. Seismic events shown from 19 August to 21 August 2000. A 5000 tonne blast in the 850 block triggered seismicity in the immediate surrounds and in the 1000 block sill pillar below.
horizons, a relatively narrow hanging wall lens (10 to 15 m wide hanging wall to footwall), and a wide portion comprising the main ore zone (up to 60 m wide hanging wall to footwall). The mining plan in the wide zone was designed to be extracted in a pillarless pyramidal sequence, the south side of the pyramid being partially stress shadowed by historic hanging wall lens mining, and the north side being a relatively highly stressed abutment area. The pyramid stoping began in late-1996 and has gone with relatively few problems to date. Figure 5 is a longitudinal view looking West of a three-dimensional block model of the 850, 1000 and 1125 north zones. The seismic events recorded in 1999 are shown as small circles, and the stopes extracted during that period are those labeled ‘99. The heavy seismic activity associated with the primary-secondary mining in the 1000 North is visible below the pyramid. A total of 11502 seismic events are displayed. As seen in the upper right part of this figure, the north side of the pyramid has had a much higher seismic event rate than the south side — seismic source parameter information also indicates that the north side is higher stressed than the south side. As mentioned, very few mining delays have been encountered in this mining region, partly due to the mining sequence retained whereby the hanging wall side of the pyramid is mined first, as a lead front, thus shadowing the bulk of the reserves behind. Also, seismicity has been observed to typically locate about one panel away from the lead stope due to stress-induced fracturing of the immediate abutment region. Seismic activity has also been observed to spread out throughout the wide portion of the ore lens, indicating that the load redistribution due to mining is also spread over a wide area. Event clustering and apparent stress are tools commonly used at Brunswick Mine to indicate highly stressed areas (Simser, 1996). Apparent stress (Mendecki et al, 1994) is defined as the ratio of seismic energy over seismic moment, times the modulus of rigidity, which is an elastic constant. High ground stresses in relatively brittle rock will typically generate high apparent stress
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seismic events. Figure 6 shows a plan view of the 850-3 sub with the seismic events recorded from January 1999 to June 2000, 30 m above and below the #3 sublevel elevation. Events are depicted as circles scaled to local Richter magnitude (only those events above a Richter magnitude –1.5 are displayed, for added clarity). The largest event in the data set was a local Richter magnitude 0.9, which had a seismic energy of 157 kJ, whereas the largest events in the mine are in the order of local Richter magnitude 2.5 with seismic energies of 107 joules. The north side of the pyramid has been able to release energy in a relatively stable manner over time, which has contributed to the lack of significant ground control problems there. The south side of the pyramid is clearly subjected to lower ground stresses, due in part to the West Ore Zone shadowing it, and to fracturing caused by historic mining to the south. This side has also had relatively few ground control problems so far. Figure 7 shows the cumulative number of events recorded between the beginning of 1999 and mid-June 2000, in both the south side and north side of the 850 North pyramid stopes. The significant increase in the event rate for the south side is due to local mining and an overall shrinking pillar size. Despite this, the north side has remained much more active throughout the 18-month data set shown. Looking at the apparent stress levels associated with the seismic events provides further insight. Figures 8 and 9 show these data as a function of time, for the same time period as in Figure 7, for both the south side and the north side of the 850 North pyramid. These figures clearly show how the seismicity on the north side is of much higher apparent stress than on the south side. As mentioned, events on the north side of the pyramid tend to locate approximately one panel away from the active lead stope. This can vary locally due to geological structures, such as contacts, joints, and small shear zones, but generally holds true, mainly due to the fact that the stopes immediately abutting the pyramid tend to fracture due to the stress concentration, shedding
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FIG 5 - Longitudinal view looking West of a three-dimensional block model of the 850, 1000 and 1125 north zones. The seismic events recorded in 1999 are shown as small circles, and the stopes extracted during that period are those labelled 99’.
FIG 6 - Plan view of the north end of the 850-3 sub showing the seismic events recorded from January 1999 to June 2000, 30 m above and below the 850-3 sub elevation. The circles represent recorded seismic events above a –1.5 Richter magnitude.
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05/01/2000
03/01/2000
01/01/2000
11/01/1999
09/01/1999
07/01/1999
05/01/1999
03/01/1999
No. Events 05/18/2000
03/18/2000
01/18/2000
11/18/1999
09/18/1999
07/18/1999
05/18/1999
03/18/1999
3000 2500 2000 1500 1000 500 0
01/01/1999
Cummulative No Events North Side of 850 Pyramid
160 140 120 100 80 60 40 20 0 01/18/1999
No. Events
Cummulative No Events South Side of 850 Pyramid
Date
Date
FIG 7 - Cumulative number of events recorded for the south and north sides of the 850 North pyramid between the beginning of 1999 and mid-June 2000.
South Side of 850 Pyramid Stopes Apparent Stress (Pa)
100000 90000 80000 70000 60000 50000 40000 30000 20000 10000 0 01/01/1999
04/11/1999
07/20/1999
10/28/1999
02/05/2000
05/15/2000
Date FIG 8 - High stress seismic events from the south side of the 850 North pyramid.
load to the next panels. Underground observations and three-dimensional inelastic modeling (Pierce et al, 1999) corroborate this effect. This lag between the mining activity and seismicity is a key aspect of pyramidal mining: pillarless pyramidal mining being usually implemented in high stress areas, it is usually associated with significant seismicity — this seismicity however usually occurs ahead of the mining front. Figure 10 illustrates this point. It shows a plan view of the 850-3 sub on the north side of the 850 North pyramid with the seismic events recorded over a Richter magnitude –1.5 during the month of June 2000. As seen on the figure, the un-mined 399-1 Stope showed very little seismic activity as it had been fractured by previous mining and was by now partially stress-shadowed by the mining along the hanging wall side of the pyramid. Also, no seismicity can be seen in the immediate abutment of the mined out region — most of it but can however clearly be seen east of this stope, about one stope span away.
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This case study illustrates to some extent the range of responses experienced at Brunswick Mine to pillarless pyramidal mining. The behaviour on the north side is unfortunately more typical, but, even though harsher than on the stress-shadowed south side, is still much more manageable than in many secondary and tertiary pillars throughout the mine This case study also illustrates well how advantageous implementing stress shadow strategies can be, ie mining certain areas first in order for them to stress shadow other regions. This is the reason the mining of the West Ore Zone at Brunswick Mine has been a top priority since 1992, as it provides easier subsequent mining in the much larger Main Ore Zone next to it.
CONCLUSIONS The three case studies presented show the type of difficulties Brunswick Mine was confronted to in the later stages of primary-secondary mining and how pillarless pyramidal mining has been associated with less ground control difficulties. On the
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North Side of 850 Pyramid Stopes Apparent Stress (Pa)
100000 90000 80000 70000 60000 50000 40000 30000 20000 10000 0 01/01/1999
04/11/1999
07/20/1999
10/28/1999
02/05/2000
05/15/2000
Date FIG 9 - High stress seismic events from the north side of the 850 North pyramid.
FIG 10 - Seismic events from June 2000. Only the events over a Richter magnitude –2.0 are shown for claritiy.
down side, this method requires extremely precise engineering and production procedures as it is not flexible, nor forgiving. This aspect can be alleviated to some extent by mining a number of pyramids simultaneously, in order to balance production. However multiple pyramid sequences advancing both vertically
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and on strike will eventually create a pillar which will offset some of the stress management benefits. Numerous other measures and guidelines have been implemented at the mine to further counter the ground control problems. For example, mining is as much as possible done in
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bad ground first, as if it is left for the end, it is usually not recoverable due to excessive mining-induced damage. Also, the ore horizons are not completely silled-out anymore, but developed drift-width. The cost of the extra drilling due to the fanned patterns and the less ideal blasthole orientation along the stope walls are largely outweighed by the much more stable back conditions during the extraction cycle. In the same line of thought, openings are kept as small as possible since the depth of failure is proportional to the span. When large spans are unavoidable, cablebolts combined with shotcrete pillars have been quite effective, providing the ground does not unravel around them. Ground support has been a crucial component of the ground control program at Brunswick Mine. As mentioned, shotcrete and cablebolting are used on a very large-scale. Hanging wall cablebolting has also proven very effective at controlling hanging wall overbreak under certain circumstances. The critical importance of adequate mine wide monitoring cannot be over-stated. Seismic monitoring, on one hand, has become an intrinsic part of mining at Brunswick. It is used to maintain safe mining conditions, understand the response of the rock mass to various mining approaches and identify in advance seismically-prone regions (thus allowing to adapt the mining sequence accordingly and to install appropriate ground support). Non-seismic monitoring techniques, such as GMMs and extensometers, are also very useful, particularly when their data are available in real time. This is achieved at Brunswick Mine by wiring these instruments to surface via the microseismic system communication network. Conditions in the mine have dramatically changed over time, largely due to the ever-increasing extraction ratio. Changes to the mining method, support practices, monitoring, and backfill method have all been necessary to ensure continued safe production.
ACKNOWLEDGEMENTS The authors would like to thank the organisers of the MassMin 2000 conference for the opportunity to publish this paper, and the management of Brunswick Mine for granting permission to publish these data.
Brummer, R, Mortazavi, A, Andrieux, P and Simser, B, 2000. Destress Blasting — A Field Trial at Brunswick Mine. Report from Itasca Consulting Canada and Brunswick Mine to the CAMIRO Mining Division. Sudbury, Ontario, Canada. Godin, R, 1987. Structural Model of Brunswick #12 Mines as an Aid to Mining and Exploration. Annual Meeting of the CIM’s New Brunswick Branch. Bathurst, N B, Canada. Hudyma, M, 1995. Seismicity at Brunswick Mining, 10th Ground Control Colloque of the Quebec Mining Association, Val d’Or, Quebec, Canada. Luff, W, 1977. Geology of Brunswick #12 Mine, CIM Bulletin, 70(782):109-119. Luff, W, Goodfellow, W and Juras, S, 1992. Evidence of a Feeder Pipe and Associated Alteration at the Brunswick No 12 Massive Sulphides Deposit, Exploration and Mining Geology, 1(2):167-185. Mendecki, A, 1994. Guide to Seismic Monitoring in Mines, ISS International, Welkom, Republic of South Africa. Moerman, A, Rogers, K and Cooper, M, 1999. Paste Backfill at Brunswick - Part I: Technical Issues in Implementation, 14th CIM Underground Operators’ Conference. Bathurst, NB, Canada. Ouellette, G, Moerman, A and Rogers, K, 1999. Paste Backfill at Brunswick - Part II: Underground Construction and Implementation, 14th CIM Underground Operators’ Conference, Bathurst, NB, Canada. Pierce, M, Board, M and Brummer, R, 1999. 3DEC Modelling of Alternative Sequences for the 1000/1125 Block at Brunswick Mine, Itasca Consulting Group technical report to Brunswick Mine. Minneapolis, MN, USA. Simser, B, 1996. Seismic Monitoring at Brunswick Mine, ISS International Seminar on Advances in Seismic Monitoring, Western Deep Levels Mine Village, Carletonville, Republic of South Africa. Simser, B, 2000. Numerical Modelling and Seismic Analysis of Events Leading-up to a Violent Wall Burst, ISS International Seminar on Modelling with Data, Stellenbosch, Republic of South Africa. Simser, B, Andrieux, P, Peterson, D, MacDonald, T and Alcott J, 1998. Advanced Monitoring and Analysis of Microseismic Activity as an Aid to Mining at Brunswick Mines, 3rd North American Rock Mechanics Symposium (NARMS 3), Cancun, Quintana Roo, Mexico. Simser, B and Andrieux, P, 1999. Seismic Source Mechanisms at the Brunswick Mine, 14th CIM Underground Operators’ Conference, Bathurst, NB, Canada.
REFERENCES Andrieux, P and Simser B, 2000. Ground Stability-Based Mine Design Guidelines at Brunswick Mine, Underground Mining Methods Handbook, 3rd Edition (Ed: Hustrulid) (Sociey of Mining Engineers: Littleton, Colorado USA).
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Evolution of Vertical Crater Retreat Mining at Mindola Mine, Zambia E K Chanda1 and C Katonga2 ABSTRACT Mindola mine is currently the deepest mine in the Zambian copper belt. Stoping is currently active to 1330 m below surface with long range plans indicating future production from 1590 m level. The orebody is tabular o and dips between 55 to vertical. The thickness varies from 8 to 14 m. The vertical crater retreat method of mining was first introduced at Mindola mine in the late-1980s following several trials that helped define the design criteria for its application. Over the last ten years the method has been continually modified to improve safety and maximise ore production. For example, due to the weakness of some rock units within the orebody formation and the instability of the hanging wall, the last VCR slices are usually between 8 - 10 m thick, and efforts are being made to strictly apply shrinkage method in order to reduce dilution. Recent advances in mining technology in the areas of equipment, blasting techniques, rock mechanics and materials handling have been employed to increase the overall productivity of the mine. Large down-the-hole machines capable of drilling 165 mm diameter blast holes, up to 80 m long are used in stoping. Each stope produces up to 20 000 tonnes of ore per month and can last for three months. Heavy systematic ground support of VCR development has significantly enhanced safety for mining crews and equipment. Crater mining has not only improved productivity and enhanced safety, but also overcome various problems associated with conventional sublevel open stoping previously employed at the mine. This paper discusses the technical and operational aspects of vertical crater retreat mining at Mindola mine during the last ten years. The changes in mining method are often accompanied by changes in mine design, blast design, mining equipment, ground support and operating practices.
systems. Currently Mindola has five production faces. VCR stope layout is standard in all mining areas with stope span of 23 m and 7 m rib pillar width. The orebody is on average 12 m wide with dip ranging from 56 - 72 degrees and overturned on the south. Blasting and drawing the blasted ground on shrinkage to control hanging wall dilution and rib pillar breach creates the 23 m stope. The crown pillar is blasted into open stope after the main stope blast. The rib pillar is normally choke-blasted. Ground is drawn from the stope using TORO 301 LHD loaders. Two retreats are using conventional grizzleys, though the 3660 L/3920 L North is being converted to mechanised draw point loading. The furthest production face is on 3400 L/3660 L and is about 2.5 km from the shaft.
GEOLOGY OF MINDOLA MINE Geological structure The orebody at Mindola Shaft forms part of the complex, tightly folded synclinal structure, the Nkana Syncline. The syncline comprises east and west limbs, connected by a highly folded complex at the base. It closes towards south of SOB Shaft and opens towards the north. The syncline generally plunges towards the North-west. The orebodies, average 8 m in width where no folding occurs and dipping to the west at an average of 60 degrees. The Mindola orebody is regular with no folding averaging 12 m in width (Figure 1). Mindola Mine is separated
INTRODUCTION Mine operations at Mindola mine started in 1935. Mindola shaft is one of the four ore sources at the Nkana Division of Zambia Consolidated Copper Mines (ZCCM) Limited, located in Kitwe the hub of the Copperbelt of Zambia. The other three sources being Central, South Ore Body and North Shafts, all located on the eastern side of a complex folded synclinal structure known as the Nkana Syncline. The long-term economic viability of Mindola mine depends on the successful exploitation of reserves below 4440 feet (1330 m) level, currently the deepest mining level at Mindola shaft. A subvertical shaft has been re-deepened for future extension of mining from the current mining level to 5150 feet level (1570 m) to exploit about 27 Mt of in situ ore at 1.8 per cent copper and 0.13 per cent cobalt, making Mindola the deepest shaft on the Copper belt. An assessment of the need to replace current down dip open stoping method with an Up dip Method with backfill is constantly under investigation. To-date three mining methods have been applied to mine the orebody at Mindola Shaft, that is, conventional sublevel open stoping (SLOS), sublevel big hole open stoping (BHOS) and the current vertical crater retreat (VCR) method. At Mindola Shaft ore has, from early-90s, been mined from seven retreats on 3660 feet (1100 m) level, 3920 feet (1177 m) level, 4180 feet (1255 m) level and 4440 feet (1330 m) level using VCR mining method with conventional grizzley and mechanised draw point loading 1.
MAusIMM, Senior Lecturer, Curtin University of Technology, WA School of Mines, Locked Bag 22, Kalgoorlie WA 6433.
2.
Senior Rock Mechanics Engineer, Zambia Consolidated Copper Mines Limited, Nkana Division, POB 22000, Kitwe, Zambia.
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FIG 1 - The Mindola orebody.
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from Central Shaft by the 1.5 km wide Kitwe Barren Gap. Towards the south from the central position of the mine, the dip becomes steep and over-turned. Towards the north the orebody tends to be shallower and dips about 56 degrees. The orebodies are primarily sulphides. The dominant ore minerals are chalcopyrite, bornite and carrollite. The deposits are contained in a group of sedimentary strata in the lower Roan group of the Katanga System. The Lower Roan is made up dominantly of a footwall formation consisting of archaceous material, conglomerates, quartzites and sandstone followed by the ore formation of argillite, shales, dolomitic shale and quartzites. Above this are more dolomitic shales with the beds of sandstone and quartzite (Table 1). TABLE 1 Rock types in and around the orebody at Mindola mine Formation
Uniaxial Compressive Strength (MPa)
Hanging wall quartzite
207
Hanging wall quartzite argillite
193
Assay hanging wall
—
Hanging wall quartzite argillite
193
Porous sandstone
131
Cherty ore
172
Banded ore
105
Low shale argillite
110
Schistose ore
50
Geological Footwall
—
Footwall conglomerate
138
Footwall sandstone
193
Lower conglomerate
170
Basal sandstone
180
Hydrogeology There are four main aquifers at Mindola Mine; a footwall aquifer and three hanging wall aquifers, these are called Near Water, Far Water and Ultra Far Water. The Footwall water is related to zones of fissuring/joint and leaching, and tend to drain fairly quickly. With caving method of mining, the aquifers, especially the Near Water and Far Water Aquifers, need to be drained below a given horizon before mining can commence. The Ultra Far Water aquifer is contained in the Ultra Far Water sediments which lie some 300 m – 350 m above the orebody, and hence, has only been probed but is not routinely dewatered as it is too far in the hanging wall. Study by Golder Associates showed that as mining gets deeper and the orebody narrower, the strata above the orebody, in particular the upper Quartzite, are unlikely to breach and will maintain their characteristics as aquicludes.
VERTICAL CRATER RETREAT MINING METHOD
Main level development The orebody is accessed on four main levels (3400, 3660, 3920 and 4440 feet) through 6 m wide x 4 m high shaft cross-cuts. The 4 m wide x 3.5 m high extraction haulages are mined towards north and south parallel to the strike of the orebody. In conventional gravity method, the location of the extraction haulage is determined by the running angle of the box raises from the grizzley level. The haulage is located approximately 20 – 25 m true cover from the orebody, and attracts heavy support in form of combined grouted re-bar, meshing and tensioned lacing, plus passive (steel arch/timber set) support in the vicinity of mined out stopes. In mechanised draw point loading extraction haulages are positioned at approximately 60 m from the orebody, within a more competent Basal Sandstone formation, away from the influence of mined out stopes, with normal re-bar and cable bolt support.
VCR sublevel development In nearly all the mining retreats, the early VCR stopes utilised the pre-existing sublevel open stoping development excavations. The pre-existing sublevel development interfered with down-the-hole drilling and adjustments were made in a number of ways. The following measures were taken:
• Drill through the existing development. Problems were
Vertical crater retreat trials Following the successful drilling and blasting of the two Blast Hole Open Stopes, the first Vertical Crater Retreat (VCR) trial was conducted at Mindola Mine in the late-1988/89 (Goel, 1987). The trial was conducted on 3400/3660 feet (1036/1116 m) Level for the 5050S stope. The stope was partially developed for Blast Hole Open Stoping, and therefore the VCR layout was modified to suit the existing development. Drilling was done using the same Mission Megamatic DTH machine with 165 mm
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diameter holes. High strength slurry explosive was used for blasting the stope. Slice advance per blast was nearly 3 m. Problems like hole deviations, machine break down and misfires due to earth leakages in electric detonator leads were experienced during the blasting period, which lasted for 20 days. A total of 13 blasts were carried out. Results in terms of stope break, stope back profile fragmentation, ground conditions and dilutions were encouraging. No closures were experienced. The method proved to be ideal for the rock conditions and depth of mining at Mindola Mine. After the trial, it was felt that the new mining method could be successfully applied to mining of the Mindola orebody. Recommendations were made to introduce the method in all mining retreats at Mindola as well as at Central and South OreBody Shafts in areas of tightly folded anticlines and synclines and inter limbs of the folded structure. A project team strictly controlled all activities and selected staff was trained in the field of charging and blasting the stopes and operating the DTH machines. However, various problems have been experienced with the method over the years. These include hole closures, bumping, premature detonation, ore collapses from stope walls, ground deterioration in sublevels and exposure of men in open stopes. As a result, operating practices and mine design have had to be continuously improved. For example, the stopes are kept as full as possible to provide a maximum support to the exposed hanging wall in order to reduce hanging wall failure and/or dilution. In all the six retreats at Mindola Shaft, the stopes are 23 m stope and 7 m pillar. The orebody width varies between 10 m and 15 m. Stope heights can be up to 75 m with average stope tonnage of 70 000 tonnes. Planned extraction and recovery figures are based on historical performance of the stopes in respective retreats.
encountered in drilling through the broken/fracture ground on the floor of cross-cuts.
• Drill in between the closed spaced sublevel big hole cross-cut development (4.5 m ring burden). Problems encountered in drilling due to ground movements. A number of DTH hammers and drill rods were lost through sticking.
• A number of blast holes closed due to pillar stress. The hole spacing and the stope sizes were adjusted to accommodate the existing development.
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After passing the pre-existing SLOS development stopes, proper VCR stope development was carried out. VCR stope development comprises the top drilling, and draw point levels, and an intermediate levels where the stope drilling height exceeds 50 m, and is established midway between the top drilling and the draw point levels. Figure 2 shows the typical VCR stope profile at Mindola. These levels are accessed by stope access cross-cuts from the haulage and ramp situated in the footwall. From the main top level developments, access cross-cuts are mined at every fourth stope position to the access drive position that can be either in the footwall, or within the orebody formations, depending on ground conditions. Drilling chambers (cross-cuts) are then mined off the drive, and extended to the hanging wall/footwall. Figure 3 shows the typical top drilling level and rigging positions in the cross-cuts. Within the waste formation the development is mined 2.4 m wide by 2.4 m wide to accommodate passage of machines (Raise borer, DTH and Loader). Within the orebody the cross-cut roof is mined to 3.5 m high as drilling height. There are two standard practices for VCR layout, one where there is a single VCR drilling chamber, slyped to 8.0 m wide, and two of 2.4 x 3.5 m cross-cuts, and another where there are two
small VCR drilling chambers, each slyped to about 6.0 m wide. The two-chamber arrangement results into improved stability of excavations. Pillar drilling and blasting is done from the safe access footwall drive. Each stope is served by three 3.5 m wide x 3.5 m high draw point cross-cuts mined off the footwall loader drive. The loader drive is mined in the competent Basal Sandstone below the weak Lower Conglomerate at approximately 50 metres from the orebody. The loader cross-cuts are then extended from the orebody footwall to the assay hangingwall. Within the orebody, all the three cross-cuts are slyped to form a single holing chamber (undercut). The main considerations in optimising the layouts are:
• stability of the pillars between excavations; • cross-cuts are more stable than drives; • within the orebody, the hanging wall strata is stronger than the footwall strata; and
• blast holes should be parallel as far as possible especially near the hanging wall. Two types of drilling layouts are used at Mindola lower levels. One is where the VCR stopes comprise a single drilling lift
FIG 2 - Typical VCR stope profile.
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FIG 3 - Plan view of DTH drilling positions in the cross-cuts.
arrangement and these are confined to the 3400/3660 L, 3660/3920 L and 3920/4180 L mining blocks. The 4180 - 444 feet Level North and South retreats have an intermediate drilling level. The 3400 - 3660 L and 3660 - 3920 L use VCR with gravity loading while the rest of the retreats employ mechanised draw point loading. The decision to have an intermediate level was to reduce on distance and thus improve on drilling accuracy; to have a level at approximate mid stope height where hanging wall support can be installed; to have draw points at approximately mid-stope height to improve draw in the stope. However, the cons resulted in increase in development and stope preparation time. Delays between the preparation of the top drilling level and intermediate level in some stopes have led to stopes being pulled down too far below the intermediate level while the top is still being prepared. Excessive over mining on the intermediate level damages the hanging wall beam, and despite having cable bolts installed, results in high dilution of the ore. The undercut in the gravity loading retreats is a conventional chamber excavation, while in mechanised draw point it is a trough mined across the entire stope on draw point elevation. The false crown pillar elevation, which is about 12 m below the mined out stope, is developed conventionally using Big Hole cross-cuts. Sometimes a DTH machine drills fanned up holes from the main drilling level. On the intermediate levels drives are mined within the waste footwall formation while the three cross-cuts are mined right through the orebody to the assay hanging wall position. The central cross-cut is then slyped to full size. Because the intermediate level can be used for drawing ore, all developments are 3.5 m wide x 3.5 m high. Cleaning of the VCR sublevel development ends is done using LHD TORO 150Ds with bucket capacity of 1.6 m3 and TOROs 301 with 3.6 m3 bucket capacity. The low availability of the smaller loaders has resulted in stope development on the top levels lagging behind. The small loaders have already done more than 25 000 hours. Unlike production loaders, development loaders are not hired in terms of maintenance and are very captive. The intermediate and draw point developments are cleaned using LHD TORO 301. Currently the mine has two
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TORO 150D loaders for development. The four big loaders have done slightly over 6000 hours and their availability is good. TAMROCK personnel carry out maintenance of the loaders. In the early stages of VCR development cleaning was done using labour intensive conventional hand lashing. Later the 0.4 cubic metre bucket capacity Micro Scoop loaders were used before the current 150D and 301 LHDs.
Mine planning The mine planning department uses data supplied by geology and survey departments to develop mine plans and layouts. Generally mine planning begins with geological section whereby the geological reserves are subdivided into mining blocks which are later subdivided into stopes. Next, the main and sublevel development required to access and extract the ore are laid out on plans and sections. The calculation of stope grades and tonnages are manually done by geology. VCR development and stope layouts are drafted manually based on geological and survey data. Long-, medium- and short-term scheduling, though partially computerised by use of spreadsheets still involves considerable manual effort to generate. These tasks involve preparation of end of mine life, five-year and one-year production schedules. Stope planning involves preparation of drilling layouts, ie precise location of drilling positions and the direction, depth and size of each hole in the ring and/or fan. This task including the calculation of ring and toe burdens and amount of explosives per hole are currently manually done. Considering the size of Mindola mine with six active retreats a considerable amount of time is spent on stope design and production scheduling. In order to overcome the shortcomings of manual planning, management introduced LYNX Mine Modelling System for mine planning at Nkana in 1992 (Chanda and Kabibwa, 1996).
GEOTECHNICAL CONSIDERATIONS Rock mass data for mine design From the Rock Mass Classification Data collected from Mindola
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Lower Levels, the mining rock mass rating (MRMR) of the ore varies from 59 per cent to 79 per cent corresponding to the orebody Schistose and the Cherty ore structures respectively. Both these constitute the lower portion of the orebody. This range of results represents ‘fair’ to ‘good’ ground conditions. The Uniaxial Compressive Strength (UCS) of the orebody generally varies from 40 to 200 MPa. The weakest units, Schistose and low-grade Argillite are susceptible to failure at low stress levels. Earlier in situ stress measurements have shown that the vertical stress is generally equal to the overburden, with the horizontal stress close to 0.9 times the vertical stress. Scan line joint surveys conducted at Mindola in the Lower Levels, are analysed using DIPS joint analysis model. There are three major sets presented over the areas sampled. Though there are four joint sets, bedding forms the main plane of weakness of the rock mass. Table 2 shows the joint set data. TABLE 2 Joint set data for Mindola mine Joint plane
Dip
Azimuth
Comment
1
66
213
Bedding
2
63
48
Joint set
3
72
2
Joint set
4
60
318
Joint set
Ground behaviour In general, ground conditions do not vary dramatically throughout the lower levels. Differences in excavation performance are partially related to subtle changes in joint distributions and rock strengths, but the major influence is the stress loading conditions as generated by the mining strategy (Jeremic, 1978). Crown pillars, rib pillars and the immediate hanging wall are the main supporting structures of the stope during ore extraction. The crown pillar and the rib pillar should both be extracted in the final phase of stope blasting. The crown pillars are located below mined out stopes and are comprised of ore sections above the false crown pillar lift which lies about 18 m above the haulage elevation and is about 10 –12 m thick. As blasting of the stope progresses upward ground conditions in the crown pillar lift deteriorates due to high stress. Drilling and blasting to recover the crown pillar becomes even more difficult due to ground movement. Delayed blasting of the crown pillar induces stress in the footwall and affects the stability of the main haulage just below leading to heavy passive support in addition to grouted rebar support. In addition seismic events in form of bumping becomes prevalent when crown pillar blasting is delayed. Rib pillars between the working stope and the old stope should be recovered by drilling and blasting from the safety drives in the footwall. Rib pillars are in a high stress state and observations show that rib pillars fracture in the direction perpendicular to the orebody, ie in the direction of the major principal stress. The fracturing causes total collapse of the stope. Since the pillars are under yielding conditions, recovery is only partial or even nil, to the detriment of over-all ore recovery. Ground condition in the abutment area can create instability in mining excavations, and unsafe especially where stope preparation takes place close to the producing stope. Support in form of grouted rebars has to be installed immediately after mining, ie before substantial ground movement takes place. Golder Associates suggested that the high stress is attributed to the lag of caving behind the lower abutment. This may be due to
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the steep dip and relative narrow orebody, and that little or no load is transmitted to the footwall by the caved material for at least 300 m and probably 600 m up dip from the lower abutment. Severe stress damage to the excavation mined close to the orebody is common. Owing to its weakness, the area within the orebody bounded by the Geological Footwall (GFW) and the Base of Banded Ore (BBO) becomes problematic in VCR drilling chambers when the reinforcement is delayed. Roof caving of up to five metres high occurs, and remedial work in form of passive support becomes expensive. The influence of induced stress in the footwall by stope extraction can be observed not only in sublevel stope developments but also on the main level developments. For example, during main level development the haulage is driven at 150 m ahead of the working stope; this heading is in relatively good ground because it encounters virgin rock stress only. Experience has, however, shown that stability is only short-term because when it advances below the position of the upper level it encounters unstable ground caused by stresses induced from the progressing stope extraction. Haulage collapses due to high rock loads as a result of stress relaxation are common in the vicinity of mined out areas, especially where grout support was not installed at the development stage, in addition to passive support in form of steel arch or timber sets.
Seismicity Seismic events have been a feature in Mindola since the 1960s with the severity ranging from loud noises only to severe bumping accompanied by spitting and spalling of rock from drives. Bumping becomes prevalent in the upper part of the stope in the late-stage of the stope blast. Introduction of a series of regional support pillars as recommended by Golder Associates in the late-seventies, to reduce stress and rock burst problems did help, though high stresses still remain a problem. However, there were little benefits in as far as clean recovery of ore is concerned and in any case, damage was caused to footwall haulages. Drop in the production rate and the yield of the weak schistose in the footwall seem to have contributed to the reduction in high stress related problems. An increase in the mining rate from the current 15 000 – 20 000 tonnes to 30 000 tonnes per month per retreat in the down dip sequence will no doubt trigger high seismicity at Mindola Mine. In the proposed up-dip mining sequence, initial mining will take place in virgin areas without any stress relief. Some bumping activities are likely to begin and then die down as mining proceeds to create stress relief zones.
Stable stope spans The hanging wall argillite forms the immediate hanging wall of the stope. Two Rock mass classification systems were used to estimate stable spans; Barton’s (NGI) and Laubscher’s (1990) system. Using properties of the hangingwall argillite, the intact rock strength averages 200 MPa and Rock Quality Designation (RQD) varies between 74 –100 per cent. The adjusted rock mass strength is between 81 and 110 MPa. These data indicate good hangingwall argillite formation. For sublevel open stoping in the 1970s, stope strike spans of 9 m to 12 m by 80 m on dip were found to be stable, permitting clean extraction of stopes prior to wrecking rib and crown pillars. This gives hydraulic radius of 4 to 5.2. The current VCR stope span prior to pillar wrecking is 23 m with a 70- degree dip, length from the draw point to the top drilling level of 60 m and to the crown pillar level of 78 m, giving hydraulic radius values of 8.3 and 8.8. Nevertheless, these spans are only maintained in a stable state through use of shrinkage method. Hanging wall peeling inevitably occurs if stopes at this span are drawn empty as has been experienced at Mindola.
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Support practices in VCR excavations Top drilling level As earlier mentioned, the top drilling level of the VCR stopes is established by using one of the following three types of access; waste footwall drive, orebody drive and hanging wall orebody drive. The type of access is dependent on the prevailing ground conditions. It is desirable that the top drilling level of a stope be fully established and supported before increases in abutment stresses result from the approaching production faces. This is considered at least two to three stopes ahead of the current producing stope. The initial drive and cross-cut development is mined to the full required height and is fully supported, except where further slyping is to be conducted in order to establish VCR drilling chambers. The support utilises both full column grouted rebars/rock studs and cable anchors (6 m long x 15 mm diameter, bird caged 25 ton Freyssi Cables), to the required standard. The rebar/rock stud support is installed on a 0.75 m spacing for the first four metres of the orebody cross-cut above the GFW in order to stabilise the weak footwall area. In the slyped area, rebar/rock stud support is installed after each slyping stage. Six 5 m cable bolts are installed during or after slyping. If the initial stub is opened in the area of the Schistose ore, this is kept as small as possible and fully supported with 2.3 m anchors and 6.5 m cables before the next slyping stage is conducted. The support of the schistose area is further re-enforced with 2.4 m x 0.3m wide Tendon Straps held in place against the rock surface using 6 m x 25 Ton Freyssi Cable Bolts with face plates and lock-off units. Tendon support has greatly improved ground condition in this area during stoping. In addition to rebars/rock stud, and cable support, all drilling excavations are meshed with tensioned lacing. Four to five timber packs (of 150 mm x 0.75 m) are installed after stope drilling has been completed and before stope blast commences.
Intermediate drilling level Although the level of damage to which the excavations on the intermediate level are subjected to is lower than that of the top drilling level, cable support is still installed in the following areas:
• VCR drilling chambers: Same support as on top drilling
is installed up to the brow slot position. For four metres up to the brow, support ring spacing is reduced to 0.5 m for the next seven rings. This serves to reinforce the draw point brow. It is recommended that the draw point area be established and fully supported before extensive blasting is conducted in the adjacent stope and before the stope undercut is established. This will counteract the effects of the increasing abutment stress.
Grizzley draw point support Support for grizzley draw points consists of 1.8 m full column grouted soft cable ropes starting 0.5 m from the brow edge at 1 m ring spacing. After the three rings of the soft cable ropes, standard grizzley cross-cut support is installed using 1.8 m rebars. The drives are supported with standard support pattern of 1.8 m looped rebars.
False crown pillar level Support of the 1.8 m x 1.8 m crown pillar orebody drive and access cross-cuts consists of five full column grouted 1.8 m anchors per ring with 1 m ring spacing. In poor ground conditions the looped rebar is replaced with 1.8 m rock studs with face plates and welded/wire mesh. The support strategy includes a ‘maximum unsupported distance’ which is the maximum unsupported distance considered safe without support during the development stage. However, the maximum unsupported distance; currently 10 m in good ground is reduced when ground conditions deteriorate. The maximum unsupported distance is reduced to 5 m in poor ground conditions and down to 0 m in the case of very poor ground, ie the support is carried out on the face. In these cases a 24 hour period from the time of installation of support, must be allowed where no blasting is conducted in order to allow the grout to set.
Stope hanging wall support The dip of the orebody in the northern extremity of the mine may reach 55 degrees. Over mining of the cross-cuts (sometimes into the hanging wall) and the slabby nature of the hanging wall, gives to premature hanging wall failure and dilution into the stope. To counteract this dilution the following two support systems are used: 1.
Shrinkage mining method: The blasted ore is kept just far enough below the retreating stope back for the next blast, ie only the swell is pulled. Thus, the blasted ore is used to confine the hanging wall, preventing bulking and unravelling and thus, dilution.
2.
Cable bolting: Rings of 10 m cable bolts are installed from the various drilling levels, which will include intermediate drilling level, the top drilling level and the false crown pillar level. Where support holes are drilled from the cross-cut within the orebody, the cables are countersunk to 1 m below the Assay hanging wall. Face plates and lock-off units are fixed to the cable and tensioned on to the rock surface once the grout has set.
level with the same increase in support along the GFW contact on a 2.0 m x 2.0 m spacing, and in the Y-intersections between cross-cuts, near the lower conglomerate.
• In extremely poor ground conditions: Pre-support with cable bolts is practised, with small excavations initially mined, before slyping takes place. However, development rate is showed down.
Loader draw point level Three support standards are given for the loader draw points of the mechanised VCR stopes based on the prevailing ground conditions in the footwall area of the draw points. The good, fair and poor ground conditions are determined on the Laubscher’s (1990) Rock Mass Classification Scale. In good ground conditions, draw point areas are supported with 2.3 m full column grouted rebar only at 1.2 m bolt spacing on a 1m ring spacing up to 2 m from the brow, and a 0.5 m ring spacing in the roof up to the brow slot. In fair ground conditions brow areas are supported by 2.3 m tendon up to 2 m from the brow slot position. 2.3 m cable strands follow this on 0.5 m ring spacing. Three rings of 3 x 6m cables are installed in the roof with the first ring at 3 m from the draw point brow. The ring burden being 2 m. In poor ground condition standard pattern of 2.3 m tendon support
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The first method is the most effective, as the hanging wall cable bolting only tends to work in the area of drilling levels. The hanging wall between the top drilling and the draw level can only be supported by the correct use of shrinkage mining method. Shrinkage ground also helps confine the 7 m strike width x 12 m dip width x 70 m stope height rib pillar to enhance its stability.
VCR STOPE DRILLING AND BLASTING Production drilling Before drilling of blast holes commences, rigging positions are
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EVOLUTION OF VERTICAL CRATER RETREAT MINING
picked and marked in the roof of the drilling chamber by Survey. Large diameter long holes for the VCR portion and the rib pillar of the stope are drilled using the DTH 6200, six feet long mast Cubex machines. Mindola Shaft has four such machines for drilling in the five retreats. The machines drill blast holes of 165 mm diameter. The rotary head of the drill is electro-hydraulic, and the down-the-hole hammer is pneumatic. The mine compressed air supply is about 4 - 5 Bars at the drilling site. However, a compressor boosts this pressure to 2500 Bars. Since the top DTH drilling level is on haulage elevation, the crown pillar level is drilled with up holes of about 15 - 18 metres. The maximum drilling length for the 165 mm blast hole is kept at 52 m and that is in a single lift without an intermediate level. In the VCR stopes with intermediate level (4180 L – 4440 L) the drilling distance is 24 – 29 m. In the 3400 L - 3660 L North where orebody dips about 55 degrees and using conventional grizzley draw points, the drilling length is 60 m, but a 16 m under cut reduces the overall length to 44 m. The actual length drilled gives about 20 to 25 t per metre of ore drilled. The decrease in t/m is due to redrilling to compensate for the deviation and blocked holes. The most critical area of the stope in terms of drilling is the VCR slot drilled with parallel holes in the middle of the stope. The blast holes are drilled 3 m apart in rings spaced at 1.5 m, with the first ring at 0.5 m from the sidewall of the chamber. The holes are staggered such that the overall spacing of holes is 3 m. The zone is timed to detonate before the rest of the lift, and breaks only at right angles to the direction of the holes. The rest of the stope on the drilling level is drilled using parallel and fanned holes from the cross-cuts on either side of the VCR chamber. The twin drilling positions are spaced at 3 m along the cross-cut. The holes break both towards the undercut and the VCR slot. The false crown pillar lift can also be drilled using the combination of drifter and DTH machines as a bench and is blasted using parallel and fanned cylindrical charges after the slot is blasted. The rib pillar is drilled with a fan of holes using a DTH rig from the footwall drill stub. The pillar is drilled before or after the VCR blasting depending on the access route for the machine. Typical VCR stope drilling section is shown in Figure 4. Drilling with the DTH machines has not been without difficulties. Owing to age of the machines and inexperience of some operators and fitters, hole deviations and stuck rods, resulting into loss of hammers, are not uncommon. The two problems have been prevalent in the northern extremity of the mine where the dip of the orebody is about 50 - 55 degrees and the footwall portion of the orebody (schistose and low-grade argillites) tend to be poor. The problem was identified and drilling through these rock units is in some cases either avoided or holes are angled to drill across laminations. However, upon exposure this part of the orebody falls into the stope due to blasting concussion.
Charging and blasting High bulk strength slurry was earlier used for sublevel big hole open stoping and VCR blasting. Available now is 110 mm x 560 mm, 6.25 kg Energex cartridge slurry explosive which has a density of 1.3 g/cc with excellent water resistance. The cartridges are slit to ensure filling of the hole and avoid decoupling. The explosive power should concentrate in a length of 6 x hole diameter to form a point charge. The stope is be blasted by the cratering technique. Average advance of the stope back varies from 2.5 m to 3 m. Plugging of the holes is achieved by the use of 110 mm diameter x 450 mm length wooden split plugs. Dry sand of uniform grain size is used as stemming material. Correct plugging of the hole at the bottom and the use of bottom and top
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stemming, and positioning of point charges and timing, are essential for efficient use of explosives. 25 kg of explosives fill approximately 0.9 m length of the 165 mm diameter blast hole giving a powder factor of 0.3 kg/t of ore. Blast initiation is achieved by the use of heavy-duty U500 HD Nonel detonators in various tube lengths (21 mm, 30 mm, 45 mm, 60 mm). It is essential that the VCR zone be kept one advance ahead of the stope so that it provides a combined effect of crater and cylindrical charge in the outer holess. After each stope blast the length of blast holes and ground level in the stope are measured to determine which holes are to be blasted next and whether or not sufficient breaking face in the stope is available. A computer program is used to show the arrangement of holes at different depths of stope blast. The problems experienced in stope blasting include hole closures due to poor ground conditions in some areas, sand baking due to too much stemming and wetness, and blow outs due to insufficient stemming or placing the charge (unintentionally) high above the correct position. Coning of the bottom of the hole is common leading to difficulties in plugging the bottom. Blocked holes are either pumped out or redrilled in order to open them. Crushing of the hole due to stress has been experienced in the lower levels. The holes attain dog-ear shape up to 3 m below the floor, probably due to end effect. The holes have to be cleaned before charging. The problem of hole closure during blasting can be very serious at times leading to delayed stope blast or complete redrilling of the stope. The critical stage of the blast is that of the final over cut sill. Usually the length of the holes in the final blast known as wrecking, ranges from 8 m to 10 m. In most stopes wrecking of the sill has been successful. Few under breaks have been experienced and where this happened it was dangerous and difficult to clear, as the over cut became unsafe due to its being under cut. Hence, the charge density in the sill has to be increased at this stage as a precaution. The sill is blasted in decked charge in each hole, in order to reduce the energy, as shown in Figure 5. Each VCR blast gives an average of 2.5 m break and gives about 2 Kt of ore. In most stopes fragmentation of ore is very good giving an average of 0.5 m x 0.4 m x 0.3 m. This is shown by the performance of the loaders in the draw points. Loader performance reaches 35 to 40 tonnes per hour.
Ore handling The broken ore is removed from the stopes using the Toro 301 load-haul-dump (LHD) units. In four retreats grizzley method is applied on 3400 L/3660 L and 3660 L/3920 L retreats. The 3660 L/3920 L which has been out of production since 1997 due to an expected water inflow is being converted to mechanised draw point loading. The draw point level development consists of a loader drive, two or three draw point cross-cuts (depending on size of stope), a tip and a ventilation return airway. One LHD is allocated to each producing stope in the retreat. Production is conducted during the afternoon and night shifts. Production target per loader per hour is averaged at 40 tonnes. The low productivity for the two loaders during most of the year is due to lack of ground from stopes as a result of late delayed stope preparation and drilling, caused by low availability of development lashing loaders and DTH machines. Long tipping distances also contribute to low loader productivity. The machines operate over dumping distances of between 50 and 150 m. Availability of the hired loaders is high, the only constraint is availability of ground at times, which results into low utilisation of the machines.
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FIG 4 - Typical VCR drilling section.
The loaded ore from the stope is tipped into a 24 m long, 2.4 m diameter raise bored ore passes of which the 2.4 m box is installed at the bottom in the lower haulage. The boxes are pneumatically operated. The ore is loaded into 12 ton tray cars on a train pulled by 15 ton Jeffrey and Batman Electric Locos.
VCR STOPE PERFORMANCE EVALUATION Vertical crater retreat versus sublevel open stoping Vertical Crater Retreat method of stoping at Mindola Mine has continued for many years now. In terms of efficiencies, comparisons are made with sublevel open stoping. Table 3 shows the efficiencies in completed stopes at Mindola shaft. 4180 L – 4440 L North and South have been producing using VCR method only. The following observations can be made with regard to applicability of VCR method in varying ground conditions:
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• On 3440/3660 N the grade factor improved by 23 per cent and metal recovery by 22 per cent and dilution reduced by 29.8 per cent. On 3400/3660 S the grade factor improved by 4.0 per cent, the metal recovery by 14 per cent and dilution reduced by seven per cent.
• On 3660/3920 LN, efficiencies remained poor in VCR stopes. Three stopes experienced heavy hanging wall failure and rib pillar breach leading to high waste dilution and ore loss.
• On 3660L/3920 S the grade factor is almost the same, though metal recovery dropped and dilution reduced by one per cent.
• On 3920 L/4180 LS, the grade factor improved by 19 per cent, but the metal recovery dropped by three per cent and dilution reduced by 40.0 per cent. Pre-existing development stopes have dilution of more than 40 per cent with full VCR stopes having dilution ranging from 0 per cent to 44 per cent,
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TABLE 3 Mining efficiencies in completed stopes. Reserves Retreat
No stopes
Ore
Factors (%) % Cu
Extraction
Grade
Recovery
Dilution
3400/3660N: - Conventional
9
401042
1.65
92.9
76.9
71.5
30
- VCR
13
786884
1.46
93.8
99.8
93.7
0.2
- Conventional
9
767816
1.95
114.9
70.8
81.3
41.3
- VCR
4
213632
2.25
128.3
74.5
95.7
34.1
3400/3660S:
3660/3920S: - Conventional
10
77562
1.57
117
83.3
97.4
20.1
- VCR
16
1118344
2.06
95.6
84.1
80.3
18.9
- Conventional
9
407464
1.77
100.5
92.7
93.1
7.9
- VCR
5
322786
1.79
94
85.4
80.3
17.1
3660/3920N:
3920/4180S: - Conventional
9
534788
3.11
133.4
60.1
80.1
66.5
- VCR
16
1305319
2.14
97.3
79.2
77.1
26.3
3920/4180N: - Conventional
8
426004
2.81
95.2
75
71.4
33.4
- VCR
17
1201265
1.99
108.1
81.5
92
17.5
2
178114
2.69
133.4
61.3
81.7
63.2
4
388966
2.77
108.7
63.9
69.8
56.5
4180/4440S: - VCR 4180/4440N: - VCR
mostly due to handing wall collapse due to presence of waste hanging wall drive. No hanging wall cable bolting was done.
• On 3920 L/4180 LN, the grade factor improved by ten per cent, metal recovery by 21 per cent and dilution reduced by 16 per cent. Four VCR stopes have 0 per cent and 13 stopes have 7.6 per cent to 60 per cent dilution.
• On 4180/4440L North and South retreats, stope performance was generally poor as is evidenced by the decreased grade factors, metal recoveries and high dilution levels. This is also attributed to hanging wall collapse weakened by the presence of hanging wall drive, low rib pillar recovery due to poor ground conditions and breaching of rib pillar before crown pillar was blasted Overall, there are some improvements in the VCR retreats, (except the 3660 L/3920 L North and 4180 L/4440 L North and South retreats), compared to sublevel open stoping. It has been suggested that performance of the VCR stopes at Mindola Mine can significantly improve, especially in the 4180 L - 4440 L block, if strict adherence to standard operational practices is made. The two retreats on 4180/4440 L have been plagued by poor stope stability that has contributed to high dilution and poor recoveries. The 3920 L/4180 L North Retreat has somehow experienced stable hanging wall and stopes are pulled clean in an attempt to improve recovery, before the pillars are wrecked. Factors contributing to low recoveries and high dilution:
• The use of a hanging wall drive results in an unstable hanging wall and thus high dilution.
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• The excessive pulling down of blasted ore in the stopes during stope blasting and prior to the blasting of the crown pillar and the rib pillar, increases the risk of dilution from both the handing wall and footwall: breaching of the rib pillar on the mined out side: sloughing of the rib pillar and subsequently stope development on the unmined stope side, as evidenced during the mining of 1250 N and 1600 N on 4 180/4440 L North.
• The late stage drilling of the rib pillars, the closure of the holes, especially at the weak footwall contact affects the recovery of the rib pillar ore (-30 per cent).
• No real draw control, in terms of generating an even draw down of the ore/waste interface is practiced. It is likely that hang-ups within the draw points affect the even draw down of the ore in the stopes.
• The current hanging wall cable bolting system is not there to support the hanging wall of the whole stope to replace the confining forces of the blasted ore, but only stabilises the hanging wall in the area of the drilling levels. The hanging wall between the drilling drives is stabilised by the blasted ore alone.
Dilution The amount of expected in-stope dilution due to hanging wall caving is not determined at the stope layout preparation stage because the orebody has a regular shape and the method should make full use of shrinkage. The estimated recoveries are based
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FIG 5 - Method of VCR spherical charging.
on the performance of previous stopes in the same retreat. The source of waste dilution include hanging wall caving and rib pillar breach due to absence of shrinkage, and waste from old stope above after the crown pillar has been blasted with no draw control. It is currently difficult to quantify hanging wall waste dilution since it may not be the only source. The occurrence of hang-ups within the draw points if not quickly dealt with, will result into isolated draw and the ore/waste interface can be excessively interrupted resulting in dilution. Dilution from the hanging wall can further be reduced by increasing the skin of ore at the hanging wall from the current 1 m to 2 m and applying shrinkage draw and ensure all deviated holes are not blasted.
Stope and pillar stability Stope and pillar stability have been identified as contributing to the low efficiencies in VCR stopes at Mindola. Stope stability has been affected in some areas by the excessive stope width caused by pillar starting to fail on the mined out side of the stope. This increase in the effective stope width leads to insufficient confining load on the hanging wall beam. The use of hanging wall drive and cross-cuts (case in almost all the retreats during the initial VCR stoping) within the immediate stope hanging wall dramatically affected the stability of hanging wall beam. In some areas rib pillar stability was affected by off-line development resulting into reduced pillar width in the immediate area; continuous orebody drive (3920/4180 L south, 4180/4440 L North and South retreats) weakening the rib pillar by increasing fracturing due to extensive loading. Pillar breach has been
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common after mining the false crown pillar due to high loading in the final closure area and that the rib pillar below the crown pillar elevation was unconfined. The breach resulted in the loss of crown pillar ore as waste ingresses before the crown pillar was blasted. In shallow dipping areas observation has shown that when drawing the ore from the bottom of the stope in a non-shrinkage scenario the ellipsoid of motion intersects with the hanging wall and moves up along the hanging wall, resulting in early reporting of waste. The waste moves over the top of the ore reaching the draw points before all the ore is drawn from that stope. Hence, the practice of pulling the stope as far down as possible (taking the risk) before the crown and rib pillars are blasted should not be encouraged, as this, apart from dilution results into loss of pillar ore. In certain stopes draw point spacing is not optimal, ie points are widely spaced with no interaction between the draw zone resulting into extensive ore losses between the draw point when waste reports at these draw points. In an attempt to improve stope performance access drives have been restricted to the footwall formation and within the orebody. In areas where the access drive is mined within the orebody, it is recommended that the drive at the pillar position be mined with smaller dimensions. However, this is rarely achieved due to lashing constraints. Efforts are made to keep stopes full of blasted ground at all times, and only the swell is pulled until the crown and rib pillars are blasted. It is vital that draw point operators adhere to draw control procedures and that occurrence of hang-ups within the draw points should be quickly dealt with so that the ore-waste interface is not excessively interrupted. A program to conduct experiments using a physical draw model to study the flow of ore in stopes and the effect of draw point spacing on ore loss has been put in place for demonstrations. Draw control charts showing the amount of draw, in terms of loader buckets from each draw points are prepared by rock mechanics and issued to respective section tramming personnel. The extraction sequence of the crown pillar lift was changed and involves mining of the slot further into the stope away from the rib pillar, to retain a temporary thick rib pillar that is mined last. Rib pillars are drilled early and pre-charged before stope blasting commences. Antistatic pipes are sometimes used to protect the blast holes in the weak collar area of the schistose ore. The draw point layout that has been adopted consists of three draw points per stope. The No 1 draw point is on the edge of the Rib Pillar to facilitate recovery of the blasted Rib Pillar. The No 3 draw point is placed 2.5 m away from the stope edge to allow the ellipsoid of motion to open up along the solid pillar edge of the next stope allowing more efficient draw. This layout provides good interaction of the draw ellipsoid resulting in an even stope draw-down. The three draw-point system is derived from optimal spacing of the draw points based on the width of the ellipsoid of motion which is approximately 9 m diameter in a VCR blasted stope. Cylindrical rather than spherical charges should be used on holes drilled close and parallel to the assay hanging wall, footwall and rib pillar for stability of the walls.
OPTIONS TO IMPROVE PRODUCTIVITY AT MINDOLA MINE The current practice of open stoping at Mindola Mine (VCR) allows only up to two faces of production per level, one on the North and the other on the South. With low productivity per face of 18 000 to 20 000 t/month, it is difficult to increase the overall shaft production. Various alternatives for increasing productivity at Mindola Mine especially below 4180 feet Level have been considered:
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1.
2.
3.
Current method continues as at present with 200 m lead-lag between each retreat and 15 000 t/month production per face, a peak of six retreats giving a maximum production of 90 000 t/month. In this method development is slow, pillar breaches are common in an effort to draw clean ore, ground conditions deteriorate near the end of blasting life of the stope due to rib and crown pillars, and lead-lag requirements slow down the down dip production build-up. Mine life of 25 years with 612 kt finished copper is expected. Increase development capacity at each face and have at least two stopes producing at any one time as during the peak production periods of early-70s. This may increase production per face to 25 000 t/month giving 150 000 t/month from six retreats. This method will have all the other advantages and disadvantages given in option 1. Pillar breach will become less serious as the rib pillar is between two stopes partly full of ore. At that production rate mine life will be 19 years with 614 kt finished copper. Partial extraction with dip pillars left for secondary extraction later. Up to eight faces may become available, as the condition of lead-lag will not be applicable. Owing to independence of mining each stope, more than one stope can operate in one retreat. Eight retreats with say 15 000 t/month per stope and two stopes per retreat can bring the peak production to 8 x 15 000 x 2 = 240 000 t/month. The method can have quick production build up, the mine life will be 15 years with 515 kt of finished copper. The stope being surrounded by pillars should not suffer from serious stress problems. However, pillars will be difficult to recover, as stress and displacement resulting in low recoveries of about 50 per cent will cause some disturbance. Waste development for the secondary stope done at primary stope may deteriorate in these pillar positions.
4.
Multiple Retreats: Open more than two retreats on a level. Four retreats per level can increase the available retreats to 12 at peak yielding 2.5 mt/year of ore at peak. In this method, by keeping the stopes full with ground, stope walls will remain stable, but pillar breaches may occur in an attempt to draw clean ore. Remnant pillars will cause ground deterioration near end of the blasting life (in upper levels) of the stope.
5.
Up dip with fill: The method was proposed by the ZCCM Nkana Expansion Project as a means of increasing production to 2.6 mt of ore annually and improving the performance of the VCR method as mining depth increases to below 1380 m. An updip progression was proposed for mining of the reserves below 4440 L down to 5150 L (1570 m, approximately below surface). The 1995 ZCCM feasibility study proposed a backfill primary and secondary stope mining method advancing upward to create a closure pillar below 4440L (Pilula and Banda, 1994).
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The benefits of using VCR with backfill include:
-
reduced development requirement; reduced dilution (small stopes and shrinkage method of draw); improved metal recovery; better ground conditions in the stopes; and low operating costs as only one tracked tramming level will operate below 4440 L and that is on 5220 level; and reduced dewatering requirement due to reduced subsidence.
CONCLUSION Vertical crater mining has been reasonably successful at Mindola mine. Improvements in grade factors, recoveries and control of dilution are clearly evident compared to conventional sublevel stoping previously employed. Stope stability is the single most important factor in the use VCR mining at Mindola. Poor stope stability is responsible for low recoveries and high dilution in VCR stopes. Options to increase productivity include opening up more than two faces per level, updip mining with fill and partial pillar extraction. To implement any of these options will require greater geotechnical in put. As mining operations at Mindola get deeper more adverse ground conditions due to high induced stresses will be encountered. With falling ore grades and low copper prices, efforts are being made to become more efficient in the mining of the low-grade massive orebody at Mindola. Hence, due care should be exercised in mine design, planning and extraction of the ore in order to enhance safety, maximise metal recovery and reduce dilution.
REFERENCES Chanda, E K C and Kabibwa, M S, 1996. The development of computer aided underground mine planning at Nkana Division of ZCCM Limited, in Proceedings of the 26th International Symposium on Application of Computers and Operations Research in the Mineral Industry, 1997, 16-20 September, Pennstate Univ, (Ed: R V Ramani), pp 89-196 (Society of Mining, Metallurgy and Exploration: Littleton, Colorado). Goel, S C, 1991. Geotechnical aspects of vertical crater mining method in a deep mine, African Mining 1991, 10-12 June, Harare pp45-53 (Institution of Mining and Metallurgy: London), . Goel, S C and Pilula, E M, 1992. The use of Vertical Crater Retreat mining at Mindola Mine, Massmin 1992, pp 179-185 (South African Institute of Mining and Metallurgy: Johannesburg). Jeremic, M L, 1978. Stress Mechanism at the Mindola Mine, Zambia, Africa, Canadian Instit Min Metall Bulletin, 7(799):77-83. Laubscher, D H, 1990. A geomechanics classification system for the rating of the rock mass in mine design, Journal of the South African Inst Min Metall, 80(10):257-273. Pilula, E M and Banda, J Z, 1994. The development of back fill mining methods at Nkana, XVth CMMI Congress, Vol 1 (Mining), pp177-188, (South African Institute of Mining and Metallurgy). ZCCM Limited, 1995. Mindola Mining at Depth Class IV Feasibility Study, 174p. ZCCM Nkana Division Rock Mechanics Department Internal Reports.
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An Experimental Study on Large-Diameter Longhole Mining of High Stope at Anqing Copper Mine Wang Renfa1, Xia Qian1 and Jiang Zhiming2 ABSTRACT All high stopes and pillars in Anqing Copper Mine were extracted by use of large-diameter longhole mining method. The height of stopes was 120 m. The forced mining was the leading concept that the safe and efficient operation was carried out in this mine, while the mining production was faced with the major technical problems on the stability of both primary ore pillars produced in the first-step stoping of stopes and the cemented fillbody on two sides of stope in the second-step stopping of pillars. Starting with mining technology, the authors emphatically describe the test results obtained in Anqing Copper Mine and make an approach to the deployment prospects and development trend of large-diameter longhole mining technology in China underground metal mines.
INTRODUCTION The large-diameter longhole mining method has advantages of high mining efficiency, low operating cost and safe stoping, etc which generally is considered as an advanced mining technique in the world so far. The study on this mining method in China began in 1977 and the pillar stoping test of VCR stope was the first success at Fankou Lead-Zinc Mine in 1984. Large workings, high stopes and open casting-like mass mining are main trends of underground mining technology development nowadays. In the past more than ten years, with the aim of forced mining, the deep scientific research on new technology/techniques/equipment for large-diameter longhole mining was made and its great technical breakthrough and progress were achieved in Anqing Copper Mine. This mining method not only filled a blank in the field of over 100 m high-stope mining in China but also made the technology for underground forced mining in China reach the international advanced level.
BRIEF DESCRIPTION OF MINING METHOD The 1# main orebody in Anqing Copper Mine was large and thick, accounting for 85 per cent of total ore reserve of the mine. The average thickness of the orebody varied from 40 m to 50 m and the dip was > 70o. The orebody and surrounding rock were fairly stable. The large-diameter longhole mining of high stope was used for the whole orebody. The stopes perpendicular to the orebody strike were arranged, and stopes and pillars were mined in two steps. The cemented filling with classified tailings was used for emptied stopes, while pillars were extracted and filled with classified tailings. The length of stopes was equal to the thickness of orebody (40 - 50 m on average), the width was 15 m, and the height of stope was 120 m. Drilling was performed in two sections, and continuous stoping was conducted in two high stopes. Simba-261 high pneumatic down-the-hole drills were used to drill Φ165 mm downward longholes. Common emulsion explosive was used for blasting. Ore was removed by ST-5C and ST-6CN LHDs. Remnant ore was recovered by remote-controlled LHDs, and the mined-out area was filled up continuously. The schematic diagram of pillar-stoping room structure is shown in Figure 1.
1.
Anqing Copper Mine, Anqing, Anhui 246131, PRC.
2.
Changsha Institute of Mining Research, Changsha, Hunan 410012, PRC.
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STUDY ON LARGE-DIAMETER LONGHOLE FORCED-MINING OF HIGH STOPE The ore quantity in high stope at Anqing Copper Mine was 0.25 0.40 Mt generally and the operating period of stope, including stoping, ore removal and filling, may be up to 14 months. The mined-out area generally was 60 000 - 100 000 m3. The single-side exposure of stope or fillbody was 5000 - 6000 m2. Highly concentrated mining operation required the mine to perform the forced mining, while maintaining the stability of stope or fillbody was a prerequisite for the forced mining. Based on the reference to successful experiences in large-diameter longhole mining , trackless ore-removal system and high-concentration tailing filling used in the developed countries, Anqing Copper Mine emphatically made a research on high stope and pillar forced-mining technology, high-concentration (cemented) tailings filling technology, stability and failure mechanism of high stope or fillbody, seismic effect of stope blasting and strata behaviour of stope, then put forward a large-diameter longhole forced-mining of high stope, which has advantages of high efficiency, low-cost and safe operation. This new technology has been widely used in the mining.
Technology of large-diameter longhole forced-mining of high stope The forced stoping, ore-removal and filling are the basic principle making the large-diameter longhole forced-mining of high stope come true. The essential purposes of forced-mining is to intensify and simplify mining process so as to realise highly concentrated mining operation. The forced ore-removal and filling themselves are a basic way to maintaining the stability of stope or fillbody, while the forced mining process, due to occurrence of larger blasting scale and strong seismic effect, has unfavourable effects on the stability of stope or fillbody. However, shorter mining period of stope is favourable to the stability of stope or fillbody. Apparently, the forced mining is the core of forced stoping process so that the mine has to search for a proper forced-mining technology to handle the contradiction between the forced mining and the stability of stope or fillbody.
Ore caving with combined column-spherical cartridges Conventional VCR has advantages of homogeneous fragmentation and smaller failure effect of blasting on the medium, etc but it obviously doesn’t meet the requirements of highly forced mining in modern mines due to expensive high-explosive, complicated detonating system, less ore caving each blasting and more auxilary blasting operation. Similarly, the side caving with the full height of stope has advantages of simple charging and lower direct blasting cost, but fails to be widely used for mines due to difficult adjusting of blasting scale, poor blasting results and obvious failure effect of blasting on the medium. Based on the deep research into blasting mechanism of spherical and column cartridges, the concept of bench mining for open pit mines was introduced to Anqing Copper Mine and a successful experimental study was made on a new technology of small-blocking cut and overhand blocking side- caving of VCR with combined spherical-column cartridges.
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FIG 1 - Schematic diagram of pillar-stope room structure.
Small blocking cut-blasting of VCR Nine to 11 blast holes were arranged generally in the area of small blocking cut of VCR. How to prevent blast holes from plugging and jetting, etc after blasting was a key to the successful use of small blocking cut of VCR, because of less blast holes, smaller section and greater constringency of blasting. The successful experiences in the small cut-blasting at Anqing Copper Mine are as follows:
• use common emulsion explosive for blasting and properly decrease a depth of cartridge centre-placement in blast hole;
• fill blast holes with the combination of sand and water-seal; • determine reasonable millisecond interval of blasting between holes; and
• accurately measure depth of blast holes after blasting. Overhand blocking side-caving The overhand side-caving took the cut as a free face. The blockings of overhand side-caving were staggeringly arranged from opposite direction at different elevations. Four to six rows of blast holes were blasted per blasting at every blocking. The height of stepped stope was 10 - 15 m and the stepped pillar-stoping room was 6 - 8 m high. Ore once blasted was 10 000 - 25 000 t generally. The successful experiences in overhand blocking side-caving are as follows:
• practise the sidehole controlled-blasting; • widely use air-decked charging structure; • staggeringly arrange adjacent side-caving blockings;
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• make use of V oblique detonation sequence; • use common emulsion explosive; and • adjust blasting scale and the maximal charge of single shooting blast hole properly.
Seismic wave propagation pattern of stope blasting In order to study the failure effects of blasting with combined spherical-column cartridges on the primary ore pillar or cemented fillbody on both sides of stope, the site observation and analysis of seismic wave propagation pattern of blasting in the cut and the overhand side caving of VCR were made extensively. The results showed that the initial peak particle vibration velocity of seismic wave was maximal but attenuated rapidly when the small blocking was cut by blasting, but the initial peak particle vibration velocity of seismic wave of blasting in the overhand blocking side-caving was minimal (about 21 per cent of the former) and its attenuation was slowest. This showed that the maximal charge of single shooting blast hole in the blasting of overhand side caving was more than that small blocking cut of VCR, but its vibration velocity and damage to adjacent rock were smaller. It can be seen that the blasting technological process of stoping can be simplified and intensified, by use of the small blocking cut and overhand side caving of VCR. This mining method not only can meet the requirement of large-scale forced-mining but also greatly reduce the damage caused by the large-scale blasting of stope. The method of ore caving with combined spherical-column cartridges has been widely used for the stoping of rooms and pillars in high stope at Anqing Copper Mine.
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Heterogeneous interface controlled-blasting of pillars in high stope The leading concept of heterogeneous interface controlled-blasting was to ensure effective ore/rock fragmenting within the side-row controlled-blasting under the prerequisite to controlling the damage caused by the blasting of side-row blast holes to adjacent cemented fillbody. The first choice of main technical methods was to reduce the charge of blast hole, but the configuration of charging was a key to the adjustment of explosive energy distribution and the controll of blasting effect. Owing to being the cemented tailings fillbody on both sides of pillar-stoping room, the boundary conditions were very complicated. Viewed from mechanics, the interface between ore/rock and cemented fillbody had a certain cohesive force, but its cohesion strength was normally lower than the tensile strength of cemented fillbody and the ratio of strength decrease was 0.33 0.67. The structure of interface transition zone proposed by Maso can prove similarly that the transition zone of interface is a weak part of ore/rock – cemented fillbody structure. As viewed from the results of observation on the seismic effect of blasting, the seismic wave in the fillbody, compared with in ore/rock, was slower in propagation velocity, lower in peak particle vibrating frequency and shorter in duration. It comes to the conclusion that the fillbody, relatively, was oft which was of function of absorbing the seismic wave of blasting and was favourable to resisting the blasting vibration. It can be known from the results of dynamic load test for the specimen of cemented fillbody that the stress-strain curve became steep and the initial linear section extended, the yield limit correspondingly increased and a ardening phenomenon of cemented filling material occurred with increase of loading rate; especially, the response of the low-strength fillbody to this phenomenon was more remarkable. Therefore, the fillbody with a low cement – tailings ratio can bear stronger blasting vibration also. Based on the analysis above, the transition zone of interface between ore/rock and cemented fillbody was a weak part of ore/rock – cemented fillbody structure. The relatively ‘soft’ fillbody was characterised by absorbing the seismic wave of blasting and producing the instantaneous hardening. This indicates that the damage of stope blasting to fillbody was reduced due to the existence of interface, while the failure of fillbody was mainly the shear failure caused by the extrusion of rock and ore. Therefore, the multi-deck and small air-space charging with small spherical cartridges was used for the controlled blasting of heterogeneous interface to avoid the failure caused by excess deck-charge. Meanwhile, this charging method can make explosion energy redistribute, reduce the effect of explosion stress wave, increase the expanding effect of explosion gas and keep the failure caused by blasting in the range provided by the blasting design. The parameters of side-hole charging and the distance from sideholes to the interface of fillbody were a key to the design of heterogeneous interface controlled-blasting. Small blasting tests were made by means of the method of quadratic factorial test of two factors (X1 – distance from side-row holes to interface of fillbody; X2 – air-space length of sid-row holes) so as to make a comprehensive evaluation of the extent of fillbody failure, the fissuring frequency, the conformability of fillbody interface, the percentage of oversize fragments and the blasting cost, etc. Based on the results of test, the optimal solutions of X1 and X2 were obtained; according to Livingston’s blasting crater theory, it was derived that the distance from side holes to fillbody interface was 1.51 m and the optimal air-space length was 0.52 - 0.54 m when Φ165 blast holes were used for the pillar stoping. In consideration of the drilling deviation and the unconformability of fillbody interface, etc the distance from side holes to fillbody was recommended to be 1.5 - 1.7 m; the air-space length was taken to be 0.6 m when the deck charge in side-row holes was 10 kg.
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A full-scale test showed good results. When observation on the mined-out area was made after all ore in stope were removed, it was found that the wall surface of fillbody on both sides of stope were conformable and several blasted side-holes inclining into the fillbody left the half-side hole mark. At the middle and later stage of ore removal in bulk, the tailings mingling percentage was rated, by means of size grading, in the ore-removal entrance with more apparent tailings. The measured tailings mingling percentage was 2.42 per cent. If the collapse of local protruding part of fillbody after blasting was deducted, the average thickness of fillbody collapse was only 0.09 - 0.10 m. The successful continuous stoping of pillars in high stope has strongly proved that the selected parameters of side-holes blasting were reasonable so as to obtain the expected results that overbreak of pillar-stoping room was controlled and the seismic effect of blasting was weakened; meanwhile, these parameters provided a reliable technical provision for the safe stoping of pillars in high stope.
Application of large-diameter longhole forced-mining of high stope Since 1991, the scientific research and the mining operation have been performed simultaneously in Anqing Copper Mine, and the large-diameter longhole forced-mining of high stope has been deployed in the whole mine. By the end of October 1998, 20 high stopes (including six pillars) were safely stoped. The total ore mined was 4.68 Mt. Anqing Copper Mine obtained remarkable economic and social benefits. Based on statistical data, overall capacity of stope in Anqing Copper Mine was basically 900 - 1000 tpd; efficiency of mining was 118.42 t/shift; per-capita productivity was 1509.6 tpa. Compared with average data taken from China similar mines, the capacity of stope in Anqing Copper Mine increased by two - three times and the efficiency of mining increased by one - two times. The experiences in successful application of large-diameter longhole forced-mining of high stope in Anqing Copper Mine can be summed up as follows: The large-scaled blasting of stope can be brought into operation by use of the new technology of ore caving with combined spherical-column cartridges, which can make the blasting of stope meet the requirement of forced mining under the prerequisite to maintaining the stability of stopes or fillbody.
• The controlled blasting of side-row holes was beneficial to reducing the damages caused by the blasting of stope to primary ore pillars on both sides of the stope or fillbody.
• Strictly controll the deviation of blast holes to reduce the overbreak and overgettings of stopes, especially rooms.
• To practise the blasting with retaining partial ore was helpful to improving the stability of stopes or fillbody.
• After blasting, the ore removal and filling of stope should be intensified to shorten exposure time of mined-out area as far as possible.
• The stoping of rooms and pillars should be performed according to the designed mining sequence.
• The monitoring and controll of ground pressure should be strengthened. According to the results of analysing the monitored data, adjust the mining sequence in time and select reasonable support pattern of stope to ensure the mine a safe operating condition.
Technology of cemented filling with high-concentration tailings Filling system The filling system in Anqing Copper Mine was a vertical-bins
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feeding – cementing of tailings — two-phase flow transpoting system which mainly consisted of five subsystems of tailings discharging, dry sand feeding, cement feeding, fill-mortar agitating and air/water supply, which made up three sets of separate tailings mortar preparing and transporting systems. The fill mortar prepared at the surface agitating station was discharged down to Filling Level -280 m via Φ125 mm filling drillhole, then flowed by gravity via the main pipeline (Φ110 mm seamless steel pipe) into the mined-out area of stope filling. The depth of drill hole was 334.5 m , the main pipeline was 1188 m long. The actual maximum ratio of transporting length to height was 4.55. The concentration of fill mortar was 73 per cent - 74 per cent. The capacity of single filling system was 70 - 90 m3/h. The capacity of two filling systems was 1200 - 1500 m3/d (by dry volume). The filling system in Anqing Copper Mine is shown in Figure 2.
• The dewatering system of corrugated pipes wrapped with 26-mesh nylon filter cloth covered with 90-mesh nylon filter cloth was used for stope dewatering. The dewatering capacity of single pipe was over 5 t/h. It can be said that the difficult problem on the filling and dewatering of high mined-out area was solved successfully.
• According to the load-bearing capacity required by fillbody and the results of filling materials test, the ingredient ratio of filling materials in layer was designed for the mined-out area at various elevation. Therefore, the filling cost was reduced greatly under the prerequisite to ensuring the quality of filling.
• Three-way drainage valve was installed between main filling pipeline and stope filling pipe to ensure that the filling water and flushing water run away from stope, and reduce the segregation of fill mortar.
Stope filling technology Stope filling technology is one of key links making the filling with high-concentration tailings come true. Advanced and reliable stope filling technology will ensure the continuity of filling with high-concentration tailings and the quality of stope filling effectively. The design of stope filling technology included: selecting the discharging point of fill mortar, designing the bulkhead, studying the method of stope filling and dewatering, and designing the ingredient ratio of filling materials. The following conclusions were drawn from the study above:
• The discharging points of fill mortar should be selected according to the shapes of stope space, the conditions of passage way and the hanging position of dewatering pipe, etc. Dewatering pipes, in principle, can not be placed on the same side of discharging point.
• Wooden props and reinforced wire-mesh were used as a flexible bulkhead which was suitable for this filling operation.
Application of high-concentration tailings filling system In Anqing Copper Mine, about 1.1 x 106 m3 filling was completed by use of high-concentration tailings filling technology by the end of October 1998. During the pillars in high stope were mined, the stability of tailings-cemented fillbody on both sides of pillars was so good as to ensure the safe stoping of pillars in high stope. Practice proved that this filling system had advantages of high automatisation, reliable operation, high capacity, less variation in transported fill-mortar concentration, homogeneous mixing of filling materials, little segregation of fill-mortar and good effect of stope dewatering, and played an important role in the improvement of automatisation and productivity in Anqing Copper Mine. Therefore, it can be said that this filling system is of a widespread use prospect for mines.
Research on stability of rockmass and fillbody The large-diameter longhole mining of high stope being 120 m high and 40 - 50 m long was used in Anqing Copper Mine.
FIG 2 - Filling system in Anqing copper mine.
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AN EXPERIMENTAL STUDY ON LARGE-DIAMETER LONGHOLE MINING OF HIGH STOPE
Therefore, makes the primary ore pillars or cemented fillbody objectively produce large exposed sides (up to 5000 - 6000 m2), no matter when the room was stoped in the first step mining or the pillar was stoped in the second step mining. The formation of mined-out space would make the stress in the primary ore pillars and the surfaces of hangingwall and footwall release. In this way, the intrinsic state of three-dimensional load balance suffered a damage. However, the stress can reach a new balance state due to displacement, deformation and failure of rock mass. The good or poor state of stope stress was directly related to the sequence and height of stoping and the conditions of mined-out space. Based on the early achievements in the investigation and statistics of engineering geology, the measurement of intact rock stress field and the determination of rock mass/filling material mechanics parameters, etc seven constitutive 3D models of orebody (rock mass) – fillbody failure were established according to the sequence and height of stoping possibly used in the process of stope mining and the conditions of mined-out space; and the analogic simulation and calculation of stress, displacement, yield and failure existing in surrounding rock, primary ore pillars and cemented fillbody in the process of large-diameter longhole mining and filling of high stope in Anqing Copper Mine were made by use of the modified version of finite element program. In order to further research into the changing law and characteristics of stress and strain in orebody (rockmass) illbody, two plane models and one analogic material stereo-model were established respectively in geometric analogic ratio as 1:100 to make the analogic simulation tests. Based on the results of analysing the tests above, the following conclusions can be drawn:
increase can be up to seven per cent. As for 120 m high fillbody, the extent of damage would be more serious. Therefore, the relative technical measures should be taken to avoid large area and long-time exposure of fillbody as far as possible.
Main technical and economic indexes Total capacity of stope
900 - 1000 tpd
Number of simultaneous ore-removing
3
Efficiency of driving
0.57 m/shift
Efficiency of large blast hole drilling
30 - 33 m/shift-drill
Average efficiency of bulk ore-removal
1440 tpd
Maximal efficiency of ore-removal
2480 tpd
Percentage of ore dilution
13.13 per cent
Percentage of ore loss
13.04 per cent
Boulder yield
Yield Stress
Stope
Shear stress
3.
Viscosity (gradient)
Hopper
Yield stress Control Level
400 200
600
0
800
Pressure 2000 kPa
Shear rate FIG 7 - Shear stress - shear rate relation for Bingham fluids.
No Flow Vertical Head = Yield Stress
Future borehole design Stope
FIG 5 - Flow process within Cannington pipeline.
Borehole wear was assessed using a pair of callipers and a camera. The calliper results showed a thin 5 - 10 mm paste coating had built up on the wall, thus protecting it and preventing wear. The images sent back from the camera showed that a smoother bottom had developed on the pipe. There was no evidence of any impact point. Visual and thickness tests on underground pipe showed that there has been little wear of this pipe. This is due to the laminar flow behaviour of the paste, where flow velocity at the walls of the pipe is minimal and the fill primarily flows as a plug of material adjacent to the annulus of low velocity material at the wall.
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Primary filling horizons have recently been extended down a further 200 m as a result of the installation of a fully cased piping system (see Figure 4). Instead of having a single reticulation system on the 325 metre level there will be horizontal reticulation on all levels at 50 m vertical intervals to 520 metre level. Planned mining in areas to the south of and below the current mining blocks require such expansion to the underground system. One major limitation of the current system is in the case of a blockage. The method currently employed for a blocked pipe is to break the pipe every 20 m which is time consuming and limits the chances of clearing the blockage. The new system to 520 level incorporates specially designed valves at every 50 m level. The pipeline will be emptied from the surface down if a blockage occurs, resulting in a minimum static pressure at each valve. Some of the features of the valve are:
• standard piece designed to withstand design pressure of 11.2 MPa;
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• minimum offset from pipe to reduce ‘dead’ zone; • sealing face material resistant to adhering of paste; and • manual opening achieved from a remote location using an
• Dynamic stability - typically 1 to 2 m of failure occurs on exposed walls from blasting.
• Ore recovery - success in exposing the fill to the widths and
heights planned is not always achieved, resulting in lower recovery of ore. This is due to a combination of correct standoff distance (from hole to fill), drill hole accuracy, drill hole performance and ground conditions.
Enerpac cylinder.
Stope void availability One of the on-going operational challenges faced at Cannington is to provide a consistent volume of void available to fill, such that average and peak demands for filling are similar. Fluctuations in demand (‘boom or bust’) can be attributed to two factors:
• Stopes of varying size. Ideally all stopes would be the same
size, however many factors contribute to the size of each stope; the two dominant factors being orebody geometry and ground stability.
• Timely turnaround of stopes. Delays in providing void
available for filling can occur as a result of excessive stope tail production times. Such delays can reduce the short-term demand for fill which must be compensated for in the future with an increase in demand.
Fluctuations in demand can cause delays in the mining sequence and therefore limit production rates. This will occur if the instantaneous peak demand exceeds the rate at which fill can be supplied.
Fill failure caused by dynamic loading has had a relatively minor effect on feed grade to-date and has depended on the success to which this dilution has been separated from the ore stream. Aside from the cost of moving additional material, from time to time significant problems have occurred in the materials handling system. The sticky nature of the material has caused problems in transfer chutes, conveyors and surge bins.
Stability modelling Stability of the fill during exposure is assessed using a modified three-dimensional version of Terzaghi’s model for arching in soils (Winch, 1999), as described in the following equation. The model uses strength test data results and determines a cement content required to stabilise the fill for a given exposure geometry and extraction and filling sequence. An example is presented in Figure 9. σ y T = σ y0
IN SITU PASTE FILL PERFORMANCE
W1W2ρg − ( c + Kσ y 0 tan φ ) 2 R(W1 + W2 ) + K tan φ
−2 R(W1 + W2 ) K tan φ( y − y0 ); y ≥ y 0 1 − e W W 1 2
Fill exposure performance To-date 12 fill masses have been exposed, with exposure widths up to 20 m and heights of 50 m. Their performance can be summarised as:
• Static stability - one significant failure of a fill mass has
occurred (up to 5 m depth), with evidence suggesting that the cause was due to undercutting by the blasthole rings (see Figure 8).
where y
= depth in fill (m)
σyT
= vertical stress at depth y in fill (kPa)
ρ
= fill density (t/m3)
g
= gravitational constant = (m/s2)
c
= material cohesion (kPa)
φ
= material friction angle (degrees)
W1,W2
= stope widths (m)
K
= lateral pressure coefficient (ratio of horizontal to vertical stress)
R
= ratio of active wall area (for arching) to total wall area
y0
= depth of fill at which arching starts to develop as fill cures (m)
σ y0
= vertical stress at depth y0 (kPa)
This model predicts the static stability of fill only and does not consider dynamic stability. Little is understood about the dynamic stability of mine backfill, yet it is seen as the primary cause of fill dilution at Cannington. It is planned to undertake research into this topic in the coming year to improve the performance of the paste fill. Fill failure caused by dynamic loading has had a relatively minor effect on feed grade to-date and has depended on the success to which this dilution has been separated from the ore stream. Aside from the cost of moving additional material, from time to time significant problems have occurred in the materials handling system. The sticky nature of the material has caused problems in transfer chutes, conveyors and surge bins. FIG 8 - Example of static instability of a paste fill mass.
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FUTURE WORK
160
There is still much to learn about the paste fill system at Cannington. The following areas for future work have been identified:
Height of Exposed Fill Mass(m)
140 120
Quality control and future strength testing
100 Unstable
An in-house strength testing facility has been developed to meet two objectives:
80
• provide regular quality control assessment of the fill strength;
60
and
Stable
40
• assess strengths of alternative fills and binders.
20
Currently, alternative binders such as slag and gypsum are being tested.
0 28
30
32
34
36
38
40
42
44
Width of Exposed Fill Mass (m)
FIG 9 - Example of using the fill stability model.
Stability modelling Stability of the fill during exposure is assessed using a modified three-dimensional version of Terzaghi’s model for arching in soils. The model uses strength test data results and determines cement content required to stabilise the exposure geometry and history analysed. An example is presented in Figure 9. This model predicts the static stability of fill only and does not consider dynamic stability. Little is understood about the dynamic stability of mine backfill, yet it is seen as the primary cause of fill dilution at Cannington. It is planned to undertake research into this topic in the coming year to improve the performance of the paste fill.
PASTE-ROCK AS MINE BACKFILL The mine currently produces over 300 000 tonnes of waste rock per annum and this is handled separately from the ore stream where possible. The licence for Cannington to operate includes the elimination of surface waste rock stockpiles at mine closure. To achieve this it is necessary to both remove the existing surface stockpiles (mostly to underground) and ensure that waste material mined remains underground. Extraction strategies and orebody geometry limit the volume of stoping void that can be filled with waste rock alone. Thus it is necessary to develop a filling strategy which allows for paste and rock to be discharged into stope voids simultaneously. Given the need for stable fill exposures, the combined material must demonstrate adequate strength. This can only be achieved consistently throughout each fill mass by ensuring that the paste and rock mix well in the stope. To achieve this, several design criteria and operational procedures have been developed, namely:
• paste and rock are to be discharged into the void via a single fill pass accessed above the stope;
• no rock is to be tipped without paste; • a maximum ratio of rock to paste will be stipulated at any given time during filling, based on stope geometry (to ensure no excessive poor quality rock paste is exposed during adjacent stope extraction); and
• fill passes will preferably be located centrally within the stope and vertical.
Some of these requirements are probably conservative and will remain until experience in the performance of the paste-rock has been gained.
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Mining through fill Future stope extraction will require mining through fill for access. Therefore a safe and efficient method of mining through the fill (both paste and paste-rock) must be developed. Fibrecrete and mesh reinforced shotcrete will be considered as the primary ground support methods. A trial is anticipated in the next six months.
Paste-rock quality To-date there is no evidence to demonstrate the success of mixing paste fill and waste rock in the stope to produce a fill of adequate strength for the duty required. It is planned to mine through a paste-rock fill mass in the next six months to assess the quality of the material.
Dynamic stability The major cause of fill dilution in the future is anticipated to be due to dynamic loading. It is therefore important that an understanding of the dynamic behaviour of the fill is developed in order to predict damage and engineer the system to minimise the effects of blasting. It is planned to install vibration monitoring instruments into a paste fill mass such that measurements can be taken during multiple adjacent blasts both within a given stope and across a sequence of adjacent stope extraction events.
Time dependency No data is available on the long-term (+90 day) strength of the paste fill. Given that historically some mine backfills (primarily consisting of high sulphide tailings) have demonstrated a reduction in strength in the long-term, it is important that the long-term strength of Cannington paste fill is assessed.
CONCLUSIONS Although its structure remains intact from commissioning, modifications have been made to the paste fill system over the past three years. These have been undertaken in response to a number of technical and operational issues in order to provide a functional and reliable system that delivers fill to the stopes at the rate and quality required to sustain mine production. It is important that the system can be used with confidence by the mining operation, as paste fill is a critical component of the mining cycle. The current and future challenges are to build on the solid foundation and improve system efficiencies whilst maintaining required end use engineering specifications.
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ACKNOWLEDGEMENTS
REFERENCES
The authors wish to thank BHP Cannington for their permission to publish this paper. The authors also wish to thank Professor David Boger and Sam Clayton of the University of Melbourne for their contribution to the research into paste fill at Cannington.
Skeeles, B E J, 1998. Design of paste backfill plant and distribution system for the Cannington project, in Proceedings MINEFILL ‘98, pp 59-63 (The Australasian Institute of Mining and Metallurgy: Melbourne). Winch, C M, 1999. Geotechnical characteristics and stability of paste backfill at BHP Cannington Mine, B Eng Thesis, James Cook University, Australia (unpublished).
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Engineering the New Olympic Dam Backfill System G Baldwin1 and A G Grice2 ABSTRACT WMC Resources Pty Ltd – Copper Uranium Division operates the large underground Olympic Dam Mine at Roxby Downs, South Australia. These operations mine and process the ore to produce refined copper, uranium oxide, gold and silver. In mid-1996, the company committed to the Olympic Dam Expansion Project (OEP) to expand production from 3.0 million tonnes per annum of ore (85 000 tonnes per annum of copper metal) to over 9.0 million tonnes per annum of ore (over 200 000 tonnes per annum of copper metal). Along with the increase in the ore production capacity of the underground mine and the surface ore handling and processing facilities, expansion of the backfill system was required. The target backfill placement rate was 2.0 million cubic metres of backfill per annum which was to be predominantly delivered as cemented aggregate fill (CAF). This paper discusses the engineering activities undertaken to assess and change the method of designing the required strength of fill exposures. In addition, the operational changes that were introduced to achieve the required fill quantities and the mix designs to meet the specified strengths, including the use of alternate binders, are discussed. This work resulted in a large increase in the fill placement rate and a reduction in the overall unit cost of fill.
INTRODUCTION WMC Resources Pty Ltd – Copper Uranium Division operates the large underground Olympic Dam Mine at Roxby Downs, South Australia. These operations mine and process the ore to produce refined copper, uranium oxide, gold and silver. In mid-1996, the company committed to the Olympic Dam Expansion Project (OEP) to expand production from 3.0 million tonnes per annum of ore (85 000 tonnes per annum of copper metal) to over 9.0 million tonnes per annum of ore (over 200 000 tonnes per annum of copper metal). Along with the increase in the ore production capacity of the underground mine and the surface ore handling and processing facilities, expansion of the backfill system was required. The target backfill placement rate was 2.0 million cubic metres of backfill per annum which was to be predominantly delivered as cemented aggregate fill (CAF). Historical practice at Olympic Dam had been placing a high strength CAF for around a decade. Fill was prepared in a surface mixing plant using various combinations of dewatered tailings, dune sand, crushed mine waste rock, cement, flyash and water. The fill was batched in standard concrete mixing bowls (6 m3 nominal capacity) and transported by open semi-tipper trucks to boreholes drilled vertically to each high-rise open stope. The hole depths varied between 350 and 500 metres depending on the stope crown location. Surface cylinder testing showed typical strengths in excess of 5 MPa and coring through fill in situ showed values over 12 MPa. The upgraded CAF Mixing Plant as part of OEP was based on continuous CAF mixing in a pugmill that was process controlled with delivery using the existing semi-trailer tipper trucks. Olympic Dam Operations undertook a further critical review of the backfill process in early-1998 and a project manager was appointed. The opportunity to review the strength requirements and mix designs for cemented fill was realised after literature 1.
Manager – Engineering and Site Services, WMC – Olympic Dam Operations, PO Box 150, Roxby Downs SA 5725. E-mail: [email protected]
2.
MAusIMM, Principal Mining Engineer, Australian Mining Consultants, Level 19, 114 William Street, Melbourne Vic 3000. E-mail: [email protected]
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searches and a visit to Mount Isa Mines in June 1998. The basis of the revised strategy was to ‘design the fill mass’ for the geometry of each stope and to develop a delivery process utilising the continuously mixed and process controlled CAF mixing plant that satisfied the specified fill mass design, in particular the specified strength for cemented fill. The review of the CAF strength design and mixing processes was expected to yield the most significant cost-savings due to the high cost of cement and flyash in CAF. Preliminary investigations of past practices and literature (internally and externally) identified the more cost-effective strategy of using zoned strengths of CAF in each stope. In general, stronger CAF would be required at the base of the stope and progressively weaker CAF placed in zones in the upper sections of the stope. The mine engaged Australian Mining Consultants (AMC) to conduct a technical review and develop a ‘fill mass design process’ with particular emphasis on CAF strength design. A ‘design study project’ was initiated to investigate the zoned strength requirements for the size and geometry of Olympic Dam stopes and to cope with the sequential multiple face exposures. This was carried out using numerical modelling software. At the same time, a spreadsheet based design tool was developed suitable for use by mine design - backfill engineers to permit design of individual fill stopes. The design study showed that the maximum strength required, with appropriate conservatism, was 4.5 MPa. The basis of the CAF strength design spreadsheet was that CAF would be available in 0.5 MPa increments from 0.5 to 5.0 MPa. In parallel with the design study, investigations were undertaken to evaluate the options for backfill mix designs. The objective was to determine the most economical mix recipes for the series of strengths ranging from 0.5 to 5.0 MPa in 0.5 MPa increments using all of the aggregate ingredients available at the mine site. This investigation required the testing of a wide range of cementitious and pozzolanic binders with the locally available aggregates consisting of underground mullock, quarried limestone, classified tailings (tails sand) and dune sand. A very recent development has been the elimination of dune sand, which has been replaced with limestone fines from the quarry.
BACKFILL DESIGN Backfill generally consists of assemblages of natural materials; fine ground mineral particles, sand, crushed rocks and a range of cementitious binder materials. These materials contain voids, which are filled with air and/or water. There are two important properties to consider for the design of backfill.
• Flow properties, which enable the delivery and placement of backfill into the underground workings. This is achieved by adding water to form either thin or viscous slurries. The water is also used to hydrate cementitious binders and excess water, where present, drains off.
• Strength properties, which enable the fill to withstand the internal stresses generated by gravitational loading of fill caused either by exposure of free fill faces or from dynamic external sources such as blasting or seismicity. This paper considers only the strength properties of backfill and will focus on the strength required for vertical, freestanding faces. The effects of dynamic loads on fill are not considered. There is little useful quantitative information in this area and this
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tends to support the opinion that well controlled stope blasting practices result in minor fill damage. Backfill is a good attenuator of blast energy and it is intended that dynamic loading conditions will be investigated in future work.
Aspects of backfill geomechanics The fundamentals of soil mechanics are extensively described by Terzaghi and Peck (1967) and Lambe and Whitman (1979). The reader is directed to them for further detailed information. The following discussion is adapted for backfill from Lambe and Whitman (1979). Relative sliding between particles is the most important mechanism of deformation within a fill mass. The resistance of fill to deformation is influenced strongly by the shear resistance at contacts between particles. Determining the magnitude of shear resistance and the factors that influence it are the keys to understanding the behaviour of backfill. A further critical aspect to consider is the interlocking of particles, which is a function of packing density. Interparticle shear resistance is frictional in nature. In some situations, part of the shear resistance between particles is independent of the normal force pushing the particles together, ie even if the normal force is decreased to zero there is still a measurable shear resistance. In such cases, there is true cohesion between particles. In backfill, this property is achieved by the use of binders such as Portland cement and pozzolans. The strength of soil is usually defined in terms of the stresses developed at the peak of the stress strain curve. The Mohr-Coulomb failure law simplifies the envelope to a straight line and can be expressed by:
FIG 1 - Theoretical assessment pf arching mechanism.
where z = vertical coordinate (m) σz = vertical stress (kPa) σx = horizontal stress (kPa) W = half width of stope (m)
τƒ = c + σƒ tan φ
(1)
ρ = material density (t/m3) g = gravitational constant (m/s2)
where τƒ = shear stress at failure along the failure plane
k = ratio of horizontal to vertical stress (sx/sz )
σƒ = normal stress perpendicular to the failure plane
c = material cohesion (kPa) φ = friction angle (°)
c = cohesion
Bloss notes that there are two assumptions;
φ = friction angle The strength of backfill can be therefore be described in terms of the cohesion and the friction angle of the material. The friction angle is a property of the backfill particulate make up and the degree of packing density. The cohesion is a function of the binder addition and the development of chemical bonds between particles. A backfill can be considered to be stable when the shear strength exceeds the current state of shear stress at any point in the fill. Terzaghi (1943) described the mechanism of arching in a two-dimensional column of soil. Terzaghi observed that the vertical stress induced by the self-weight of the soil is transferred to adjacent soil or containing walls. This phenomenon of pressure transfer is known as arching. For a given prism of soil located above a void a stable condition exists when the self-weight of the soil is resisted by the shearing stresses along its lateral boundaries. He stated that the ultimate pressure on the base of the soil is practically independent of the depth of the layer of the soil. If the soil has a trace of cohesion, the soil will not fail into an underlying void. Bloss (1992) describes the two-dimensional force equilibrium equation for a mass of fill in a stope as illustrated in Figure 1. Solving, Bloss gives: (Wρg − c ) σz = 1 − e K tan φ
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− k tan φz W
(2)
• uniform stress distribution; and • full mobilisation of the shear strength of the fill (c + σx tan φ) at the wall contacts. Askew et al (1978) caution that the use of the cohesion term is not referring to the cohesive strength of the backfill but the cohesion developed between the fill and stope wall contact. Where the walls are subvertical and rough, this cohesion should not be less than the backfill cohesion. Where walls are inclined, however, this contact cohesion may be less than the backfill cohesion. The effect of arching and the comparison with self-weight is plotted in Figure 2. The arched vertical stresses are considerably lower than the corresponding depth stress values. Cowling (1983) describes the state of the initial stresses developed in backfill as a result of placement in stopes. He notes that when cemented fills are placed, they are in a liquid state and can have high water contents for some time after they have been placed. Hence, each fill pour shows an initial hydrostatic vertical stress profile. The fill commences curing and the stiffness increases linearly with time. This locks in the stress profile for the lift and the pattern is repeated with each discrete fill run. The modulus continues to increase with time. This generates a characteristic zigzag of shear and horizontal stresses down the centre line of the fill during modelling. This feature is noted by Barrett et al (1978) and Bloss (1992) and is illustrated schematically in Figure 2. The features shown are a function of the modelling process and would be different for discrete fill runs.
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FIG 2 - Comparison of vertical stress in arched and unarched modes.
There are very few in situ stress measurements to back these assumptions of backfill behaviour. Coulthard and Dight (1979) measured hydrostatic loading during fill placement but Cowling (1983) reported that depth stresses (in cured fill) have not been recorded in situ and that arching effects substantially reduce the vertical stress values. Although the description of the zigzag loadings is generally accepted, Cowling comments that there is little direct evidence to support the notion. In summary, the effect of arching significantly reduces the stress levels within fill. Since the stress at any point in the fill can be determined, it is therefore possible to zone the fill strength such that the cement content is varied through the height of the stope in order to match the required performance at that elevation. This then permits the design and placement of a cost optimised backfill.
(3)
where σz = unconfined compressive strength (kPa) ρg = bulk density of backfill (KN/m3) z = depth of backfill from the surface (m)
Vertical slope Mitchell et al (1982) discuss the vertical slope method where φ = 0 for a constant strength fill: σz = ρg H / 2
(4)
where
Backfill design techniques Various authors have published backfill design techniques and these are briefly reviewed in the following section. The techniques discussed are:
• • • •
σz ≥ ρg z
H = overall height of exposed fill (m)
Limit equilibrium wedge– Mitchell Mitchell et al (1982) used physical centrifuge modelling to develop a model for a single exposure, three-dimensional fill mass.
freestanding vertical faces, vertical slopes, limit equilibrium wedge, and
σz =
arching.
Free standing vertical face Mitchell et al (1982) note that the largest shear stresses in backfill are caused by self-weight. The cemented backfill can be designed as a freestanding vertical face. A freestanding wall where the unconfined compressive strength required at any depth in the fill is given as:
γH H 1 + L
(5)
where σz = design UCS H = height of fill exposure, and L = width of fill exposure.
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There are four assumptions with this model:
• the mobilised wall shear stress is equal to the material cohesion;
• the material friction angle is zero; • the fill height is much greater than fill depth; and • stresses are evenly distributed. This formula is commonly applied to backfill design strength estimation. Stone (1993) presents a methodology for the design of backfill using the limit equilibrium approach in the form of standardised charts. A limitation of this technique is that it only applies to single exposures of a fill mass. Even in small and narrow primary-pillar orebodies, most backfill masses are exposed on two opposing faces. The design technique must assume that two wedges are laid back to back and do not interfere with each other. The technique cannot cope with the exposure of adjacent faces, a common occurrence in large mines such as Olympic Dam.
Arching – Terzaghi Askew (1982) and Bloss (1998, personal communication) independently adapted the Terzaghi equation for regular three-dimensional square plan area stopes. Their solutions are shown below as Equation 6.
σz =
( N . k . z .tan φ ) W c W ρ g − 1 − e N W
k.tan φ
AMC has therefore incorporated two factors into the calculation of average vertical stress.
• The loss of contact area due to successive fill exposures is (6)
modified in line with the results reported by Bloss (1998). The loss of arching is reduced to allow for the benefit from fill to fill arching.
• The resultant average vertical stress calculated at a particular
where N = Number of exposed equal sized fill faces W = Width of exposures (m) AMC has further modified the equation to cope with irregular prismatic stopes as commonly encountered at Olympic Dam. The modification is illustrated in Figure 3. For Olympic Dam, prismatic stope shapes are the normal design occurrence. It can also be seen that there will be a difference in stress estimation depending on the sequence of large and small face exposures made. AMC therefore applied a further modification that changed the number and size of exposures to that of contact surface area. During initial fill placement and curing, the fill is confined by all the surrounding rock walls. As each face is successively exposed, the contact area is reduced progressively from 100 per cent contact until all faces are open. This would imply that a stope which has been exposed on all four walls should be considered a free standing prism and therefore subject to depth stresses. Bloss (1998) investigated this issue using numerical modelling techniques. He concluded that, for a regular set of stopes:
• the arching mechanism is still dominant in stabilising the fill even when adjacent extraction is complete;
• the effect of Poisson’s ratio in developing confinement is significant in stabilising the fill;
• if the fill remains primarily elastic, the final stresses after complete extraction are similar even though the path to the final result may vary as a result of differing extraction sequences; and
• for the stope geometries considered, plastic strain levels are close to or at the limit for stability proposed by Bloss (1992). This suggests that, even though vertical stresses are below that for failure, the arching mechanism is on the verge of breaking down due to excessive shear stress adjacent to the wall contacts. This can lead to complete failure of the fill.
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FIG 3 - Modification of Terzaghi analysis to incorporate prismatic stope shapes.
elevation is multiplied by a factor of safety but limited by the maximum depth stress calculation at that elevation. The selection of the factor of safety is a function of the degree of conservatism that is required and a scaling component that converts small-scale test strengths to real world scale exposure performance. AMC currently recommends a factor of safety of 2.0 for the Olympic Dam situation. In the absence of failure data for ODC stopes this is considered to be conservative and it is expected that this can be relaxed with improved experience.
Design tools summary Of the existing techniques, the only method suitable for development and application at Olympic Dam is Terzaghi’s arching. This method can cope with the shape and multiple exposure issues and lends itself to a spreadsheet based design tool. However, before the method was advanced further, numerical modelling was undertaken to investigate the distribution of stresses in Olympic Dam backfill. The results of the previous modelling and the current work were then compared with the Terzaghi results to ensure that the AMC Terzaghi Spreadsheet would provide a simple proxy solution to the complex real world problem.
Numerical modelling Numerical modelling was carried out using MAP3D boundary element software and validated and calibrated to earlier work conducted by Olympic Dam Operations. Backfill material properties were not available from Olympic Dam backfill material at the start of the project. The material properties used in the preliminary analyses were therefore based on published data from Mount Isa by Cowling et al (1983), from Broken Hill by Askew et al (1978) and from Canada by Cundall et al (1978). These parameters are listed in Table 1.
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TABLE 1 Backfill material properties. Material Property
Symbol
Value
Unit degrees
Friction angle
φ
34
Cohesion
c
250
kPa
Modulus
E
500
MPa
The initial modelling was conducted to repeat earlier work and this was used to validate and calibrate the model. Purple 54 was then selected as an example of a tall, slender Olympic Dam stope. The exposures modelled were cumulative in nature. No fill was placed back into the adjacent void and the benefit of fill to fill arching was not included. Therefore these results contain a degree of conservatism. The results of the modelling are shown in Figures 4a to 4d and are graphed in Figure 5. Figure 4a shows the geometry of Purple 54 stope, 140 m high by 15 m long by 85 m wide. The stope has been modelled in seven segments, each 20 m high. The stope is to be exposed on the two widest opposing faces. Figure 4b shows vertical stress distribution prior to exposure of any fill faces. Note that significant arching is occurring which keeps stresses levels at around 300 kPa with a concentration at
the toe of around 600 kPa. This is well below depth stress of around 3350 kPa. The zigzag vertical stress profile has been locked into each of the 20 m segment boundaries representing discrete fill pours. In Figure 4c the first wide face has been exposed. There has been a small increase in stress at the toe of the exposure and a small decrease in the upper part. At this point, around 57 per cent of the fill perimeter is in contact with rock and arching is significant. In Figure 4d the second face has been exposed and only 15 per cent of the fill perimeter is in contact with rock. Vertical stress has increased to 1200 kPa but is still well below depth stress levels.
Comparison modelling
of
Terzaghi
Method
with
numerical
The numerical modelling using MAP3D confirmed magnitudes of stress level development in the backfill as successive exposures were made. Figure 5 plots the vertical stress distribution in Purple 54 against the Terzaghi method and correlation in the distribution and magnitude of stresses can be seen. For the one and two exposure cases the maximum average vertical stress at the foot of the fill mass calculated by Terzaghi method and MAP3D are similar. At all other points Terzaghi estimates a higher value. It can be seen that in the case of the second exposure, arching is starting to break down and the maximum values are around one-third of depth stress values.
FIG 4A, B, C and D - Results of numerical modelling of Purple 54 stope. (Refer to the CD-Rom for colour explanation).
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FIG 5 - Comparison of Terzaghi method with numerical modelling.
AMC concluded that the Terzaghi method provides a valid approximation to the distribution and magnitude of average vertical stresses in backfill. Since the values derived are conservative, it can therefore be used as the basis of a simplified design tool.
Development of the backfill design strength tool The primary objective of this part of the project was to develop an effective method to determine the required strength of backfill for use in open stope mining at Olympic Dam. Most of the backfill design techniques described earlier do not allow for the complex geometry and sequencing of exposures of fill masses as experienced at Olympic Dam. Neither is it practical to run numerical modelling for every fill exposure option. The results of the study show that the Terzaghi method can meet all of the design requirements. It was therefore decided to develop a spreadsheet based version of the 3D Terzaghi equation. The software was designed to run as a simple Excel worksheet without macros. The following sections document the underlying geomechanics assumptions and the software design criteria.
AMC Terzaghi spreadsheet The AMC Terzaghi Spreadsheet is designed to print out on a single landscape sheet of A4 paper and contains no macros. There are a number of calculation areas outside of the main print area. The sheet is split into several functional areas and each of these will be discussed in turn. In general, text in blue on an indented pale green background represents data entry. Text in black on an indented pale yellow background represents calculation results. Other numbers are intermediate, advisory calculations. Units are indicated where appropriate. Figure 6 illustrates the general assumptions.
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The key geotechnical data entries are friction angle and Poisson’s ratio. The fill density is the wet bulk density of the fully drained backfill. Depth zone increments permit the user to control the thickness of fill zones in the backfill strength model. The choice of this number is one of convenience and the practicality of control of fill placement into stopes. AMC would tend to recommend ten metres as a lower limit and 25 metres as the upper limit. CAF is prepared to a number of discrete recipes of generally increasing binder content with an increase in strength. Depending on the controls available it is possible to mix batches with incremental strengths of between 250 and 500 kPa UCS. For Olympic Dam the batch increment selected is 500 kPa, ranging up to 4500 kPa maximum. The final entry is the Factor of Safety, a term which loosely describes the degree of conservatism in the design process. AMC favours a probabilistic approach to failure assessment where the design process, the lab scale test data and their statistical variance and back analysis to a large data set of full sized fill exposures which exhibit various degrees of performance and failure are considered. Olympic Dam has no history of major fill failures. This, combined with evidence of very high test sample strengths, renders this approach unsuitable at this time. It was considered that a Factor of Safety of 2 should be applied until real performance data emerged. At some time in the future the Factor of Safety could then be adjusted or the process modified. Figure 7 shows the entry of stope information. The geometry of the stope is entered in the stope information section. The geometry is designed to cope with regular prismatic vertical stopes. Length is defined as the stope plan distance along strike of the orebody and width is the plan distance in the perpendicular horizontal direction. Height is generally the full vertical stope height. Frequently however, the exposure of the fill mass is not over the full height of the stope. In this case, the height of the exposure can be entered. In the examples that
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FIG 6 - General Assumptions.
FIG 7 - Stope information.
FIG 8 - Exposure geometry.
follow a square plan stope, 40 m wide by 160 m high has been selected with four successive exposures. Figure 8 shows the exposure geometry. Details of sequential fill exposures are entered in the exposure geometry section. For each of up to four exposure faces, the width is entered in sequence of exposure. This value is then adjusted to permit fill to fill arching modification in the following calculations and a value for the length to width factor (B`) is computed. The perimeter removed is then calculated as a percentage and factored back into the number of intact walls. Figure 9 shows the results of the calculations. The average vertical stress is computed for each depth increment for successive exposures. The maximum average vertical stress at each depth increment is recorded along with a
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corresponding value for depth stress for comparative purposes. Figure 10 shows the computed average vertical stresses for each exposure plotted as a graph. Note that the fourth exposure is still exhibiting a curved shape that shows the continuance of fill to fill arching. The strength required at that depth increment is then calculated as the lesser of the average vertical stress multiplied by the Factor of Safety and the depth stress. In order to determine the strength of the fill to be placed in that depth sector, this value is rounded up to the next fill strength increment. Figures 11a to Figure 11d show the fill strength zones required for one, two three and four exposures stopes. Taking the height of each of the computed zones it is then possible to develop a schedule of fill placement and a cumulative volume of fill required to be placed in each zone.
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BACKFILLING STRATEGY AND MIX DESIGN DEVELOPMENT: PHASE 1 Overview of the Olympic Dam mining method and backfilling requirements. The expansion strategy for Olympic Dam has evolved from the proposals considered by a scoping study team in 1993 to the detail designs that have been constructed and commissioned by OEP in 1998 and subsequently put into operation by ODO.
Early in this review process, determination was reached that retained the long hole open stoping mining method. As a consequence, mining of ore continued with this historical practice, which similarly enshrined CAF as the predominant fill material. With this mining method, stopes are generally production drilled from access drives which are generally 50 - 60 metres apart vertically and allow both uphole and downhole drilling. A typical stope consists of a bottom level extraction drive which allows LHDs to ‘bog ore’ from multiple extraction points and a top level drill drive which also allows the ‘slot’ (internal rise) to be drilled which creates the void to allow the blasting of the ore to occur.
FIG 9 - Calculation of average vertical stress.
FIG 10 - Graph of average vertical stress.
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FIG 11A, B, C and D - Graphs of average vertical stress for one, two, three and four exposures.
The typical geometry for a stope is an upside down milk carton except the stope is generally rectangular rather than square. Dimensions taper up from approximately six metres wide at the base (for extraction purposes) and attain a maximum stope width of say 35 metres, a length up to 40 metres and a height from 60 to 270 metres determined by orebody characteristics and mine design considerations. The ‘fill void’ for a stope ranges in volume from a minimum of 50 000 m3 to the largest designed to date of approximately 250 000 m3 (270 m high). Traditionally stopes have been filled with CAF which were of a constant strength, mixed to attain at least 5.0 MPa and probably achieving in excess of 10 MPa (based on very limited in situ coring of one or two stopes). The predominant fill mass historically was CAF with some fill void being filled with underground development waste or mullock where there was no adjacent stope to be mined (and hence the stope face not being exposed). Wherever CAF is exposed by an adjacent mined stope, it is required to remain fully supported with no face fall off.
Summary and observations: Phase 1 During the pre-feasibility and feasibility study stages for the expansion of Olympic Dam, particular emphasis was placed on defining CAF properties, utilising existing aggregates available on site. Particular emphasis was placed upon defining the CAF strength characteristics when utilising ‘tails sand’ to replace dune sand as the regular fine aggregate and a significant amount of the coarse aggregate. A typical graphical representation of test results is given in Figure 12 where UCS (Uniaxial Compressive Strength) on the vertical axis is plotted against per cent binder (cement and flyash) on the horizontal axis. In general the following observations can be made from Figure 12 and the Phase 1 investigations.
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• Testwork was undertaken with 30 mm crushed mullock (and limestone where tested) which was not truly representative of the crushed rock used in the actual CAF mixing conducted in the old batch plant and planned for the new mixing plant. The aggregate top size was determined by the test cylinder size rather than investigating the intended process being utilised. The on site fixed crushing plant was operated in an open circuit configuration which resulted in maximum lump approaching 120 - 150 mm. Whilst this top size is too big, it was important to understand the maximum lump size available from closed crushing and a properly graded aggregate - generally, the smaller the rock top size, the greater the crushed rock cost and this is a non-linear cost as rock size reduces.
• All testwork results were recorded and reported on a percentage dry basis. Simple testwork using a small (builder’s) concrete mixer revealed that water requirements varied considerably between crushed mullock and limestone as well as the grading of the course aggregates. It was apparent that ‘mix designs’ required specification of all the ingredients inclusive of the water added (both directly and indirectly).
• All testwork undertaken was on the basis of cement and flyash being used in the proportions of 1 to 2. No testwork results/reports existed to substantiate this nor define an appropriate relationship of strength versus percentage binder (cement and flyash) with variable proportions, eg 1 to 2, 1 to 3, 1 to 4, etc. The economics of landed cement and flyash cost meant that for a given CAF mix design strength, the quantity of flyash should be maximised to meet the strength requirement because the landed cost was five to six times cheaper.
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FIG 12 - Strength variations with cement content for samples with crushed mullock.
• All testwork strengths were for 56 days; three test cylinders
• CAF mix designs to attain the minimum specified strengths
were crushed at 14, 28 and 56 days and the results averaged to yield a single strength value for the elapsed curing time. There were no results/reports, which investigated and defined post ‘56 days’ curing times. A literature search revealed that this was most probable and clearly required defining. The most appropriate way to represent this phenomenon was to graph UCS (MPa) versus curing time (days). This knowledge should have been established and defined.
did not differentiate between crushed mullock and crushed limestone. Simple testwork with the concrete mixer had revealed that there was at least a two to three per cent w/w (weight wet) difference in water requirements between CAF mixes with mullock compared to limestone.
• All testwork was undertaken using traditionally available binders (cement and flyash); cement was supplied from Adelaide and flyash (Zone 1) ex the Port Augusta Power Station (coal fired). It is well documented that ground granulated blast furnace slag when processed correctly will develop pozzolanic properties. A well-established blast furnace exists at Whyalla (BHP) and no serious discussions had been entered into that resulted in an appropriate test program utilising GGBFS and the locally available aggregates at Olympic Dam. The testwork undertaken in 1995 had provided some valuable and useful information regarding mix designs, which had been carried forward to process and plant design. However, there seemed as many unanswered questions as those that were answered. The recommendations that were carried forward are in essence summarised in Figure 13. The main features are:
• CAF was to be mixed at the respective strengths of 2.0, and 3.5 MPa using readily available aggregates from site. Whilst a general strategy for selecting the CAF strength existed (ie stronger at stope bottom and weaker towards the top depending upon number of exposed faces), there was no well-defined systematic design method for determining the minimum specified CAF strength for all stope mining situations.
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The real concern was the over simplification of the mix designs and thus the subsequent process and actual plant design. These and other important factors revealed from critical review conducted in July 1998 prompted a rapid investigation and testwork program. At that point in time the upgraded plant was being constructed, with commissioning planned for early to mid September 1998. The clock was well and truly ticking and the delivery of CAF to fill the underground fill (stope) voids was essential to the highly planned and well-coordinated mining at Olympic Dam. Any serious delay to the reliable delivery of CAF would have impacted significantly on the mining operations during the critical underground production ramp-up phase to meet the new production capability of the process plant. Any delay to a stope going into production due to late backfilling was untenable.
BACKFILLING STRATEGY AND MIX DESIGN DEVELOPMENT: PHASE 2 Overview: Phase 2 By early August 1998 the high level strategy was to:
• Establish a clearly defined design process, which allowed the fill mass to be engineered. The outcome of the design process was to determine the minimum specified CAF strength.
• Develop mix designs that allowed CAF to be mixed continuously in a modern process controlled plant and delivered using semi-trailer tipper trucks.
Brisbane, Qld, 29 October - 2 November 2000
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ENGINEERING THE NEW OLYMPIC DAM BACKFILL SYSTEM
FIG 13 - CAF Plant process flow diagram.
The outcome of the testwork was the development of mix designs using mullock as the main coarse aggregate ingredient due to the significant surface stockpiles available from the OEP mine development and normal mining operations. Preliminary work was commenced on mix design testwork with limestone, which would be mined from a nearby quarry and required after the surface mullock stockpiles were exhausted. Similarly, testwork was undertaken using the small concrete mixer.
• Define a stope filling strategy which attempts to maximize the uncemented fill mass relative to the cemented fill (CAF) mass. The intent here would be to refine the stope filling design process so that uncemented fill could be maximised. Thus, utilisation of surface rockfill and hydraulic (sand) fill were potentially suitable uncemented fill types that could be delivered quickly and efficiently from the surface, in addition to the underground development waste. These refinements were not the immediate priority due to the need to get the upgraded CAF mixing plant constructed and commissioned as quickly as possible. It was obvious that the promised completion dates would not be achieved and the work program was split into more sensible packages that gave greatest priority to completing all facilities directly associated with CAF mixing.
Operating experience: November 1998 to April 2000 A simplified flow schematic for CAF production is shown in Figure 14. This schematic is applicable to the time frame from
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November 1998 to April 2000. This was when the fixed crushing and screening plant was operational. From mid April 2000 an in-pit track mounted (mobile) two-stage crushing and screening plant has been operational. It was supplied by Nordberg and is owned by Brambles Industrial Services and supplies crushed rock under the ‘Backfill Services Contract’ along with the mixing and delivering of CAF. For the stated period of time, the following milestones and observations are relevant Upgraded Backfill (CAF Mixing) Plant Commissioning Milestones:
• CAF Mixing in pugmill: 1 September 1999 (plan), 1 November 1998 (actual);
• New Binder (cement and flyash) Silos: 1 October 1998 (plan), 15 January 1999 (actual); and
• Neutralised Tails Sand Supply System: 1 November 1998 (plan), 20 August 1999 (actual). The pugmill was re-designed and upgraded in April 1999 with respect to wear materials for the mixing chamber. The pugmill was further redesigned and upgraded in August 1999 with respect to bearings and mixing chamber sealing along with a complete overhaul. The mix designs are entered into a recipe menu in the process control system for the required CAF strength as specified by the ‘CAF delivery schedule’. Specific mix designs exist for CAF strengths from 0.5 MPa to 5.0 MPa in 0.5 MPa increments, utilising the following ingredients:
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G BALDWIN and A G GRICE
FIG 14 - Simplified flow schematic for CAF production.
• crushed mullock/dune sand/cement and flyash/water; • crushed mullock/dune sand and tails sand/cement and flyash /water;
• crushed limestone/dune sand/cement and flyash/water; and • crushed limestone/dune sand and tails sand/cement and flyash/water. Tails sand is supplied as neutralised tailings sand slurry at around 50 - 53 per cent solids (for the current single stage neutralisation circuit). Cement and flyash are normally mixed in the ratios of 1:2 or 1:3 depending upon the availability of flyash from the Port Augusta Power Station. Properly graded crushed rock is essential to ensure the repeatability of CAF mixing using process control and delivery using semi-tipper trailers to dropper holes. The fixed crushing and screening plant was successfully operated in closed circuit from March 1999 onwards after the successful commissioning of a ‘trash removal facility’ within the circuit. This capability is essential whenever underground waste mullock is used so as to remove the significant quantities of metallic and non-metallic rubbish that reports with the mullock from underground. The fact that the crushing and screening plant did not have this capability was a significant oversight. Twelve test cylinders are taken on every shift so that testing can be conducted at 14, 28, 56 and 90 days elapsed curing time. These cylinders need to be cured in a controlled environment water baths are preferred at Olympic Dam. The water bath temperature is controlled at 35°C to approximately represent the
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‘in-stope temperature’ where curing of the CAF occurs. The rock temperature is generally above 40°C in the ore zones and CAF is cured in a hot moist environment which is also affected by aquifer water running into the stopes (it is intersected by the CAF dropper hole). Despite the benefits of a continuous, process-controlled CAF mixing plant, CAF consistency (hence water content) is best determined by the simple and reliable ‘slump test’. This test is conducted on an hourly basis when CAF mixing occurs and the results logged - it is part of the contract key performance indicators (KPI) for the ‘Backfill Services Contract’. Brambles Industrial Services were the successful tenderer for the new contract awarded in September 1999. Slump of the CAF is specified to be 150 - 180 mm and when ingredients are mixed and controlled correctly, the slump of the CAF can be consistently between 160 - 170 mm. Water management against mix design is the most critical parameter to be controlled to ensure delivery of the minimum specified strength for CAF. Owing to the delivery of CAF in semi- tipper trailer trucks (payload currently 12 - 13 m3 per trailer), correct slump means no segregation occurs when the CAF is transported over gravel roads where a loaded one-way trip can be in excess of 2 km. The purpose designed semi-tipper trailers (supplied under the new contract and owned by Brambles) are capable of a 15 m3 payload. This transport method is very flexible and effective when a prompt stop of CAF delivery to any stope is required for whatever reason.
Brisbane, Qld, 29 October - 2 November 2000
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ENGINEERING THE NEW OLYMPIC DAM BACKFILL SYSTEM
Drilling of the ‘CAF/ Fill Dropper Holes’ is now always an eight inch percussion drilled hole from the surface, which is then reamed to 12.25 inch diameter. Wherever possible, holes are drilled directly into a stope to provide the simplest and most cost-effective delivery method. Hole depth is between 350 - 500 metres and to an accuracy of ±1 per cent of depth. For a period of nine months dropper holes were drilled at eight inch as a simpler and lower cost delivery hole. Despite the improvement in the crushed rock grading and quality from the closed circuit operation of the crushing and screening plant (implemented March 1999), some dropper holes continued to fail prematurely by blocking up and the practice was abandoned in June 1999. The ‘CAF/ Fill Dropper Holes’ are steel cased to hard rock from the surface (generally 40 - 50 m) and a tundish (= a funnel with screening grates) is installed on the top of the casing into which the CAF delivery trucks tip directly. The tundish design is now at ‘mark 3’ and is very reliable from a wear perspective. The dropper hole below the steel casing is in the natural rock and generally lasts between 80 000 - 100 000 m3 of CAF and/or surface rockfill delivery. The design of the new semi-tipper trailers and tundish allows both CAF and SRF (surface rockfill) to be delivered to the stope in accordance with the fill design and delivery schedule. The ‘process control system’ was upgraded in July/August 1999 to accommodate the supply of the neutralised tailings sand as a slurry (NTSS). The basis of the process design was that a dry sand product would be available to supplement the NTSS supplied from the process plant. This was required due to the variable density and supply reliability from the process plant deslime circuit. The quantity of tails sand and neutralised liquor is calculated from the slurry flow and density and compared with the mix design values for ‘sand’ and ‘water’. The shortfall in tails sand is made up with a dry sand product and similarly with the water. There has been a ‘trim water’ circuit added to allow the operator to adjust the water requirement set point due to change in the ‘slump’ of the CAF. Surface rockfill (SRF) was first delivered under the new design process to Purple 80 (stope designation) from October to December 1999. The stope was filled to a defined level as per the fill design with CAF. Then, CAF and SRF were simultaneously delivered down separate fill dropper holes so that a CAF plug was created on one side of the stope and straight rockfill on the other. This was due to the stope only having a single face exposure. This method means stope fill void is filled with a significantly cheaper fill mass for the SRF (approximately 60 70 per cent cheaper) compared with the CAF.
Operating experience from April 2000 and future developments To improve the quantity, quality and reliability of tailings sand supply from the process plant, the ‘deslime circuit’ was redesigned and simplified. This was implemented and commissioned in May/June 2000. The quantity of tails sand has now reached in excess of 50 per cent of the sand requirement for the CAF mixes. The ‘process plant deslime circuit’ supplies acid tailings sand slurry (ATSS) to the backfill plant where it is neutralised with a ‘milk of lime slurry’ prepared in a dedicated lime slaking plant. The neutralised tailings sand slurry (NTSS) is then supplied to the pug mill on demand from storage tanks. A plan has been developed to reduce the dune sand consumption to zero for cost and environmental reasons. This will ultimately require the ‘mix design sand per cent to be supplied as tails sand at 0.75 and limestone fines at 0.25 which will replace the dune sand. The limestone fines will be produced by screening (scalping) out the fines in the quarried limestone drill and blast feed rock to the new mobile in-pit crusher. Field
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trials have confirmed around 30 per cent by weight can be screened out which will sustain the intended replacement process. The new in-pit mobile crushing and screening plant was supplied by Nordberg to Brambles who are the owner. It is capable of 600 tph and utilises a Nordberg Lokotrak Configuration of LT 125 primary crusher and LT 1500 secondary crusher. Mixing trials with the pug mill have also been conducted to confirm the process control and CAF mixing consistency using the limestone fines. To increase the tailings sand supply to the backfill plant so that the 0.75 proportion can be achieved, a new upgraded capacity pump station will be constructed and commissioned at the tailings disposal area. This pump station will supply tailings to the ‘deslime’ circuit reliably and at the required rates, even at the upgraded CAF mixing rates required over the forthcoming years. The CAF Mixing Plant production amounts on a per annum basis (actuals/ forecast) are given in Table 2. TABLE 2 Fill schedule. Year
Fill volume m3
Fill tonnes
1999
1 256 000
2 888 000
2000
1 450 000
3 371 250
2001
1 850 000
4 347 500
2002
1 950 000
4 582 500
2003
2 050 000
4 817 500
2004
2 150 000
5 052 500
2005
2 200 000
5 170 000
It is believed that the CAF Mixing Plant maximum reliable production capacity is around 2 200 000 m3 per year. This equates to a continuous delivery of 335 m3 per hour at 75 per cent use, which is considered a very achievable figure with the current redesigned pug mill and binder delivery feeders. It will be necessary to conduct some very minor upgrades to achieve this eg conveyor, feeder and pump speed-ups. The current neutralisation process is single stage using ‘milk of lime slurry’ produced from slaking quicklime. It is effective from a process perspective; however, the costs are high due to the cost of lime. Waste products from a soda ash works in Adelaide, which have neutralisation properties, are available. Laboratory scale testwork has revealed their effectiveness in neutralising ATSS (acid tailings sand slurry) from 1.0/1.5 pH to 4.0/5.0 pH. This will allow the neutralisation of the ATSS to be effected in a two-stage process. The existing lime slaking facility will be used to trim the pH value up to the set point of 7.0 which is then classified as NTSS (neutralised tailings sand slurry) and ready to be supplied to the pug mill for CAF mixing. Technical investigations have been conducted into the use of ‘ground granulated blast furnace slag (GGBFS)’. An agreement has been entered into with BHP and testwork conducted by the CSIRO on GGBFS from both Newcastle and Whyalla. This product will be able to in essence replace cement as the main binder in CAF mixes. Depending upon the economics, it may be possible to replace some or all of the flyash. The other application for this product is in neutralisation of acid tailings sand slurry and acid liquors. At the time of writing this paper, BHP - Whyalla and WMC - ODO have tendered for the supply of GGBFS on a contract basis. A final decision on the use of GGBFS will be made late in year 2000.
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CAF Displacement Rockfill (CDR) is being developed for stopes which are currently planned to be 100 per cent CAF. This method is based on the design requirement of only having a 10 m CAF pillar. This need has been developed and defined for stopes where CAF is at one end and rockfill at the other (as was the case for Purple 80). Thus, it is possible to introduce SRF (surface rockfill) into the centre of a stope and generally have the SRF remain in the stope centre. SRF is delivered by semi-tipper truck along with CAF to a defined ratio, eg 1:10. Whilst this method may only displace around five per cent CAF, the cost savings are appreciable if this can be applied to only 50 per cent of stopes currently receiving CAF. For 50 000 m3 of CAF displaced, the estimated savings would be $ 500 000 - $ 600 000 per year.
CONCLUSIONS The results of the design study project have been to define the required minimum CAF strengths and produce an interactive design spreadsheet able to be used by mine design/ backfill engineers on site. The CAF mix designs to meet the required minimum CAF strengths have been developed on-site for a wide range of aggregates that are available. These have been mixed in the new CAF mixing plant and delivered to the U/G stopes by well proven methods developed at Olympic Dam This investigation and implementation process has generated some new insights into backfilling at Olympic Dam, especially for cemented fill (CAF) and for surface rockfill (uncemented fill). The project has delivered a simple and effective design tool, which is used on a routine basis to determine zoned CAF fill strengths. The design mixes that have been developed ensure that these strengths are then achieved at the most economical cost. The unit cost of CAF as backfill has reduced approximately 15 20 per cent since the new design, mixing and delivery processes have been optimised. Monitoring of the performance of the new backfill system is in progress. Olympic Dam has successfully engineered the backfill system to meet the needs of the expanded mining operation. This has been achieved through improved design processes, rigorous investigations and the introduction of an innovative approach to backfill mixing and placement. As future mine fill needs increase in quantity and change in requirements, the knowledge base of personnel about the backfill system ensures that Olympic Dam is well positioned to meet those changes and continue to deliver a cost-effective fill mass at a reducing unit cost per cubic metre of void filled.
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ACKNOWLEDGEMENTS The management of WMC – Olympic Dam Corporation and Australian Mining Consultants are thanked for giving their permission to publish this paper. Ken McNabb of MINCAD Systems Pty Ltd carried out the modelling of the backfill and the authors acknowledge the previous work of both Dick Cowling and Martyn Bloss in particular, towards the understanding of cemented rockfill behaviour and upon which much of this work has been based.
REFERENCES Askew, J E, McCarthy, P L and Fitzgerald, D J, 1978. Backfill research for pillar extraction at ZC/NBHC, in Mining with Backfill, 12th Canadian Rock Mechanics Symposium, Sudbury, Ontario, May, CIM Special Volume. Askew, J W, 1982. Lecture notes, Rock mechanics, Western Australian School of Mines. Barrett, J R, Coulthard, M A and Dight, P M, 1978. Determination of fill stability, in Mining with Backfill, 12th Canadian Rock Mechanics Symposium, Sudbury, Ontario, May, CIM Special Volume. Bloss, M L, 1992. Prediction of cemented rock fill stability – Design procedures and modelling techniques, PhD Thesis, Department of Mining and Metallurgical Engineering, University of Queensland. Bloss, M L and Greenwood, A G, 1998. Cemented rock fill research at Mount Isa Mines Limited, 1992 - 1997, in Proceedings Minefill ‘98, Sixth International Symposium on Mining with Backfill, pp 207-215 (The Australasian Institute of Mining and Metallurgy: Melbourne). Coulthard, M A and Dight, P M, 1979. Numerical analysis of fill pillar stability – Two dimensional calculations of initial stresses (Support and stabilisation of stopes project), CSIRO Division of Applied geomechanics Technical Report No 86 (CSIRO: Melbourne) Cowling, R, Auld, G J and Meek, J L, 1983. Experience with cemented fill stability at Mount Isa Mines, in Proceedings International Symposium on Mining with Backfill, Lulea, 7-9 June, (Balkema). Cundall, P, Shillabeer, J H and Herget, G, 1978. Modelling to predict backfill stability in transverse pillar extraction, in Mining with Backfill, 12th Canadian Rock Mechanics Symposium, Sudbury, Ontario, May, CIM Special Volume. Lambe, T W and Whitman, R V, 1979. Soil mechanics, SI Version, (John Wiley and Sons, Inc). Mitchell, R J, Olsen, R S and Smith, J D, 1982. Model studies on cemented tailings used in mine backfill, National Research Council of Canada, Can Geotech J, Vol 19. Mitchell, R J, 1983. Earth Structures Engineering, (Allen and Unwin Inc: 9 Winchester Tce, Winchester, Mass 01890, USA). Stone, D M R, 1993. The Optimization of Mix Designs for Cemented Rockfill, in MineFill 93, pp 249-253 (South African Institute of Mining and Metallurgy: Johannesburg). Terzaghi, K and Peck, R B, 1967. Soil mechanics in engineering practice, (John Wiley and Sons, Inc). Terzaghi, K V, 1943. Theoretical Soil Mechanics, p 510 (John Wiley: New York).
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Backfill Research at the Western Australian School of Mines C Wang1 and E Villaescusa1 • determination of detailed fill properties covering Uniaxial
ABSTRACT This paper summarises research work carried out at the Western Australian School of Mines on mine backfill properties in the last few years. The fill properties presented relate to cemented aggregate fill and cemented tailings fill. Emphasis of these studies has been placed on the composition of backfill materials including aggregate particle size distribution, cement dosage and influence of cement admixtures. The unique features of the backfill research carried out at WASM during the last few years are the accomplishment of a study into the properties of large cemented rockfill samples by triaxial testing and the investigation of the properties of fill material made from highly saline ground water prevailing in the Goldfields region of Western Australia. Exploring different cost-effective binding agents for mine backfill was also carried out at WASM. Special effort has been put on the development of fill materials compatible with hydraulic and paste fills. During the process, specific methods and procedures for laboratory testing of backfill material have been formalised. The on-going backfill research on sample scaling is summarised and the future direction of WASM’s backfill research focus is outlined.
INTRODUCTION With a reputation as the most practical mining school in Australia, the Western Australian School of Mines (WASM) has been consistently seeking technical developments to facilitate the mining industry in Western Australia to keep up with the advancement of mining technology around the world. The school is situated in Kalgoorlie, the centre of a very active mining area in Western Australia. In this scenario, advanced mine backfill technology is in urgent demand in order to increase stoping recovery ratio and control ground stability. Consequently, in early-1996, WASM started a mine backfill research program. To-date, three subprograms have been completed:
• properties of cemented aggregate fill; • properties of cemented highly saline tailings fill; and • high water and quick setting cementitious material based aggregate or tailings fill. The development of specific facilities for carrying out backfill research has also been one of WASM’s research endeavours. The establishment of a triaxial testing system for fill samples up to 200 mm in diameter and 400 mm in length is an exclusive feature of WASM’s backfill research capability. Pursuing high standard geotechnical backfill design by incorporating WASM’s research results on backfill sample scaling into numerical modelling is now another focus of WASM’s backfill research commitment. Over the last few years, a fully equipped backfill research laboratory has been set up at WASM. This laboratory is capable of undertaking standard testing of properties for different types of backfill. Its capability mainly consists of the following three components:
• fill material preparations including rock crushing, particle size distribution (PDS) analysis, tailings solids and water separation, permeability determination and material drying;
• fill sample preparations composed of materials mixing, fill sample moulding (with diameters of 38 mm, 50 mm, 70 mm, 100 mm, 150 mm, 200 mm and 300 mm), fill sample curing with both temperature and humidity control (either in curing tank or curing room) and sample end trimming; and 1.
Western Australian School of Mines, PMB 22, Kalgoorlie WA 6430.
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Compressive Strength, Triaxial Compression Test (at a confining pressure range of between 0 to 1.0 MPa), Tensile Strength, Elastic Modulus, Poisson’s Ratio, Friction Angle, Cohesion, Moisture Content, Bulk Density, Void Ratio, Slump and Workability. One of the difficulties encountered by backfill researchers has been the lack of identified test standards to follow for laboratory testing. Therefore, part of the effort of the backfill study carried out at WASM has been focused on rationalising laboratory test methods and procedures both for cemented rockfill and cemented tailings fill. With assistance from external experts and combined with research experience accumulated at WASM, a set of standards for laboratory backfill testing has been completed (see Appendix 1).
ASPECTS AFFECTING PROPERTIES OF CEMENTED AGGREGATE FILL (CAF) A proper evaluation of the fundamental properties of CAF is very important to conduct numerical analysis, economical evaluation and quality control of a backfill operation. The scope of this study was to determine the potential factors influencing the fundamental properties of CAF, mainly strength character.
Influence of aggregate strength on CAF strength Generally, strong rocks are used to produce aggregates for CAF as it is commonly recognised that the stronger the aggregates, the higher the strength of the cemented aggregate fill. However, laboratory tests indicate that the strength of CAF made of a combination of strong aggregate and relatively weak aggregate is noticeably higher than that obtained from a fill made solely of either strong or weak aggregate. This was discovered while an Australia-China joint research project, the Daye Iron Mine Backfill Project, was being carried out at WASM in late-1996. Owing to the restraint of fill material resources (Golosinski and Wang, 1996, 1997), only two types of suitable waste rock were available at the Daye mine site, namely diorite and marble. As a result, it was decided to limit the tests to pure diorite and to a mixture of equal proportions of diorite and marble. Because of insufficient aggregates supplied by Daye to WASM, locally available aggregates, dolerite and marble, from Kalgoorlie area were used to conducted the tests. The UCS of dolerite and marble used were 127.0 MPa and 68.6 MPa, respectively. The influence of aggregate type and strength character on the CAF strength was revealed by the testing results of three batches of CAF made of either dolerite, marble or mixture of equal amount of dolerite and marble with the same particle size distribution at the same cement dosage. As presented in Table 1, a CAF made of dolerite having four per cent cement content has a 28-day UCS 26 per cent higher than that of a CAF made of marble with the same amount of cement. This confirms that strong aggregate can result in higher strength of CAF. However, Table 1 also indicates that the UCS of samples made of a mixture of equal proportions of dolerite and marble is about 30 ~ 55 per cent and 60 per cent higher than that of a sample made solely of either dolerite or marble, respectively.
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TABLE 1 Properties of CAF made of different aggregates. Aggregate type Cement content (% by weight of aggregate)
Dolerite
Marble
Dolerite and marble
3
4
5
6
4
3
4
5
6
3-day UCS (MPa)
2.7
3.7
4.8
6.2
N/A
4.0
4.7
6.2
8.7
7-day UCS (MPa)
3.1
4.5
5.8
7.8
N/A
4.3
5.4
7.1
9.5
28-day UCS (MPa)
4.3
6.3
9.2
12.8
4.98
5.3
7.5
10.2
14.3
Tensile strength (MPa)
0.7
0.8
1.1
1.8
N/A
0.8
0.9
1.2
1.5
Bulk density (tonne/m3)
2.24
2.26
2.29
2.34
2.41
2.25
2.28
2.32
2.37
0.2366
0.2201
0.2096
0.1875
0.152
0.2270
0.2059
0.1885
0.1601
void ratio
An investigation (Golosinski and Wang, 1997) into the factors that made a CAF sample consisting of a mixture of strong and weak aggregate stronger than a CAF sample made solely of either strong or weak aggregates was conducted in the laboratory. This was achieved by simulating the procedure of CAF sample preparation. Dolerite aggregates and a mixture of equal proportions of dolerite and marble aggregates having the same PSD, with neither water nor cement, were mixed separately in a rotating tumbler for five minutes in order to match the process of CAF sample preparation. The aggregates were placed separately in a cylindrical mould similarly to the CAF sample preparation process. As no cement and water were used in the simulated tests, sieve analyses were able to be carried out. This indicated an obvious difference between the modified PSD of the combined aggregates and that of an individual type of aggregates as given in Table 2 and Figure 1. The results suggested that the aggregates resulting from the simulated mixing and compacting generated more fines. This investigation revealed that more fines are produced during the preparation of CAF samples with aggregate containing equal proportions of dolerite and marble than that of CAF samples made of dolerite only. This generation of additional fines during mixing and moulding enables the CAF samples to be densely packed resulting in a higher bulk density and a lower void ratio as shown in Table 1. Hence, the UCS of such samples will be higher than that of samples made purely of dolerite. Consequently, it is reasonable to predict that the use of a combination of hard and relatively weak aggregate in mine backfill practices can lead to a higher quality fill mass because the procedure of fill preparation, delivery and placement at a mine site can bring about the same effect on aggregate composition as that described on the simulated tests.
FIG 1 - Aggregate particle size distribution.
Influence of aggregate PDS on CAF strength The influence of aggregate PSD on the properties of CAF was also examined during the Daye backfill property study.
TABLE 2 Particle size distribution of aggregate (percentage retained, per cent, by weight). Sieve size (mm)
PDS of original aggregate Trial
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PDS after mixing and molding
Adjusted
Dolerite
Dolerite/Marble
40
0
0
0
0
26.5
20
10
9.6
8.1
19.0
40
25
24.4
21.3
9.5
60
45
44.0
39.5
4.75
75
65
63.8
61.6
1.7
85
80
79.5
77.7
0
100
100
100
100
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BACKFILL RESEARCH AT THE WESTERN AUSTRALIAN SCHOOL OF MINES
The uniaxial compressive strength of CAF, made of dolerite with a trial PSD (initially designed based on the backfill practice experience at Kidd Creek (Yu and Counter, 1983; Farsangi and Hara, 1993) and Mount Isa mines (Bloss, 1996; Grice, 1989) was firstly determined. Based on these UCS results, the PSD was adjusted to give an improved CAF UCS. The trial and adjusted PSD are given in Table 2 and graphically presented in Figure 1. Two aggregate to cement ratios were used in the investigation, namely, 100:4 and 100:3 by weight. Compressive strength of cylindrical CAF samples of 150 mm diameter and 300 mm in length at a curing age of three, seven and 28 days are presented in Table 3. The results indicated that, for a four per cent cement content (by weight of aggregate), the aggregate samples with an adjusted particle size distribution gave significantly higher strength than those samples with a trial particle size distribution. Similarly, a CAF sample having three per cent cement by weight of aggregate and adjusted PSD had a 28-day UCS higher than that of a sample having four per cent cement and trial PSD. The comparison clearly shows that an optimum PSD of CAF can enhance the properties of the fill, and also has a potential for reducing cement content, therefore reducing the cost of fill materials and the overall fill operation.
Tailings addition and CAF properties The effect of sand addition to cemented rockfill was substantially studied in Kidd Creek’s backfill practice. T R Yu (1983) reported that a five per cent addition of sand to the cemented rockfill would significantly reduce segregation of coarse aggregate and result in a 40 per cent increase in compressive strength of cemented rockfill. When more sand was added, the increase in the surface area of the aggregate, which must be coated with the same amount of cement paste, caused the strength to decrease monotonically. Around many mine sites there is often either no sand available or an insufficient supply for mine backfill purposes. It is therefore necessary to make use of tailings as an alternative, which normally have a much finer particle size than sand. The optimal addition of tailings to the cemented aggregate fill remains unknown and needs to be investigated and specified. A laboratory study for this particular objective was therefore carried out at WASM. The effect on fill strength due to the introduction of tailings into cemented rockfill is shown in Figure 2. The results indicate that for fill samples, 100 mm in diameter and 200 mm in length, made of aggregates having a nominal maximum particle size 20 mm (at a cement dosage of four per cent), the highest strength was gained when a tailings addition of ten per cent was adopted. With the same cement dosage, for both cases where cemented rockfill has no tailings or has more than ten per cent tailings, a lower strength was achieved. This is because mixing tailings into a cemented rockfill with no tailings had a poor cement coating effect, whereas a high tailings addition increases the surface area of the particles which require cement coating and binding.
FIG 2 - Influence of tailings addition on the strength of cemented rockfill (Sample size 100 mm diameter × 200 mm length).
Application of cement admixture for improvement of CAF properties The effect of cement admixtures on mechanical properties of cemented fill has been investigated. An admixture dosage of 0.3 per cent, 0.4 per cent and 0.5 per cent was used in cemented aggregate fill without tailings and an admixture dosage range of from 0.3 per cent to 0.9 per cent was used in cemented aggregate/tailings fill. The results are presented in Figures 3 and 4. The general trend from the experimental results is that a 0.3 ∼ 0.4 per cent admixture dosage is capable of increasing CAF strength by around 12 per cent. The results also show that a dosage of 0.5 per cent would indicate an adverse impact on the development of cemented rockfill strength. Figure 4 presents triaxial testing results of cemented tailings/aggregate fill with a recipe of tailings:aggregates:cement = 32:64:4. Admixture dosages used were 0.3 per cent, 0.6 per cent and 0.9 per cent by weight of cement. An increase in the failure stress of around 35 per cent was achieved when a 0.3 per cent and 0.6 per cent admixture was used for both cases of 100 kPa and 300 kPa confining pressures. A 0.9 per cent admixture dosage increased the failure stress by 55 per cent for both tests (100 kPa and 300 kPa confining pressures). The slump
TABLE 3 CAF compressive strength (dolerite based). PSD
Cement (%)
3 days
7 days
28 days
Void Ratio
Trial
4
3.0
3.1
3.7
0.337
Adjusted
3
2.7
3.1
4.3
0.236
Adjusted
4
3.7
4.5
6.37
0.220
MassMin 2000
Compresive strength
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C WANG and E VILLAESCUSA
Total Dissolved Solids (TDS) of 18000 mg/L. Details of the main chemical elements and their contents in the mine water are presented in Table 4.
TABLE 4 Composition of saline mine water. Items Total Dissolved Solids (calc)
FIG 3 - Influence of admixture on cemented rockfill.
0.20
Sodium, Na
5800
Potassium, K
110
Calcium, Ca
630
Magnesium, Mg
430
Chloride, Cl
8500
Carbonate, CO3
100
9900 LEVEL
centimetre s
FIG 12 - Convergence monitoring data from 9900 m level to April 1999. Refer to the CD Rom for colour explanation.
Closure monitoring in planning
CONCLUSIONS
The modelling results confirmed that vertical pillar stress on levels from 9900 m Level and deeper was greatly reduced compared to 9920 m Level due to the levels being overmined (de-stressed). Stress magnitudes for at least the following few levels (9870 m and 9860 m Levels) were determined by modelling to be similar to 9900 m Level. Closure monitoring results for 9900 m Level could therefore be used for predicting closure in these levels, in general and around major shear zones. The calculated closure results for 9870 m and 9860 m Levels were used to determine the time available from development up until 100 cm of closure occurs, at which stage expensive rehabilitation would be required to widen the cross-cuts for LHD access and safety clearances. The available working life calculated for cross-cuts varied from 16 to 24 months, these figures being used in conjunction with other cave draw limitations for optimising tonnage draw limits. MAP3D modelling of deeper levels including 9760 m and 9400 m Levels using the current 17.5 m cross-cut spacing indicates that stress magnitudes will rise to those experienced on 9920 m Level. The rock mass generally improves with depth and the expected closure are expected to be manageable. Localised peaks in stress along the hangingwall shear will, however, result in problem sections and possible increases in cross-cut spacing and/or expected working life. Rock mass rating projections for deeper levels have been used to estimate cross-cut closure and hence excavation life. These limits on excavation life have been used as restrictions in the cave drawdown strategy.
MassMin 2000
That the mine remained operational, and even increased throughput, during such a difficult period in its history in the mid-1990s is testament to the skill of all those involved, especially the operations staff who dealt with severe mining difficulties on a day-to-day basis.
ACKNOWLEDGEMENTS The authors gratefully acknowledge the permission of WMC Resources Ltd to publish the paper, and the support and contributions of colleagues in both WMC Resources, AMC, MINCAD Systems, and Terry Wiles of Mine Modelling Ltd.
REFERENCES Barton, N and Grimstad, E, 1994. The Q-System following twenty years of application in NMT support selection, in Felsbau, 12(6):428-436. Bieniawski, Z T, 1974. Geomechanics classification of rock masses and its application in tunnelling, in Advances in rock mechanics, 2(A)27-32, Washington DC National Academy Science. Grasso, P, Rossler, K, Muccan, S and Xu, S, 1996. The construction, ground reinforcement and monitoring of a large cavern in poor rockmass in NW Italy, in Eurock’96, (A A Balkema: Rotterdam). Obert, L and Stephanson, D E, 1965. Stress conditions under which core discing occurs, SME Transactions, September. Paillet, F and Kim, K, 1987. Character and distribution of borehole breakouts and their relationship to in-situ stresses in deep Columbian River basalts, J Geophys Research, 92, B7, June. Stacey, T R and Harte, N D, 1989. Deep level raise boring - prediction of rock problems, in Proceedings International Symposium Rock at Great Depth, (A A Balkema: Rotterdam).
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Struthers, M and Keogh, J, 1996. Performance of mesh and fibre-reinforced shotcrete under high ground pressures, in Shotcrete Techniques, Procedures and Mining Applications, Kalgoorlie, 1996. Wiles, T, 1997. ‘MAP3D User Manual’, Mine Modelling Ltd, Victoria, Australia.
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Wood, P, Jenkins, P and Jones, I, 2000. Sub-Level Cave Drop Down Strategy at Perseverance Mine, Leinster Nickel Operations, in Proceedings MassMin 2000, pp 517-526 (The Australasian Institute of Mining and Metallurgy: Melbourne).
Brisbane, Qld, 29 October - 2 November 2000
MassMin 2000
Stope Design Based on Realistic Joint Networks M Grenon1 and J Hadjigeorgiou1 ABSTRACT
STOPES IN JOINTED ROCK
The stability of open stopes in underground hard rock mines is influenced by several factors including the shape and size of the excavation, surrounding stresses as well as the degree of jointing. This paper presents a design methodology for stope analysis in jointed rock masses. The proposed approach relies on geological structural data to construct realistic three-dimensional joint network models. The stability of an open stope is determined based on a limit equilibrium analysis.
A rock mass is defined by intersecting discontinuities that can lead to the creation of distinct rock blocks at the periphery of a stope. The presence of such wedges is a main source of structural instability in the excavation. Consequently any design effort should aim to predict the risk and cost of any structural failure, as well as identify all measures to provide adequate reinforcement. Traditional analyses rely on rock classification systems to provide a unique index for the investigated surface (stope back, hanging wall). The limitations of traditional classification systems to fully capture the structural regime of a rock mass have been demonstrated, Hadjigeorgiou et al (1998). The use of traditional limit equilibrium analyses as discussed in Hoek and Brown (1980) is also limited. Such approaches rely on three intersecting discontinuities of infinite length to define a wedge. They do not provide any mechanism to identify the critical discontinuities nor the number and location of possible wedges along a stope surface. Recently Esterhuizen and Ackermann (1998) investigated the stability of a stope in jointed rock. Monte-Carlo simulations were used to generate a series of typical blocks, which were randomly positioned over a stope surface. A series of checks were conducted to determine the efficiency of a given reinforcement pattern. The first check determined whether a bolt could support a block. The second test verified if the block was within the support limit of two adjacent bolts. Several thousand simulations allowed for the development of probability predictions on the stability of the excavation. This approach is promising and arguably an improvement over previous efforts. It is however limited by the assumed spatial distribution of the blocks as well as the assumption that joints are of infinite length. Furthermore this approach does not take into consideration the presence of random joints.
INTRODUCTION Several factors control the stability of open stopes including the shape, size of an excavation, surrounding stresses as well as the degree of in situ fragmentation. While several tools are currently available the most successful design techniques are empirical methods such as the Stability Graph method, Potvin (1988). The success and popularity of the Stability Graph method can be traced to its use as a reliable design tool during the pre-feasibility and feasibility stages. As more geomechanical data become available it is possible to refine any stope design and better evaluate the need and efficiency of support systems. This paper presents a design methodology for open stopes in jointed rock. This is an extension of previous work by Hadjigeorgiou et al (1995) on rock mass characterisation. The method relies on collecting structural field data by traditional scanline mapping which are used to develop realistic three-dimensional joint network models. The generated rock mass is linked to a structural analysis in-house software package that permits the identification of wedges defined by the joint networks and stope dimensions. This allows the evaluation of the stability of any particular stope. The advantages of the proposed approach are that it can account for finite length as well as random joints. Consequently, this results in more realistic wedge representations. A further advantage of this approach is that it lends itself to incorporating risk as an integral part of the design process, allowing for the optimisation of stope design and support. The methodology is demonstrated by referring to a documented case study in a Canadian underground hard rock mine.
EMPIRICAL STOPE DESIGN The most popular exploitation method currently used in the Canadian mining industry is open stope mining. The Stability Graph is the primary tool to design the size, shape and reinforcement of a stope. The stability of an excavation is a function of a stability index N’ and the hydraulic radius of the studied excavation. The index is obtained from the rock mass classification scheme proposed by Barton et al (1974), as well as the geometry of the stope and geotechnical conditions of the rock mass. The hydraulic radius is defined as the ratio of the area of the stope over its perimeter. The stability graph method is also routinely used to provide guidelines for the design of cable reinforcement.
1.
Department of Mining, Metallurgical & Materials Engineering, Université Laval, Quebec City G1K 7P4, Canada.
MassMin 2000
REALISTIC DISCONTINUITY NETWORKS The generation of a three-dimensional discontinuity network based of field measurements is believed to be the most adequate mean of quantifying the discontinuous nature of a rock mass. Over the last ten years the Stereoblock generator was developed at Université Laval to provide a working tool to generate and validate the realistic 3D joint networks. The model is based on the original work of Baecher et al (1977) and has drawn from the work of Villaescusa (1991), Lessard (1996) and Grenon (1997) and others. The developed methodology is summarised in Figure 1. The first step is to conduct a sampling survey of the discontinuity network. This can be achieved through scanline mapping or oriented core logging, accounting for orientation, position and trace length. A statistical analysis is then performed to quantify, for every set, the distributive nature of orientation, spacing and trace length. In the case study presented in this paper the authors employed the results of scanline mapping. The results of this analysis allow the generation of a three-dimensional model of the discontinuity network. This is possible by converting the mean trace length and standard deviation into the diameter of circular joints. The methodology for this conversion has been based on Waburton (1980), Chan (1986) and Villaescusa (1990). The generation of a 3D network relies on using virtual scanlines in the model. Joints are generated along these scanlines until the
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M GRENON and J HADJIGEORGIOU
FIG 1 - Methodology used to generate and validate a realistic discontinuity network.
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STOPE DESIGN BASED ON REALISTIC JOINT NETWORKS
spatial frequency of the generated joints equals the recorded joint frequency recorded during in situ mapping. More details are provided in Grenon and Hadjigeorgiou (2000). An important element of the developed procedure is validation. This is possible by conducting a scanline survey through the simulated network. It is then possible to analyse and compare the statistical properties of the simulated joint network to the one recorded in the field. Meaningful statistical correlation supports the validity of the simulation. The whole process is integrated into Stereoblock, a software package developed at Université Laval.
STRUCTURAL ANALYSIS OF A REALISTIC DISCONTINUITY NETWORK Once a joint network is developed it is possible to introduce any form, size and shape of an excavation and investigate its stability. In the case of open stopes it is possible to quantify all blocks formed at the periphery of the excavation. Figure 2 shows a section along the investigated stope wall. The geometrical and mechanical properties are known for all intersected discontinuities. Thus it is easy to identify all tetrahedral blocks formed along the investigated section. It is the author’s experience that tetrahedral blocks are the most common type observed in the field. This is independently supported by the work of Windsor (1999). If during the geological mapping a major discrete feature such as fault was observed it is possible to introduce this feature in the model. For every generated rock block, the base area, apex length and volume are defined. Apex length is the perpendicular distance between the base of the block and its altitude. For any stope surface it is possible to determine the resulting distributions for every block property. The stability of every individual block is evaluated through a limit equilibrium approach as described in Hoek and Brown (1980). This facilitates the determination of an individual safety factor for every block and the associated risk for any given stope size and orientation. This information provides the framework for the choice and design of a reinforcement scheme. The efficiency of any reinforcement pattern is based on a methodology developed by Li (1991). At the design stage this approach can be used to compare different reinforcement strategies.
CASE STUDY The developed methodology is best illustrated by a case study at a mine in Northwestern Quebec. Three scanlines were run perpendicular to each other. All encountered discontinuities were recorded for orientation, position and trace length. The collected structural data were entered into a commercially available stereonet software package. Following a visual identification of the dominant discontinuity sets, Figure 3, their mean orientation and coefficient of dispersion (K) were determined. For every discontinuity set, the mean normal spacing value and the mean trace length were also evaluated. The results are summarised in Table 1. Given that random joints represent 31 per cent of all recorded joints it was important to consider them. In order to account for the possible influence of random joints, the total spacing and the mean trace length of these discontinuities were evaluated. While the model has capabilities for accounting for joint censoring in the presented case study this was not undertaken as 96 per cent of the joints had a length inferior to 3.5 m and were not censored. The rock mass is characterised by three dominant discontinuity sets. The first joint set has a higher dispersion than the other two sets. The average trace length of the sets would be described as low persistence with spacing value defined as moderate to wide. The mean total discontinuity spacing is 0.15 m, resulting in a Rock Quality Designation of 85 per cent, calculated by the Equation 1 proposed by Priest (1993). This is indicative of good quality rock mass. RQD = 100e-0.1λ(1 + 0.1λ)
(1)
Table 2 summarises the distribution of the in situ discontinuity set characteristics. For every discontinuity set parameter a theoretical distribution is tested against the collected data using a χ2 test. Ten out of the 11 distributions tested were accepted. Based on these results, the Stereoblock model was deemed applicable. A 3D joint network was generated using the more recent version of Stereoblock. The simulated rock volume was 35 x 35 x 60 m. The stope has a width of 15 m, a height of 30 m and a length of 50 m. The recorded joint set orientation, spacing and
FIG 2 - Section cut through the simulated joint network.
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FIG 3 - Stereonet of the collected data identifying the two major joint sets.
TABLE 1 Statistical analysis of mapped discontinuity network. o
Orientation ( )
K
Average trace length (m)
Trace length standard dev (m)
Normal spacing (m)
Set 1
61/037
13
1.40
1.30
0.23
Set 2
49/301
69
1.00
1.10
0.71
Set 3
90/136
43
0.83
0.91
1.00
Random
—
—
1.00
0.90
0.91
TABLE 2 Validation of in situ discontinuity network. Distribution
Number of classes
χ2 test
Accepted
Orientation
Fisher univariate
7
11.02
Yes
Spacing
Exponential negative
7
0.82
Yes
Trace
Lognormal
8
16.50
No
Orientation
Fisher univariate
5
1.18
Yes
Spacing
Exponential negative
6
7.21
Yes
Trace
Lognormal
7
11.84
Yes
Orientation
Fisher univariate
5
6.55
Yes
Spacing
Exponential negative
5
2.45
Yes
Trace
Lognormal
8
7.22
Yes
Spacing
Exponential negative
6
4.60
Yes
Trace
Lognormal
6
10.59
Yes
Parameter Set 1
Set 2
Set 3
Random
768
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MassMin 2000
STOPE DESIGN BASED ON REALISTIC JOINT NETWORKS
trace length distributions are used to generate a 3D joint network for a given rock volume. Following validation exercises the generated was accepted. Table 3 demonstrates how the generated distributions fitted the distributions observed for the in situ data. Ten out of the 11 distributions tested were equivalent and the results were deemed satisfactory. Once the discontinuity network was generated, following the methodology described in section 5.0, the generated blocks were characterised based on their geometric properties at the stope back. The geometric properties of the blocks are presented in Figures 4 to 7. A closer inspection reveals that the majority of blocks are relatively small in size. In fact, very few blocks have an apex length exceeding 3 m and no block has a volume greater than 0.5 m3. It should be noted that the undertaken analysis has also accounted for the presence of random discontinuities. In previous works these were often ignored to facilitate analysis. In Figure 7 it is demonstrated that in the present case more then 22 per cent of the created blocks are formed with at least one random discontinuity (joint set 0). Figure 8 shows the distribution of safety factors for each block. For the purposes of this preliminary analyses a cohesion of 0 was used and an angle of friction of 300. Recent modifications to the structural analysis software allow the use of any series of distributions for the frictional properties of the discontinuities.
500
Number of Blocks
400 300
200
100
0
0
1
2 3 Apex Length (m)
4
5
FIG 4 - Number of blocks versus apex length.
600
Number of Blocks
500
Influence of reinforcement The developed structural analysis package has the facility to investigate the influence of reinforcement. The mine uses a standard reinforcement pattern for stope backs of:
400 300 200 100
• 10 m cables on a 2 x 2 m pattern; • 2.1 grouted rebars on a 1.2 x 1.2 pattern; and • Gage 9 welded mesh.
0
The first attempt to evaluate the efficiency of the reinforcement system relied on determining the resulting factor of safety for every block. These were plotted in Figure 9. It is of interest to note that there are two extremes in the histogram. In the first place there are a number of blocks now stable with reinforcement enjoying a high factor of safety. These blocks were identified as the larger blocks generated in the rock mass with their length and area distributions plotted in Figures 10 and 11. It is hence shown that the support system of cables and rockbolts is more than
0
0.2
0.4 0.6 Area (m2)
0.8
1
FIG 5 - Number of blocks versus base area.
adequate in supporting the larger blocks on the stope back. The number of unstable blocks is at first glance alarming (79 per cent of the generated blocks). A closer inspection reveals that these blocks however are very small in size thus posing no threat. It
TABLE 3 Validation of the simulated discontinuity networks.
Set 1
Set 2
Set 3
Random
MassMin 2000
Parameter
Distribution
Number of classes
χ2 test
Accepted
Orientation
Fisher univariate
5
9.50
Yes
Spacing
Exponential negative
6
1.46
Yes
Trace
Lognormal
7
6.87
Yes
Orientation
Fisher univariate
6
15.76
No
Spacing
Exponential negative
5
9.24
Yes
Trace
Lognormal
6
3.69
Yes
Orientation
Fisher univariate
6
6.57
Yes
Spacing
Exponential negative
5
1.87
Yes
Trace
Lognormal
6
1.99
Yes
Spacing
Exponential negative
5
1.74
Yes
Trace
Lognormal
5
3.07
Yes
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M GRENON and J HADJIGEORGIOU
700
500
500
400
Numbber of Blocks
Number of Blocks
600
400 300 200 100
300
200
100 0
0
0.5
1
1.5
Volume (m3)
0 -1
0
1
2
FIG 6 - Number of blocks versus volume.
3
4 5 6 Safety Factor
7
8
9
10 11
FIG 9 - Safety factor quantification. 1200
40
800
35 30
600
Number of Blocks
Number of Discontinuities
1000
400 200 0 -1
25 20 15 10
0
1
2
3 4 5 6 7 Discontinuity sets
8
9
10 11
5 0
0
1
FIG 7 - Number of discontinuities per set.
2 3 Apex (m)
4
5
FIG 10 - Number of reinfoced rock blocks versus apex length.
500
60 50 300
Number of Blocks
Number of Blocks
400
200
100
40 30 20 10
0 -1
0
1
2
3
4 5 6 Safety Factor
7
8
9
10 11
0
0
FIG 8 - Safety factor distribution.
0.2
0.4 0.6 2 Area (m )
0.8
1
FIG 11 - Number of reinforced rock blocks versus area.
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MassMin 2000
should also be noted that the mine also employs a gage 9 screen. Based on the work of Tannant (1995), the support capacity of the mesh was introduced in the stability analysis. It was then shown that less then one per cent of the unstable blocks will exceed the strength of weld mesh. These results indicate that the combination of cables, rebar and weld mesh is sufficient to control the stability of the back of an open stope. Finally, it was decided to determine whether the reinforcement pattern and bolt length were adequate. The standard mine reinforcement pattern was simulated and its performance evaluated. Based on the block apex length dimensions it seems clear that the cable length are sufficient to reinforce 100 per cent of the blocks while the rebar length is sufficient to stabilise 98 per cent of them. Figure 12 shows the number of bolts that intercept a rock block. The results indicate that, out of the 746 reinforcement units installed, 90 intercepted a block. This represents only 12 per cent of efficiency. Of the 90 intersecting bolts, 51 are rebars and 39 cables.
600
Number of Blocks
500 400 300 200 100 0 -1
0
1
2
3 4 5 6 7 Number of Bolts
8
9
10 11
FIG 12 - Number of blocks intercepted by a bolt.
Once the stope was blasted and mucked, a cavity monitoring system (CMS) was used to quantify the stability of the back and walls. In this particular case it was shown that the back remained stable. This is well in agreement with the analysis.
CONCLUSIONS Block size is a critical factor in the stability of an excavation. As the span increases, it is even more important to use realistic block sizes. This necessitates the use of realistic joint networks to define block size. Consequently, trace length is an important input parameter. A further observation is the importance of random joints in the analysis. Use of the joint networks results in a large number of blocks. Smaller blocks may not be critical and it can be argued that they will fall during blasting or be scaled down. It was also shown that welded mesh can provide adequate support for the small blocks in the investigated stope. Cables and rock bolts used to stabilise the larger blocks not only provide adequate support but are also of sufficient length to allow the development of an adequate embedment length.
MassMin 2000
The analysis undertaken has only focused on the influence of structure and has not investigated the stress regime, which may have an important role to play. The 3D joint network model lends itself to stress analysis as well. Grenon et al (2000) report on exporting cuts from the model and importing them in a distinct element program.
ACKNOWLEDGEMENTS The authors would like to acknowledge the technical contribution of Jean-François Lessard. The financial support of the National Science and Engineering Council of Canada and the Institut de Recherche en Santé et Sécurité au Travail of Quebec.
REFERENCES Chan, L P, 1986. Application of Block Theory and Simulation Techniques to Optimum Design of Rock Excavation, PhD thesis, University of California Berkley. Baecher, G B, Lanney, N A and Einstein, H H, 1977. Statistical Description of Rock Properties and Sampling, in Proceedings 18th U S Symp on Rock Mechanics, Colorado, pp 5C1.1-5C1.8. Barton, N R, Lien, R and Lunde, 1974. Engineering classification of rock masses for the design of tunnel support, Rock Mech, 6:189-239. Esterhuizen, G S and Ackermann, K A, 1998. Stochastic keyblock analysis for stope support design, Int J of Rock Mech and Min Sci, 35:4-5, Paper no 122. Grenon, M, 1997. Utilisation de la blocométrie du massif rocheux dans les mines souterraines. Mémoire de Maitrise, Université Laval. Grenon, M, Hadjigeorgiou, J, Harrison, J P and Banas, R, 2000. An Investigation of Drift Stability in Moderately Jointed Rock Mass Proceedings of the Fourth North American Rock Mechanics Symposium, Seattle, Washington, USA. Grenon, M and Hadjigeorgiou, J, 2000. A 3D joint network model (in preparation). Hadjigeorgiou, J, Lessard, J F and Flament, F, 1995. Characterizing in-situ block distribution using a stereological model, Canadian Tunnelling Journal, pp 201-211. Hadjigeorgiou, J, Grenon, M and Lessard, J F, 1998. Defining in-situ block size, CIM Bulletin, 91(1020):72-75. Hoek, E and Brown, E T, 1980. Underground Excavations in Rock, p 527 (The Institution of Mining and Metallurgy: London). Li, B, 1991. The stability of wedges formed by three intersecting discontinuities in the rock surrounding underground excavations, PhD thesis, University of Toronto. Lessard, J F, 1996. Evaluation de la blocométrie des massifs rocheux, Mémoire de Maitrise, Université Laval. Potvin, Y, 1988. Empirical open stope design in Canada, PhD thesis. The University of British Columbia p 350. Priest, S D, 1993. Discontinuity Analysis for Rock Engineering (Chapman and Hall), p 473. Tannant, D, 1995. Load capacity and stiffness of welded-wire mesh, 48th Canadian Geotechnical Conference Vancouver Canada. Villaescusa, E, 1991. A three dimensional model of rock jointing, PhD Thesis, University of Queensland. Warburton, P M, 1980. A Stereographical Interpretation of Joint Trace Data, Int J Rock Mech Sci & Geomech Abstr, 17:181-190. Windsor, C R, 1999. Keynote lecture : Systematic design of reinforcement and support schemes for excavations in jointed rock, in Proceedings of the international symposium on ground support, Kalgoorlie, Australia, p35-58, (Balkema).
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Seismicity Control of Induced Seismicity at El Teniente Mine, Codelco-Chile
E Rojas, P Cavieres, R Dunlop and S Gaete
775
Support Appropriate for Dynamic Loading and Large Static Loading in Block Cave Mining Openings
T R Stacey and W D Ortlepp
783
Seismicity at Big Bell Mine
M Turner and J Player
791
Study of the Geodynamic Regime of the Region Under Large-Scale Mining in the Apatite-Nepheline Deposits in the Kola Peninsula, Russia
A A Kozyrev, V I Panin and V A Maltsev
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Experience in Block-Pillar Mining Under Rock Burst Conditions
A A Kozyrev, V A Maltsev, V V Rybin and V S Svinin
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Control of Induced Seismicity at El Teniente Mine, Codelco-Chile E Rojas1, P Cavieres1, R Dunlop2 and S Gaete3 ABSTRACT
STARTING THE PRIMARY ORE EXPLOITATION
El Teniente Mine started the exploitation of the primary rock in the beginning of the 1980s decade, using caving methods. Since the beginning the mining activities have induced seismicity including some occasional rockbursts. Owing to the relevant damages, the mining operations were stopped in some production sectors. This paper describes the concepts that have been used in order to control the induced seismicity based on the strategies for controlling the rockmass breaking process, which depend on mining parameters. It also presents the main concepts applied in the Teniente Sub6 sector design and the experience collected during its sector exploitation.
After preliminary tests during the 1970s and the 1980s decades the primary rock exploitation commenced and the induced seismicity started. By 1981, the long-term mine plan of the El Teniente considered two sectors for the expansion of the mine operations using a mechanized panel caving method, with columns height around 140 m. The primary ore exploitation proved to be more difficult and more expensive, and on the other hand, the grades in primary ore grades are lower than in the secondary ore. These factors were pulling down the revenues. At that moment, the development costs were considered as depending only on the area and a natural way to lower the costs was to increase the column height. The geomechanical condition and the behavior of the rockmass exposed to the caving were defined to be independent of the column height according to the results available at that time. Two parameters considered column height dependent were the dilution and the ore recovery. Using these concepts, the decision was made to apply a mechanised panel caving method with a higher rockmass column. Accordingly, a new sector, Teniente 4 South, was developed with a total column of 180 m. The main part of this column was classified as secondary ore. As the exploitation progressed to the South, an increasing portion of the column was primary ore. By 1990, the column height of the production area reached a 280 m value and it was a 100 per cent primary rockmass volume.
INTRODUCTION The El Teniente Mine is a Codelco-Chile underground copper mine. It is located in the first elevations of the Andes in the central zone of Chile, about 70 km SSE from the capital city, Santiago. The El Teniente porphyry copper orebody is one of the largest known copper deposit in the world. It includes andesite, diorite and hydrothermal breccias of the Miocene era as the main lithologies. The main structural feature of the orebody is a stockwork of multi-directional veins and veinlets. The veins are principally cemented with anhydrite, quartz and sulphides. A chimney of subvolcanic breccias known as the ‘Braden Pipe’ postdates the copper-molybdenum mineralisation. It has an inverted cone shape and the hydrothermal mineralisation is distributed around this pipe over a variable radial extension of 400 m to 800 m, with mineralogical associations of variable strength. The mineralisation has two very different forms, the secondary ore is located near the surface and the primary mineralisation is at greater depth. The primary ore can be described as a high cohesion and impermeable rockmass. The stockwork veins, containing the original mineralogy, are sealed. According to a geomechanical behavior, the primary rockmass has been characterised as harder than the secondary. Its caving results in a large fragmentation. Primary rockmass could exhibit brittle, often violent failure under a high stress condition.
EXPLOITATION OVERVIEW Since 1906 more than 1100 Mt have been mined out. The current daily production is close to 96 000 tpd with an average grade of 1.1 per cent. Since 1945, the exploitation methods have been the block caving and later panel caving. The initial production extracted the secondary ore close to the mountain surface. The first test searching for a mining method to be applied to the primary ore extraction was conducted by the middle of the 1970s. The first sector exploiting a mainly primary ore started its production in 1982. 1.
Mining Engineer, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
2.
Geophysicist, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
3.
Matemathical Engineer, Codelco-Chile, El Teniente Division, Av Millan 1020, Rancagua, Chile.
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TOWARDS A PRIMARY ORE MINE, THE SUB6 SECTOR The long-term plan required the deepening of the exploitation as a way to lower the costs and to expand the scale operation. The north sector of the orebody was defined to be exploited. The Sub6 sector was the first project including the deepest productive area with a higher column of a 100 per cent primary rockmass. The long-term mine planning defined the Sub6 production to be the 50 per cent of the total mine production by the year 2000. The design concepts available at the Sub6 engineering stage (1984) recommended the application of a mechanised panel caving method using LHD equipment for the intermediate transport in the production level. Production drifts separated by a 30 m distance defined the production level layout. These drifts were intersected by a set of 15 m spacing parallel galleries (trenches) at a 60º angle, forming the El Teniente type drawpoint mesh. The extraction rate was set as 0.75 t/m2-day as an average. The only extraction rate restriction was the transport capacity of the corresponding level located under the production level. An indication was stated in order to control the dilution using the extraction rate and draw regularity. A relevant design concept was to reduce the length of galleries per extracted ton of ore. This concept directly introduces the increase in the column height in the panel caving method. The geomechanical main recommendations were:
• the orientation of the galleries according to the principal stresses;
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• the undercutting sequence in agreement with the principal stresses; and
• the influence of the subsidence on the upper level infrastructure. With the Teniente 4 South exploitation induced seismicity had been was evident. No indications were stated that the initial stage of the caving progress in the primary rockmass would be affected by a related relevant rockmass seismic response.
SUB6 SECTOR COMMISSIONING The initial developments for the Sub6 sector began in 1985. The access from surface and the underground access were first developed. The area preparation started in 1986 and production was started by the middle of 1989, using a conventional panel caving. Six months after the beginning of productive activities, the Sub6 sector was affected by induced seismicity with associated major rockbursts. As a global result we can state that the operations were driven by the rockmass behaviour. All the tactical activities were leaded by the rockmass response, and the mining became a reaction to the rockmass response. In 1992, it was evident that the problem was a strategical issue, and the global long-term plan had to be changed as pointed out by Cavieres, Gaete and Karzulovic (1994). The main events are described as follows.
18 January 1990 According to the original plan, the Sub6 operations started on August 1989, to reach a 10 600 m2 undercut area and 3500 tpd production by the middle of January 1990. A 3.6 Richter magnitude scale seismic event and a relevant rockburst generated wide spread damage. One of the more important damages affected the temporarily established transport gallery. The Sub6 operations had to be stopped. No major changes were made to its production plan
2 July 1990 A 3.2 Richter scale seismic event and the associated rockburst created a so much related damage that all the Sub6 mining activities were stopped for a nine-month period. A systematic repairing of all the damages was accomplished during that period. A safety zone 100 m wide was defined in order to separate the preparation activities and the production operations. In that way, the sector personnel were exposed to a reduced rockburst risk. The zone dimension was estimated from some geomechanical analysis of the area and from the Teniente 4 experience where a 100 m zone was observed as affected by the undercutting front.
23 May 1991 A sequence of seismic events affected the area, initiated by a 4.0 Richter scale event. Different damages were generated around the previously reinforced area. The production activities were stopped during a five-month period. After this rockburst, some actions were assumed in order to get some control of the seismic rockmass control. A partial and temporary closing of the area was established after undercut blasting. A relative operational control was defined using a seismic alert criteria based on event frequency.
25 March 1992 A 3.7 Richter scale event severely damaged the main access to the production area.
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The global mining plan was changed, no production requirements were planned for the sector until a new approach to sector the exploitation would be defined. The production was resumed in January 1994, under the application of an experimental production plan.
CONCEPTUAL FRAMEWORK FOR THE INDUCED SEISMICITY In January 1992, a global digital seismic network was installed, monitoring the induced seismic activity in the whole mine. Additionally, the regional network was improved to get better covering of the tectonic activity, due to some hypothesis relating the mine situation with the tectonic generated seismicity. The regional monitoring showed no evidence of relevant events in the mine area and on the other hand, the one year recording revealed the global relation between the mining and seismic location. According to Dunlop and Gaete (1995), during 1992 - 1993 a conceptual framework was developed in order to relate the mining parameters to the rockmass response characteristics. This relation would allow the control of the induced seismicity by modifying the mining parameters. The caving methods are initiated by the blasting of the bottom volume of a rockmass column. The broken material is mined out creating cavities that allows gravity to continue the fracture process of the rockmass, producing new broken material. The subsequent production generates the continuity of the breaking process, propagating the rockmass fractures to the upper levels. In general terms, the fractures correspond to the disruption of the structural pattern of the jointed rockmass. A seismic event corresponds to the radiated energy associated with a rockmass rupture. Then, the induced seismicity is always associated to a rupture process affecting a competent rockmass. The characteristics of the induced seismic events will be determined by the spatial and temporal distribution of the mining activities and conditioned by the geometrical, geological and structural characteristics of the mined rockmass. The caving methods do not always allow an adequate control of rupture extension. According to the mining and the rockmass parameters, caving could generate large ruptures, ie high magnitude seismic events that can radiate enough energy to produce damage to the surrounding excavations. The actions for controlling the seismic rockmass response, particularly the maximum event magnitudes, are based on the modification of the mining parameters, such as the undercutting advancing rate and the extraction rate, in relation to the rockmass characteristics. This framework and some empirical evidences indicate that the initial caving situation (ie before the caving cavity connects to intact rockmass upper surface) corresponds to the period of greater seismic risk when using caving as the mining method. Once such a connection has been achieved, the changes in the mechanical stability conditions of the rockmass would create a more favorable seismic response. These concepts explain the seismic rockmass response in general terms allowing a global control of the induced seismicity as shown in Dunlop and Gaete (1997). However, local interactions between the rockmass geometrical or structural singularities and short-term mine operations could be the source of related relevant seismicity. Currently, different strategies are under development in order to control this local scale seismicity.
TENIENTE SUB 6 EXPERIMENTAL PLAN Based on the previous concepts, an experimental production plan was implemented in the Sub6 sector starting on January 1994. The plan aimed to gradually expand the rockmass breaking process in order to reach an old upper undercut level, Teniente 4, already exploited, covering 240 m of column.
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To develop this plan with a controlled seismicity, the Sub6 production was modified as follows:
b.
c.
The productive area was limited to an area already with undercut level blasting and with a shorter column of an intact rockmass (only 240 m), temporarily closing the open area with a higher column.
Period
The spatial and temporal distribution of the extraction was maintained as uniform as possible in order to minimise the rupture sizes and, consequently, the seismic event magnitudes. At the beginning of the plan, very low production rates were accomplished and they were gradually increased. Starting production rates were close to 0.1 t/m2-day. An extraction rate range was established for the sector, which has been slightly modified during the mining of the area.
d.
The extraction was made using only remote controlled equipment until 1995.
e.
The extraction rate was indirectly limited by criteria based on a time weighted seismic activity average of the last two weeks.
Production (t)
Undercutting (m2)
Relevant rockbursts
89/07/18 - 90/01/18
384 000
9174
1
90/04/01 - 90/07/02
107 000
480
1
91/04/01 - 91/05/23
175 000
120
2
91/11/01- 92/03/25
160 000
10 410
2
94/01/02 - 97/08/30 (EP)
3 100 000
2160
0*
97/09/01 - 99/12/31
6 500 000
10 790
0**
(*) 2 minor rockbursts (**) 1 minor rockburst
Figure 1 shows the Sub6 production since 1989 and the number of rockbursts affecting the area. The Sub6 production has not been significantly affected by rockbursts since 1994.
According to estimations and some empirical evidences, the complete column would be fractured after an extraction of 30 per cent of the solid column. In that condition, no further ruptures would affect the rockmass column, then no seismic activity response would be generated. Table 1 presents the results of the initial mining periods and the total production during the experimental plan (EP).
CURRENT STATE After the end of the experimental plan, the Sub6 sector was subject to a transition stage, including an undercutting test. It was integrated to the long-term production plan during 1999. In the 1997 - 2000 period few minor rockbursts have occurred with slight damages.
14000
7
1
2
3
4 6
10000
5
8000
4
6000
3
4000
2
2000
1
0
0
7
19
99
-1 1
3
-0
-0
99
99
19
7
-1 1
98
19
19
3
-0
-0
98
98
19
7
-1 1
97
19
19
3
-0
-0
97
97
19
7
-1 1
96
19
19
3
-0
-0
19
96
-1 1
96
95
19
19
3
7 -0
-0
95
95
19
7
-1 1
94
19
19
3
-0
-0
94
94
19
7
-1 1
93
19
19
3
-0
-0
93
93
19
7
-1 1
92
19
3
-0
-0
92
19
92
19
7
-1 1
91
19
19
3
-0
-0
91
91
19
7
-1 1
90
19
19
3
-0
-0
90
90
19
19
7 -0
89
19
89 19
-1 1
tpd
12000
# Rockbursts
a.
TABLE 1 Sub6 production and relevant related rockbursts
Date tpd
1
INITIAL OPERATION
2
EXPERIMENTAL MINING
3
Rockbursts
PRE-OPERATIONAL STAGE
4
NORMAL OPERATION
FIG 1 - The Sub6 mining periods.
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Table 2 presents the current extraction rates for the Sub6 area. At present, the sector production is close to 10 000 tpd, and the undercutting rate is close to 12 000 m2 per year, the exact number depend on the geometry and the rockmass. TABLE 2 Maximum production rates (tpd/m2). Extraction %
Initial caving stage
0-5
0.26
Steady caving stage 0.28
5 - 10
0.29
0.34
10 - 15
0.33
0.40
15 - 20
0.38
0.47
20 - 25
0.43
0.55
25 - 30
0.50
0.65
ANALYSIS OF THE RESULTS The results of the mining activities at the Sub6 sector have shown that the most critical stage developing a caving operation in a new sector is the period before the connection to an upper crater or an abandoned upper sector. This condition is created by the mechanical response of the rockmass under an initial caving process. A mechanical response involving the complete rockmass under the influence of the mining activity and the rockmass searching for a stable equilibrium condition as an arch appear to be the main factors generating this relevant seismic response. It is important to consider different alternatives to reach the connection in order to reduce as much as possible the initial caving stage.
As a back-analysis, the Figures 2, 3, 4 and 5 present the rockmass conditions existing at the moment of the four main rockbursts. The differences between the estimated geometry of the rockmass corresponding to the applied production rates (1989 - 1992 period) and the recommended rates for the experimental plan can be appreciated. Figure 6 presents the estimated situation during the sector connection to the upper exploited level by the middle of 1996. Some relevant seismic parameters are the radiated energy and the influence distance. According to empirical evidence, damage is associated with events having a radiated energy higher than 107 [J]. A 100 m distance has been established as the minimum distance in order not to get damage from a high magnitude event according to the moment magnitude scale used as a local scale. The sector has suffered high magnitude events but located more than 200 m away from Sub6 excavations without getting damage. In a caving exploitation, a critical point is to reach a steady caving condition. Beyond this point, the mining business starts. The critical issues from this point of view are the mining parameters during the initial caving stage and some permissible mining rates according to the characteristics of the exploited rockmass and its response to the mining. A basic idea is to limit the rockmass volume in an unstable condition. Mining has to maintain the control of the equilibrium rockmass conditions. It means a controlled unstable rockmass condition. Regarding the engineering process of the Sub6 project, the applied standards were in agreement with those used by the world mining industry. In spite of this situation, the seismic response to the initial caving stage was not anticipated. As a consequence, the El Teniente Division was forced to assign a large amount of resources, including external consultancies, in order to generate a mining methodology to extract the primary ore using a caving method.
N
Simulated Extracted Ore Using an Initial Controlled Caving Rate
A-A Section View (Looking at the North) Extracted Ore
A
A
Production Level
FIG 2 - Sub6 sector on 18 January 1990.
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N
A-A Section View (Looking at the North)
Simulated Extracted Ore Using an Initial Controlled Caving Rate
Extracted Ore
A
A
Production Level
FIG 3 - Sub6 sector on 2 July 1990.
N
A-A Section View (Looking at the North) Simulated Extracted Ore Using a Initial Controlled Caving Rate
Extracted Ore
A
A
Production Level
FIG 4 - Sub6 sector on 23 May 1991.
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N
Simutaled Extracted Ore Using an Initial Controlled Caving Rate
A-A Section View (Looking at the North) Extracted Ore A
A
Production Level FIG 5 - Sub6 sector on 25 March 1992.
N
Ten-4 Sector Abandoned Area Fragmented Ore
A-A Section View (Looking at the North)
Extracted Ore
A
A
Production Level
Simulated Extracted Ore Using an Initial Controlled Caving Rate FIG 6 - Sub6 sector on 30 June 1996.
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ACKNOWLEDGEMENTS The authors wish to acknowledge the authorisation given by Codelco-Chile, El Teniente Division to publish this paper.
REFERENCES Cavieres, P, Gaete, S and Karzulovic, A, 1994. Geotechnicalgeomechanical evaluation of the Teniente Sub6 sector and its impact in the medium and long term. Codelco-Chile, El Teniente Division, Internal Report # EM-10/94.
MassMin 2000
Dunlop, R and Gaete, S, 1995. Seismicity at El Teniente Mine, in Proceedings 4th International Symposium on Mine Planning and Equipment Selection. (Eds: Singhal et al), Calgary, Canada, pp 865-869 (Balkema: Rotterdam). Dunlop, R and Gaete, S, 1997. Controlling induced seismicity at El Teniente Mine: the Sub6 case history, in Proceedings 4th International Symposium on Rockbursts and Seismicity in Mines, (Eds: S J Gibowicz and S Lasoki), Krakow, Poland, pp 233-236 (Balkema: Rotterdam).
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Support Appropriate for Dynamic Loading and Large Static Loading in Block Cave Mining Openings T R Stacey1 and W D Ortlepp2 ABSTRACT Seismicity associated with deep block cave mining operations in hard rock can give rise to dynamic loading of the mining openings. A description is given of the source and damage mechanisms associated with such seismic events, and the possible demands placed on the rock support by the resulting dynamic loading. In addition, production level drifts are often subjected to large static deformations during mining. Appropriate support, able to sustain large static deformations, and capable of absorbing the energy released during dynamic loading, is described. The performance of these support types has been determined by means of large-scale dynamic loading tests, and large deformation static laboratory tests. Finally, a support system which is considered to be appropriate for extraction drifts in block cave layouts subjected to dynamic loading is described.
INTRODUCTION Cave mining of deep, hard rock orebodies, involving removal of large volumes of rock, will inevitably lead to the generation of mining-induced seismicity, which may lead to rockbursts. A rockburst may be understood to be ‘a seismic event which causes violent and significant damage to tunnels and other excavations in the mine.’ There are no constraints on the magnitude of the seismic event. Thus, the event can range from a strainburst, in which superficial surface spalling with violent ejection of fragments occurs, to a mining-induced ‘earthquake’ involving slip along a fault plane. The range in Richter magnitudes for these two limits is from about -0.2 to 5.0. This seismicity can lead to dynamic loading of the rock surrounding mining openings and may also cause rockbursts. The main types of rockburst source mechanism which have been identified (Ortlepp and Stacey, 1994) are summarised in Table 1. It should be noted that some types of seismicity do not necessarily require high stress levels for their occurrence. Stacey (1989) found that, from a review of the occurrence of this type of seismicity, particularly with massive rock conditions, strain bursts can occur in tunnel development when the field stress is as low as 15 per cent of the uniaxial compressive strength of the rock material. More recent information suggests that this figure could be as low as ten per cent for very brittle rocks. The location of the source of the seismicity and the location of the rockburst damage may or may not be coincident. In the larger magnitude events, the separation of the two locations may be hundreds of metres. The factors determining the intensity of the seismic impulse include the following (Ortlepp, 1997):
• • • • •
the amount of energy available;
1.
Director, SRK Consulting, PO Box 55291, Northlands 2116, South Africa. E-mail: [email protected]
2.
Associate Consultant, SRK Consulting, PO Box 55291, Northlands 2116, South Africa.
the rate of liberation of energy; the source distance and dimension; the peak particle motion (shear, compression or other); and the ray path properties, influenced by the geological structure, or major excavations, intervening to cause reflection, shielding, channelling, etc of the seismic waves.
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TABLE 1 Suggested classification of seismic event sources with respect to tunnels. Seismic Event
Postulated Source
First Motion from Seismic Records
Guideline Richter Magnitude ML
Strainbursting
Usually Superficial spalling with undetected; could violent ejection of be implosive fragments
-0.2 to 0
Buckling
Outward expulsion of pre-existing larger slabs parallel to opening
Implosive
0 to 1.5
Face crush
Violent expulsion of rock from tunnel face
Implosive
1.0 to 2.5
Shear rupture
Violent propagation of shear fracture through intact rock mass
Double-couple shear
2.0 to 3.5
Fault-slip
Violent renewed movement on existing fault
Double-couple shear
2.5 to 5.0
Factors which influence the response of the excavations to the seismicity include excavation geometry (size and shape); site amplification factors (stress intensity, stress distribution); characteristics of the surrounding rock (strength, brittleness, fabric, structure, intensity of induced fracturing); characteristics of existing support (length, strength, density, yieldability, quality of installation, quality of containment support). Types of rockburst mechanisms which have been identified include ejection, buckling, gravity enhancement, shake-out, fall of ground associated with large, distant seismic events, disruption and displacement, and convergence and heave.
IMPLICATIONS OF SEISMICITY FOR CAVE MINING LAYOUTS In hard rock, deep cave mining conditions, it can be expected that the following types of seismicity will be experienced: Strain bursting from the walls and faces of tunnels could be expected to occur intermittently on any level during primary development. This is hazardous, since sharp edged fragments, often plate sized, are ejected violently, but does not result in major stability problems. An example of this type of occurrence, which caused the failure of three rockbolts, is shown in Figure 1. During undercutting, strain bursting could be expected to occur with greater frequency in all undercut development, but particularly in the region of the abutments and advancing undercut front. This behaviour should reduce substantially or even stop once the cave has propagated through to surface. Owing to the major changes in the stress distribution in the cave back and surrounding rock mass during the development of the cave, and subsequently in the surrounding rock mass as a result of the creation of the ‘destressed cave void’, stress conditions may be conducive for the generation of buckling, fault slip, and possibly shear rupture types of seismic events. These events could involve large amounts of energy, and their effects
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FIG 1 - Strain burst area, showing fractured rockbolt surfaces.
could be manifested in the production level excavations, and openings such as ventilation, rock breaker, crusher and other service excavations adjacent to or below the production level. The damage associated with such events has involved roof, sidewalls and floor of excavations, in many cases resulting in complete closure of the tunnel. It has typically been observed that approximately a metre thickness of rock from the walls of the excavation is violently ejected. An example of such an event is shown in Figure 2. From back analyses of the ejection velocities of several of these occurrences, it was suggested that an appropriate velocity for support design purposes could be 10 m/s (Ortlepp and Stacey, 1994). The undercut level will be substantially protected from these types of events. There are three potential approaches to the alleviation of problems due to rockbursts. These are:
• prevention of seismicity, and hence rockbursts; • prediction of rockbursts, and timely evacuation of personnel
FIG 2 - Approximately a metre thickness of rock ejected violently.
and equipment; and
• containment of rockburst damage with appropriate support. It is considered that the only practical option of these three is that of containment of damage. Success with prevention and prediction is likely to be only partial, and the implication is that there will always remain some risk of rockbursts which will have to be addressed by means of appropriate rock support. In the following sections, the results of dynamic testing of different types of rock support are described. In addition to potential dynamic loading, deep caving layouts will be subjected to significant static stresses and stress changes, and support will be required to contain large static deformations. The results of some large deformation static testing of reinforced shotcrete will also be given.
PERFORMANCE OF SUPPORT UNDER SEVERE LOADING Several programs of dynamic support testing have been carried out over the past eight years, involving both retainment support such as rockbolts and cables, and containment support, including various types of wire mesh, wire rope lacing, and shotcrete. Retainment support testing, containment support testing and static load testing will be dealt with separately in the sections below.
Dynamic testing of retainment support Using explosives as the impulse force, Ortlepp (1994) carried out tests on rebar and smooth bar rockbolts, and special yielding rockbolts. The results of these tests are summarised in Table 2.
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From these results it can be seen that the rebar rockbolts were not able to contain the energy, and this type of support failed in all cases. In contrast, the yielding cone bolts performed well, and none of the bolts was broken. The 16 mm cone bolts failed as a system under the impulse of a large amount of explosive, but the bolts were undamaged. The smooth bar bolts also performed well. Thus, in spite of the fact that the rebar rockbolts had the greatest strength capacity, they were unable to resist the dynamic loading. Using a drop weight impact system for loading, dynamic testing has been carried out on the following (Stacey and Ortlepp, 1999):
• 16 mm diameter rebar (fully grouted, rigid); • 16 mm diameter smooth bar (partially bonded, semi-yielding);
• destranded hoist rope strand (lubricated, helix); • 18 mm compact strand cable (fully grouted, prestressed, very stiff);
• 39 mm diameter Split Sets (percussively driven, friction contact); and
• standard Swellex bolts (inflated, friction contact). The elements were grouted into thick-walled steel tubes, or installed in boreholes formed in simulated rock in steel tubes. The dimensions of the tubes were chosen so that the tubes provided confinement of the same order as that provided by the rock mass. Each test specimen consisted of two lengths of tube, butted together. The butt provided the joint at which failure of the rockbolt element could take place.
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TABLE 2 Results of impulse load testing of rockbolts. Rockbolt Type
Total Resistance (kN)
Amount of Explosive (kg)
Maximum Height (m)
Maximum Velocity m/s at distance S mm
Damage to Tendons
16 mm cone
290
0.5
0.33
6.5/180
Nil
16 mm cone
290
1.0
1.80
8.6/80
All pulled out but totally undamaged
16 mm smooth
700
0.5
0.38
7.3/20
Nil
16 mm re-bar
725
0.5
2.03
8.0/50
All broken
22 mm cone
1035
1.0
0.50
12.8/55
Nil
25 mm re-bar
1350
1.0
4.65
10.2/500
Two pulled out three broken
-
0.5
1.8
6.5/20
-
-
1.0
5.20
10.2/100
-
Calibration tests
Large input energies, sufficient to fail all of the element types, were achievable by means of the drop weight impacting on a swing beam. Measurements of the impact velocities showed that velocities in excess of 20 m/s were achieved in some cases. In all of the tests, the behaviour of the rockbolts was characterised by one, or a combination of two or all, of the following mechanisms:
Rockbolt Type
Rupture Energy kJ
Rebar 16 mm
4.8 ± 0.7
-
-
• If the rockbolt was strongly bonded, or prevented from
Smooth bar 16 mm
-
13.9
3.0 × le 0.15*
Vaal Reefs hoist rope strand
50°) - Massive Open Stoping (Combination of closely spaced reefs) Cut and Fill Stoping (Wide reefs) Long Hole Open stoping (Narrow reefs) • Medium dips (9°)
- Mechanised Narrow Vein Stoping (Narrow tabular reefs)
EVOLUTION OF MINING METHODS The mining methods were first conceived during a feasibility study done in 1998, where the geometry of the orebody was based on a conceptual model. The conceptual orebody was deemed to represent the Target reefs well, as a surface borehole confirmed virtually the identical characteristics found in the Loraine Gold Mines, Weltevreden Fan. The methods were progressively altered to suit the latest geological information, and mine design process as detailed below.
The original mining methods The mining methods comprised the following mining methods as depicted in Figure 2:
• long hole open stoping, • massive open stoping, • cut and fill stoping,
Drift and Fill (Wide reefs) Cut and Fill stoping (Wide or multiple reefs) • Shallow dips ( than average and non slaking
>10
6 4.5 - 8
4. Caving 4.1 Multi-Pass Lw Soutirage 4.2 Single-Pass Lw Soutirage 4.3 Sub-Level Stoping
5. Cut and Fill 5.1 Ascending Lw 5.2 R&P 6. Mechanised Stoping (R&P)
+2 MTPA
Low
Medium
Caution: strong/ massive roof As for 2.1
Other
+2 MTPA +2 MTPA
Medium Medium High
Weak
High
Weak
High
Weak
Strong roof and floor
High
Weak
200
200 - 800
>800
Capacity (t/h) Frequency (Hz)
15.7
50
23.8
Amplitude (mm)
2-5
2-4
1-3
Excitation angle (°)
-
30
30
(0 - 20) × 2
(0~50) × 2
(0~50) × 2
3×2
2.25 × 2
1.5 × 2
Weight (kg)
2518
1889
1198
Length (mm)
3800
3600
2800
Width (mm)
2200
1950
1920
Height (mm)
2030
1000
-
-
-
570 - 600
Excitation force (KN) Power (KW)
Grizzly width (mm) Installation angle (°) Lump size (mm)
20
0
21
≤1100
≤1100
1 and b0 >0), Cm,n, the wavelet coefficient, can be evaluated from the wavelet transform as: C m , n = W f ( m, n) = ∫ f ( x ) ψ m , n ( x ) dx
(3)
The values of m and n are used to adjust the frequency and time location of the wavelet in Equation 3. A small value of m produces a high frequency (contracted) wavelet when high time-resolution is required, whereas a large m produces a low frequency (dilated) wavelet when high frequency resolution is needed. The WT’s superior time localisation properties are due to the finite support of the analysis wavelet. As n increases, the analysis wavelet traverses the length of the input signal, and m increases or decreases in response to changes in the signal’s local time and frequency content. Finite support implies that the effect of each term in the wavelet representation is purely localised. This feature also sets the WT apart from the Fourier Transform, where the effects of adding higher frequency sine waves are spread throughout the frequency axis. Discrete methods are required for computerised implementation of WT. The Discrete Wavelet Transform (DWT) is derived from the CWT through discretisation of the wavelet Cm,n(x). The discrete wavelet transform using multi-resolution signal decomposition is described in (Mallet, 1989). Initially, the signal is filtered through a low pass filter (LPF) and its quadrature mirror filter (QMF). For the next level of resolution, the output of the low pass filter (LPF) is decimated and again filtered. The output of the QMF at each level m, is termed as the wavelet coefficient at that level of resolution. The coefficients Cm,n have a very desirable property: if the signal is smooth, the coefficients are small in magnitude, and if there is a jump in the signal, the magnitude of the coefficients show a significant increase. The coefficients are the detailed part of the signal. Therefore, Cm,n is expected to be large at the locations where impulses start in a vibration signal with a worn bit. Although the index n indicates time, the index m, does not directly indicate frequency, instead it represents a scaling factor which can be related to a multiple of mother wavelet frequency.
898
EXPERIMENTAL SETUP AND PROCEDURES In order to test the feasibility of using the Wavelet Transform approach for detecting bit wear, laboratory scale experiments were conducted. Figure 1 is a schematic of the experimental configuration which consists of a rotary drill with core bit, a block of rock, accelerometers, non-contacting sensors (laser vibrometer), and a data acquisition/analysis system. Several intact rock samples of basalt and limestone were used for the drilling tests. The rock block (basalt or limestone) of dimension 0.61 m × 0.61 m × 0.91 m (2.0 ft × 2.0 ft × 3.0 ft) was cast in concrete and secured on a work bench with metal plates and bolts to reduce the vibrations of the rock block during drilling. A rotary drill with core bit was used to drill 3.81 cm (1.5 inch) diameter holes to the full depth, 0.91 m (3.0 ft) of the rock sample. In order to measure the transverse vibrations of the of the drill core barrel, two laser vibrometers were set up such that each laser beam was projected normally to the drill core barrel and 0.3 m (1.0 ft) away from the barrel. A tri-axial accelerometer was mounted on the drill head, and another on the rock sample. The laser vibrometers and the accelerometers on the drill head monitored the vibrations of the rod and drill respectively, and they provided pertinent information on the health of the bit during drilling. The accelerometers on the rock sample monitored the vibrations and the response of the rock sample during drilling. Prior to conducting the tests, the total system response was calibrated including the accelerometers and vibrometers using the ratio calibration method. Data from the accelerometers and vibrometers were monitored and analysed on-line using a data acquisition and analysis system. Since it was anticipated that it would take a considerable length of time for the bit to wear, the tests were conducted initially with a new core bit and a worn bit. Later tests on bit wear were conducted where bit wear was accelerated by grinding the core bit by a known amount after the first two holes were completed. The process of grinding the bit by a known amount at intervals of two holes continued until the bit was completely worn. Data acquisition and processing were continuously carried out during the entire process. Pull-down force and rotary speed were invariant as drilling progressed.
SIGNAL PROCESSING AND ANALYSIS Vibration signals from the drill, core barrel and rock were processed using a 16-channel data acquisition system and analysed using the Wavelet Transform (WT) method. The WT method has been shown to be effective in the detection of any defects or malfunction in a mechanical system such as tool breakage in milling or bearing defect (Mallet, 1989). Typically vibration signals from a defective mechanical system consist of a representation of the defective frequency and the natural frequency at resonance. The energy of the vibration from a defective system is normally greater than the system vibration. Therefore, in order to detect a defect, such as bit wear, it is necessary to determine the natural frequency of the resonance and the defective frequency. The later frequency produces periodic impulses. Bit failure in rock drilling may occur due to normal wear, where failure occurs gradually or due to fatigue where failure occurs abruptly. In the case of normal wear, a gradual increase of the overall energy occurs with progression of the defect being evident. For fatigue failure, there is no significant increase in the overall energy prior to failure. It has been shown by several researchers (Martin and Thorpe, 1992; Springer, 1988), based on studies on bearings, that bearing health should be monitored in the frequency domain, whereas defect period or frequency should be identified in the time domain. Wavelet Transform is a joint frequency-based and time-based analysis technique which should be very suitable for detection of bit wear, as well as monitoring bit health.
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EXAMINATION OF BIT WEAR IN ROCK DRILLING THROUGH WAVELET ANALYSIS
Theoretical analysis and simulation (Pandit, Joshi and Paul, 1993) show that early detection of defect depends not only on the signal-to-noise ratio of the signal, but also on the color of the noise. For colored noise with a given signal-to-noise ratio, the residual spikes produced by the WT method are very prominent. In the presence of a low signal-to-noise ratio defective signal buried in noise, WT identified the defect by means of frequency
decomposition. Therefore, the WT method provides the time and frequency information simultaneously for detecting bit wear and monitoring bit health. WT provides for the accurate estimation of spectrum from noisy data, power decomposition in frequency bands, characteristic residuals, auto-correlation and impulse response of the system.
FIG 1 - Schematic of experimental set-up.
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F O OTUONYE
RESULTS The results of tests for a new bit at the early stages of wear and a worn bit show there were five bands of frequency appearing around five center frequencies. However, the spectra of the two bits show an apparent difference on the distribution patterns of peaks. For the new bit, most of the energy was concentrated over high frequency bands, while for the worn bit, the energy was distributed more evenly over the whole frequency range. This can be explained by the fact that, for the new core bit, the comparatively uniform diamond inserts excited the accelerometer, and therefore the spectrum of the signal represented more concentrated peaks. Although the difference has been observed over the spectra of signals for the two bits, the time information embodied within the signals was not observed. It was impossible to determine when the change in the spectrum pattern occurred. Wavelet Transform sets out to overcome this drawback by presenting the vibration signals simultaneously in time and frequency domains.
In Figures 2a and 2b, wavelet transform coefficients, Cm,n are given as spikes distributed on different resolution levels m. The level implies the frequency embodied within the signal and that is given by (2|m|/2N)*Fs if the block size is 2N and the sampling frequency is Fs. Therefore, the maximum possible absolute level is (N-1) according to the criterion of Nyquist frequency. The Wavelet Analysis actually decomposes the signal into approximately two categories: the lower level category and the higher-level category. The former constitutes the components of the stationary part, while the latter constitutes the components of the non-stationary part. Since we are concerned with the transient part, attention is paid to those components within the higher-level category. The coefficients of signal Discrete Wavelet Transform (DWT) are presented with a pike distribution map along the time axis on different levels, in which the level measures the discrete scale of wavelet and can be equivalent to frequency in the Fourier Transform. Two pike maps of the accelerometer signals are
FIG 2 - DWT of accelerometer signal for new and worn bit.
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EXAMINATION OF BIT WEAR IN ROCK DRILLING THROUGH WAVELET ANALYSIS
FIG 3 - Three-dimensional scalogram of accelerometer signals for new bit and worn bit.
shown in Figures 2a and 2b for a new bit at the early stages of wear and a worn bit respectively. The comparison between maps of the two bits indicates that the coefficients of DWT for the worn bit are comparatively larger at level 11 than those at the same level for the new bit. The DWT at level 11 measures the contribution of the components near 256 Hertz and is sensitive to the state of wear. The pike map is presented along the time axis, so that the evolution of the pattern can be definitely related to bit wear by an index obtained from the coefficients of DWT. At the same time, a scalogram can be constructed on the time-frequency plane based on the Continuous Wavelet Transform (CWT). The scalogram is a representation of normalised values of squared CWT in a format of contour or image, showing the distribution of energy over the plane. Two three-dimensional images are presented in Figures 3a and 3b for the same signals as in Figures 2a and 2b respectively. The images demonstrate a great difference between the signals of the two bits. Besides the fluctuations in amplitude observed in the high frequency ranges, the peaks in the low frequency ranges turn out to be prominent for the worn bit. This analysis indicates that bit wear tends to uniformise the energy distribution of the signal over the time-frequency plane.
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CONCLUSION A wavelet-based method has been presented as a diagnostic tool for evaluating and monitoring the state of bit wear as drilling progresses. Impulsive responses appear in the vibration signals at the onset of bit wear. The DWT of the vibration signals is a sensitive index of the impulsive responses. The values of the wavelet coefficients during these impulsive responses increase as bit wear increases. Although the preliminary results are encouraging, additional studies are needed to confirm the laboratory results and to extend the studies in a more practical way to field studies so that both DWT and CWT can be combined to provide a reliable and accurate on-line monitoring method for bit wear. DWT will provide an approach to quantify the state of bit wear while CWT will provide a visual evaluation of bit wear. With the advent of high-speed CPU computers, it is possible to acquire, process and present signals in image format by synchronising the data acquisition system with the CWT algorithm. Further, the development of an on-line visual C-based computer program will accelerate the implementation of real-time monitoring of bit wear using CWT and DWT technique.
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F O OTUONYE
REFERENCES Brown, E T, Carter, P and Robertson, W, 1984. Experience with prototype instrumented drilling rig, Geodrilling, 24:10-22. Chui, C K, 1992. An introduction to wavelets (Academic Press: London). Clark, G B, 1982. Principles of rock drilling and bit wear, Colorado School of Mines Quarterly, Part I and Part II. Cooper, G A, Lesage, M, Sheppard, M and Wand, P, 1987. The interpretation of tricone drill bit vibrations for bit wear and rock type, Rapid Excavation and Tunneling Conference, (8):202-218. Larsen-Basse, J, 1973. Wear of hard metals in rock drilling: A survey of the literature, Powder Metallurgy, 16:1-32. Mallet, S G, 1989. A theory of multi-resolution signal decomposition: The Wavelet representation, IEEE Transactions on Pattern Analysis and Machine Intelligence, 11(7):674-693.
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Martin, K F and Thorpe, P, 1992. Normalized spectra in monitoring rolling bearing elements, Wear, 159:153-160. Pandit, S M, Joshi, G A, and Paul, D, 1993. Bearing defect detection using DDS and Wavelet methods-I: Theory and simulation, in Proceedings Symposium on Mechanics, WAM’93, ASME, PED, 63, pp 285-293. Pfister, P, 1985. Recording drilling parameters in ground engineering, Geodrilling, 34:8-14. Springer, C W, 1988. The role of time domain in analyzing bearing defect, Vibrations, 4(3):14-15. Wang, W J, 1996. Application of wavelet to gearbox vibration signals for fault detection, Journal of Sound and Vibration, 192(5):927-939. Warren, T M, 1984. Factors affecting torque for a roller cone bit, Journal of Petroleum Technology, (9):1500-1508.
Brisbane, Qld, 29 October - 2 November 2000
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Selection of Hoisting Wire Ropes Used in Indian Mines — An Economical Approach A K Basu1 and M K Singh1 Basic components of a wire rope
ABSTRACT Mining is nothing but an art - besides being a science - of extracting valuable minerals from the womb of the mother earth. In this gigantic task, a large number of machinery and equipment are employed judiciously. Wire ropes play a vital role in lifting or hoisting men and materials under hostile/adverse working conditions and constitute a major item of investment on machinery/equipment. The economy of using a wire rope is a concern for engineers, users and the planners. The economic use of wire ropes for mining purposes largely depends on the following factors; 1.
Mechanical design considerations - which include tensile strength of wires, elasticity, resistance to abrasion and crushing, fatigue values, static and dynamic loads on wire ropes, factor of safety, fleet angle, etc.
2.
Metallurgical consideration - Composition and heat treatment of wire materials, defects like inclusions and voids, surface quality factor, lubrications, micro structural parameters.
3.
Operational consideration - Seizing and cutting of rope, storage, care and maintenance, cutting back, etc.
4.
Environmental consideration - Temperature and humidity, pit condition, dust, corrosive atmosphere, etc.
A wire rope has three basic components - the wire, the strands and the core. A wire essentially consists of high carbon killed steel, cold drawn to a designed degree. Depending upon the application several such wires are wrapped round a lubricated core which may be of fibre or a metallic wire (IWRC) to form a strand. A number of strand are twisted in the form of a helix. The dia of the rope and that of the wire are related as d = 1.5dw√n, where d = dia of rope, dw = dia of wire and n = No of wires and assembled together to form a rope. In case of a locked coil, full or half locked, the individual wires are laid layer by layer. A wire is, in fact, an assembly of a number of moving parts.
Unique property of wire ropes Because of the cold drawn fibres developed due to a specialised heat treatment called patenting, the wire ropes achieve a unique property of anisotropy capable of taking directional load. The flexibility of structure enables a wire rope to withstand off longitudinal stress like bending, torsion, etc in mining and other engineering applications. Tensile strength of wires gets enhanced remarkably due to the effect of cold drawing, eg heat treated block of steel 1200 Mpa, after cold drawing, achieves strength of 3500 Mpa. The various properties of high carbon steel in different structural shape having equal cross-sectional area shown in Table 1. All the above stated reasons have made wire ropes find wide applications in the mining industry, chiefly for lifting or hoisting men and material under different environments.
In the present paper, an attempt will be made to suggest steps for achieving optimum economy while using the wire ropes for extraction of precious minerals and coal keeping in view the points noted earlier. Once technical viability is established, economy is the ultimate consideration for an industry for safe hoisting of men and material.
INTRODUCTION What is a wire rope? By definition, a wire rope is an assemblage of parts that transmit forces, motion and energy from one part to another, in some predetermined manner and to some desired end. A wire rope is thus a machine of complex construction but of sophisticated nature requiring lot of care, consideration and lubrication for its longevity.
1.
Construction of wire ropes Round strand wire rope of different construction are available. Formerly considerable flat strand rope was used in vertical shaft but, in the main, round strand ropes of 6 x 19 classification are now employed. All have fibre core and most of them are made of improved plow steel and have highest strength. In the 6 x 19 classification, there construction are specified 6 x 19 seale (1/9/9), the 6 x 21 filler wire (1/5/5/10) and 6 x 25 filler construction (1/6/6/12) . In addition to this, construction of round strand ropes of 6 x 7, 6 x 37 and 8 x 19 are also available. 6 x 37 ropes are used for mine drum hoists of medium capacities and in shafts vulnerable to atmosphere corrosion. 6 x 7 construction ropes are used where effects of bending fatigue is negligible. They are used more as haulage ropes than winding ropes. 8 x 19
Scientists, Central Mining Research Institute, Barwa Road, Dhanbad - 826001, Bihar, India. E-mail: [email protected] or [email protected]
TABLE 1 Properties of high carbon steel of different shapes. Shape
MassMin 2000
Size in mm
Modulus
Strength, Ten
Mpa Comp
Bending behaviour
Torsion behaviour ---
3.2 x 0.4
200
850
800
??
1.13 x 1.13
200
1400
1400
--
--
1.27 Dia
200
2100
2100
+
++
1 + 6 + 12 x .3
200
3500
—-
+++
??
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A K BASU and M K SINGH
classification are not recommended for multi-layer coiling on drums. Most of the round stranded wire ropes are being replaced by flatten strand ropes due to more contact area on the drum/sheave result into less frictional effect on the rope. Locked coil ropes do not fall under the conventional classification as there no individual strands and the whole rope forms a strand. Locked coil ropes of two types (I) full locked and (2) half locked. Only full locked coil are used on drum and friction hoists. They are made in size up to 65 mm. Locked coils ropes offer greater resistance to abrasion, and size for size, they have a greater metallic section. They have greater strength than stranded ropes of the same grade of steel. Moreover, locked coil ropes are non-rotating under all loading conditions. They are free from internal corrosion. The average life of a locked-coil lies between 200 000 and 300 000 trips per rope as compared to average life of 72 000 trips per rope for stranded ropes.
3.
For maximum resistance to bending fatigue, a flexible rope with smaller outer wire is required, whereas to obtain maximum resistance to abrasion, a less flexible rope with large outer wire is needed. It is, therefore, necessary to compromise between the two factors. An ideal rope should be one which is sufficiently flexible to run smoothly over the sheaves or pulleys and round the drum and yet composed of largest dia which will correspond to its flexion efficiently to enable the maximum life and wear to be obtained from the rope. Fatigue failure, especially of the bending fatigue type, is very common in wire ropes. A rope passing over the sheaves or pulley and round the drum has to undergo repeated cycles of bending and unbending. This phenomenon introduces bending fatigue over the entire length of rope. Datas derived from the test laboratories support field experience that the effect of variation in pulley/rope bending ratio, rope tension and rope construction are predetermined in fatigue performance of a rope. Preforming increases a rope’s resistance to bending fatigue. Surface quality factor which contributes to fatigue in the wire and rope, has a predominant role in the performance behaviour of wire ropes (Basu, Mandal, Singh and Singh, 1999).
Purpose of using hoisting rope - safety of men and material The sudden and premature snapping of a rope, while in use, may generate a lot of troubles including administrative problem, payment of compensation to the victim(s) of the accident, stoppage of the work schedule, damage to the equipment, loss of productivity and cost of its replacement. All the above factors do adversely affect the economy of the operation. Safety of men and material is therefore, should be goal in deciding economic use of wire ropes employed in hoisting/hauling of load from / in the mines. 4.
SELECTION OF ROPE
Resistance of wire to bending fatigue is inversely proportional to its tensile strength and to the size of the outer wires of the rope strands.
Ratio of pulley/rope dia (D/d)
Wire rope of different shape, size, construction and strength are available. It rests on the discretion of the user to select the proper rope to suit a particular set of conditions so as to achieve maximum economy out of its use. It is unwise and uneconomic to fit a round nut in a square hole. Economy of any enterprise is basically related to the cost of input involved in it. While selecting wire ropes the following points need careful consideration:
This ratio can be defined as D/d and is called the Bending ratio
1.
mechanical design consideration,
2.
metallurgical consideration,
3.
operational considerations, and
4.
environmental conditions.
Ropes operating with high factors of safety and low bending ratio will deteriorate faster than those with lower factors of safety and higher bending ratio. In order to reduce the bending stress in the rope, the sheaves and drums dia are commonly recommended are given in Table 2.
Where D = Dia of the pulley, sheave or drum d = nominal dia of the rope
Mechanical design consideration From mechanical design points of view size, shape, construction of a rope are decided considering the pit depth, load to be lifted/hauled, frequency of cycles of operation rope speed, etc. There are a number of technical aspects which should be considered for safety and efficiency of steel winding ropes for design of installations. Many a designer of hoisting ropes have a tendency to call for ever increasing tensile strength (TS) in wire ropes to achieve maximum overall economy in production. However, they will do well to consider the following points. 1.
2.
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TS of a wire rope depends upon chemical composition, particularly of C and Mn and subsequent processing in wire mill. It has been shown in a (Mukherjee, Chakraborty and Roy, 1979) research paper that rope life is not always increased in proportion to strength of wire. Resistance of a wire to wear and abrasion is proportional to its TS and the size of the outer wires of the rope strand. However the higher the tensile strength of a wire, the lower is its ability to withstand the effects of surface work hardening following heavy abrasion.
TABLE 2 Sheaves and drums dia for different types of rope. Types of rope
Preferred
Minimum
6x7
72 d
42 d
6 x 19
45 d
30 d
6 x 37
27 d
18 d
8 x 29
31 d
21 d
Locked coil
100 d
100 d
For deep shaft:
100d
For shallow shaft:
80d
For underground:
50d
Rollers used to support the hoisting rope in an inclined shaft should not less than eight times the rope diameter. Rubber faced roller prolong the life of the rope as well as of the rollers.
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SELECTION OF HOISTING WIRE ROPES USED IN INDIAN MINES — A ECONOMICAL APPROACH
5.
6.
7.
8.
9.
Sheave/Drum grooves: It is important that the sheave and drum grooves match the diameter of the rope. Too small a groove will pinch and bind the rope causing excessive wear, fatigue and wire breakage in stranded rope or distortion in locked coil ropes, thereby shortening the rope life. Too large a groove will not provide adequate support and the rope may be distorted. The groove dia should be slightly larger than the nominal rope dia. At least one-third of the rope circumference should be supported by the groove. Depending upon the fleet angle and the angle between flanges, angle of contact for proper support lies between 130° to 150°. The depth of the sheave groove should be at least 3/2 times the rope dia. When the groove wears appreciably, it must be trimmed or regrooved or the sheave replaced before a new rope is installed. Fleet Angle: Experience has shown that the best wire rope service is obtained when the fleet angle is not more than 1°30’. If the fleet angle is lower than the recommended limit, the rope will not pull away from the flange soon enough, enough, tends to pile up against the flange head for 2/3 wraps and drops down suddenly, ie causes overwinding. Overwinding is harmful both to the rope and drum. If, other hand, the fleet angle is too large, the rope will be too quick to pull away from the drum flange head leaving gaps into which the coil of the next layer will fall. There will be excessive rubbing of the rope against the sheave groove. The net result of a larger fleet angle is crushing and abrasion of the rope. Preforming and postforming of ropes improve the resistance of high tensile rope wires to severe bending stress. The use of preformed and post formed ropes is advised if load condition demand high tensile wire in the ropes.
safety selected should be sufficiently high to allow for its reasonable factor of safety till the rope is discarded or deteriorated in service. However in Indian mines regulation prescribe the factor of safety for hoisting for hoisting ropes are given in Table 3. TABLE 3 Service
FS
Rope Speed (rpm)
8 - 12
75 - 375
2. Mine Hoist a. up to 100 m b. 100 m to 150 m c. 150 m to 300 m d. 300 m - 600 m e. 600 m - 900 m f. Over 900 m
10 8 7 6 5 5
— 360 - 480
3. Misc hoist
5
60
4. Haulage ropes
6
600
1. Elevators
The rope bending ratio induces secondary bending stresses in the wires and the factor of safety exert greater influence on the life of a rope. The British regulations take into account the bending ratio in the formula based on experience for calculation of FOS for any given shaft depth. The minimum FOS for friction hoist ropes are given by F2 =
4.5 (R + C) for mineral / material hoisting R(1 + 0.005 L ) − 13.5
Resistance of a rope to crushing over drums or sheaves is directly proportional to the tensile strength of the wires and inversely proportional to the number of wires and no of strands. In this case, the improved service life of a rope can be achieved by using preformed and post-formed ropes. In so far as the resistance to crushing/abrasion of a rope is concerned, it is better to use rope of higher TS and more no of wires and strands.
F2 = 1 +
Elasticity or residual ductility is the most desired property in a rope, which, in actual practice, is submerged in the vision of having a higher Factor of Safety (FOS). But FOS does not necessarily reflect the capacity of a rope to meet the combined stress of a particular service condition. An important factor governing the life of a rope is its ability to ‘absorb’ stresses which it can do by virtue of its length, size and volume in relation to the condition under which it operates unless the operational stresses are capable of being consistently absorbed by the rope. Merely numerical higher initial factor of safety (FOS) are not without any reasoned meaning.
=
FOS = Static Load =
Minimum breaking load of a rope Maximum static safe working load wt of rope + wt of hoisting load +Bending stress + stress due to change of speed
To obtain useful economic life of a hoisting rope, operating under given service condition, a FOS with reasonable safety margin should be allowed. Use of a rope excessive large length or one of higher strength with too low a factor of safety will result in excessive rope cost. The factor of
MassMin 2000
750 - 900 900
4.5 (R + C) for man hoisting R(1 + 0.005 L ) − 13.5
Where R=
Ratio of dia of sheave/drum dia
C=
35 where no rope deflecting sheave is used
L=
43 where rope deflecting sheave is used Vertical distance (meters) between level of the top of the highest driving sheave/drum and level at which the hoisting ropes meet the suspension gear of the conveyance when at its lowest position.
The German regulation prescribe the following FOS FOS,
ν ≥ 7.2 - 0.0005 H
for mineral hoisting
ν ≥ 9.5 - 0.001 H
for man hoisting
Where, H=
hoisting distance in meter
According to the German standard, the load bearing capacity of a rope is used instead of its rated breaking load. Selection of FOS with drum and koepe hoisting in the USA, Canada and Europe is given in Figure 1. 10. Constructional stretch: Permanent elongation of rope occurs due to lengthening of rope lay, compressing of strands on the core and wires in the strands during the first few days of operation in the field. It depends upon the type of core, rope construction, laylength and material composition. Allowable constructional stretches are given in Table 4
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A K BASU and M K SINGH
FIG 1.
between Design Factor and Relative service life is shown in Figure 2. The value of design factor 5 is desired frequently.
TABLE 4
Rope Construction 6 Strand FC 6 Strand IWRC 8 Strand FC
½% - ¾% ¼% - ½% ¾% - 1%
11. Vibrational fatigue is due to cyclic loading operation. At starting hoisting rope has low swing, but the frequency of swing becomes gradually high as the cage reached at the top. The vibration is torsional as well as transverse. The vibrational enery is absorbed by the point of contact on the rope. So a consideration for the vibrational fatigue may be taken into design. At last but not the least Design factor (Wire Rope Users Manual) is defined as the ratio of the Nominal strength of wire rope to the total load it expected to carry. The relationship
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METALLURGICAL CONSIDERATION
Approx Stretch
Considering from the metallurgical point of view a good rope must conform to the following points: 1.
The wire of the rope are made from good quality high carbon killed steel, cold drawn through well lubricated die giving 80 - 90 per cent reduction in area through a number of passes.
2.
The steel should be free of harmful inclusion and voids.
3.
The heat treatment applied to rod material, ie ‘patenting’ should be done in a controlled manner so as to generate the desired fibre structure in the rope.
4.
On micro-examination, the wire material should reveal uniform distribution of sorbitic pearlite in a ferritic matrix. There should not be any pronounced decarburisation at the periphery of the wire.
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SELECTION OF HOISTING WIRE ROPES USED IN INDIAN MINES — A ECONOMICAL APPROACH
FIG 2.
5.
The surface quality of the wire must be good without any visible surface flaws, cracks, scratches, nicks, corrosion pittings, etc. A bad surface quality of wire breeds fatigue in the wire material.
6.
The fibre core of the rope should be well lubricated to avoid formation of internal corrosion on the surface of the wires in the strands in contact with the core.
7.
The rope should be subjected to periodical non-destructive testing to assess the health of the rope and condition monitoring.
8.
The rope should have adequate ductility and toughness to absorb impact load during operation.
OPERATIONAL CONSIDERATION Seizing and cutting of a rope The life of a rope may be significantly shortened if proper seizing and cutting operation aren’t performed. When a rope is carelessly seized before cutting, the ends become distorted and flatten after cutting. Non-formed ropes require two seizing on either side of the place of cut, the second seizing being away by six times rope dia from the cut. In case of performed ropes, one seizing is enough.
Storage The life of a new rope depends, to a large extent, on the method of storage. It should be stored in a clean dry place free from smokes, fumes and protected from water care should be taken against moisture also. The place of storage should be protected from direct sun rays and rain. Rolling of the reels should be avoided. Wire ropes should be stored by methods of stacking layer by layer.
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Care and maintenance of the hoisting ropes Hoisting ropes are often subjected to many forms of abuse due to selection of incorrect rope size, construction and grade, improper design and maintenance and incorrect operation. By taking care and maintenance, some of the causes of abuse can be eliminated/minimised which will provide economical rope life. Thus unreeling of the rope and winding the rope on drum calls for maximum care and expertise. Proper lubrication of the rope at regular intervals to replenish the lubricant lost in the normal operation is essential for obtaining maximum rope life as it will reduce wear and retard corrosion. A suitable lubricant specially made for wire rope application with necessary characteristics must be selected for the operating conditions.
Cutting back Cutting short length (2 m) of the rope subjected to severe wear and fatigue, new section of rope are exposed localised abuse. A short length of the rope may be cut off at the rope - end attachment or anchorage end or at both the ends. In order to increase the service life of a rope in this way, it is necessary to install a rope slightly longer than normal required.
ENVIRONMENTAL CONDITIONS The characteristics of the environment in which the rope is working must be clean, dry, free of flue, dust and foul gases like SO2, NO2, CO2, CO, NH3, etc. Humid atmosphere aids corrosion attack on the surface of the wires. Wet pit water should not be acidic. To resist atmosphere corrosion, galvanised wire should be used. Ambient temperature of the pit is also an important point to be considered.
CONCLUSION From the foregoing discussion, it can be concluded that to make economic use of wire ropes, the followings points need to be considered.
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1.
A hoisting rope for a specific duty at a specific working environment should be selected after considering the load to be lifted, depth of the pit, condition of the pit, FOS to be provided, construction and grade of the rope, No of cycle of operation, rope speed, acceleration and deceleration, etc.
2.
Care must be taken for bending fatigue which is likely to affect the performance of the rope.
3.
The wire of the rope should be examined metallurgically before installation and after a specified period of use as a regular check-up of its health condition.
4.
The design of the installation has to be checked properly.
5.
Proper care and attention must be given for storage of the rope, unreeling, seizing and cutting, handling and transportation of the rope.
6.
There must be periodical check-up for internal lubrication of the rope and fibre core. The rope should be tested by NDT to assess its condition monitoring.
7.
Winding ropes, excavating ropes and hoisting ropes are made of high tensile wire, harder to handle, have a greater tendency to kink. Such ropes should be preformed. Preforming makes them more flexible, easier to handle and full advantage of their higher tensile properties can be utilised. Preforming also provides resistance to bending stress.
8.
The rope selected should have maximum reliability and low consumption.
According to the Indian Mines Regulations, the maximum permitted life of a drum winding rope is 3.5 years and that for friction winder it is two years. In Indian mines, locked coil ropes are mostly used. These ropes are discarded after the expiry of the
scheduled period without investigation as to its ability for further use. There are reasons to believe that if residual ductility/residual stress of these ropes are ascertained before discard, the result might prove economic. To achieve economic use, attempt should be made to extend the rope life as well as proper selection of the rope (Appendix 1). Final selection of a rope may be modified in the light of actual experience gained from use of different rope conditions. It has been roughly estimated that if rope life is increased by one year the total hoisting cost is saved by 33 per cent of the cost of a new rope. Proper selection and handling of the rope, its care and maintenance are key factors in economic use of a rope. All the factors which affect rope life and ultimately the economy of its use are listed in Tables 5, 6 and 7.
ACKNOWLEDGEMENT The authors express their sincere thanks to Dr A K Dube, Director, CMRI for his kind consent in allowing the paper to be published. The views expressed in this paper are of the authors and not by the institute they belong to.
REFERENCES Australian wire ropes - Published by Australian Wire Industries Pvt, Ltd. Basu, A K, Mandal, N, Singh, M K and Singh, T N, 1999. The Impact of the Surface Quality Factor in the Performance Behaviour of Steel Wire Ropes for the Mines, in 1999 Conference proceedings of The Wire Association International, Inc Wire & Cable Technical Symposium (WCTS), 69th Annual Convention, Atlanta, Georgia, USA, May 1999, Trans, pp211- 216 DGMS Circular, 1986 Dove, A B (ed). Wire Rope Hand Book, Vol II, pp10 -17, (The Wire Association, Inc, Branford, Conn).
TABLE 5 Design data. Rope size & Breaking load
Rope loading condition
Rope Construction
Type and amount of loading
Lay-Right of left hand
Lay - ordinary/Lang’s lay
Rope bending stresses
Wire grade and Quality
Lay length-strand & rope
Rates of acceleration & deceleration
Wire arrangement & size
Core - Type, size & Quality
Speed of Rope operation
Equipment care & skill in manufacture
Rope - shop lubrication
Angularity of Rope operation
Fabrication
preformed / Non-preformed
Statutory and operational factors of safety and regulation
TABLE 6 Equipment limit, handling and fitting. Drum and sheave conditions
Handling, fitting and installation
Tread and groove diameters of drum and sheaves
Method of transport and storage
Drum and sheave materials Arc of contact of rope on sheave
Size, arrangement and type of equipment
Abrasive wear
Unreeling & uncoiling method
Type of rope service
crushing and cross over in drum
Sizing, cutting & preparation
rope speed, acceleration and braking
cropping, end for ending and retention
Vibration and whipping of rope
Hot, Wet, Gritty and dirt conditions
Arrangements of drum and sheaves
Lubrication and care in storage
Fleet angle between drum and sheave
Drum anchorage and breaking in
Unic bearing pressure of sheave materials
Sizing, cutting & preparation
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TABLE 7 Operational conditions
type of loading Steady or impact)
Corrosion
Regularity of inspection
Lubrication in the field
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SELECTION OF HOISTING WIRE ROPES USED IN INDIAN MINES — A ECONOMICAL APPROACH
Ganguli, K K et al, 1997. Some case studies into inadequacy of safety in mine transportation equipment, Proceedings of 27th International conference of safety in Mining Research Institutes (CS-MRI ’97’, Feb ’97, New Delhi (Oxford & IBH Publication Pvt Ltd). Kejriwal, R Km 1980. The Wire Rope - An Incredible Machine, Paper presented in Colloquium on Wire Ropes at CMRS Auditorium on 10th August, 1980. Mukherjee, D, Chakraborty, R P and Roy, P R, 1979. Role of Surface Quality on Fracture Susceptibility of Cold drawn high carbon steel used in Mines, New Sketch, Annual Number.
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Ramlu, M A, 1996. Mine Hoisting, pp54-125 (Oxford & IBH publishing Co Pvt Ltd, New Delhi). Rope Man’s Hand book - Published by N C B, U K. Sengupta, T K and Roy, P R, 1968. Performance Study of Indigenous Mine winding Ropes - R R No 48, CMRS E6/48, Feb’68, pp 20-21. Walker. H A and Garrens, J, 1950. American Wire Rope practice in Mining - Paper No - 6 622.673.6(73) - proceedings of a conference held at Ashrome Hill, Leamington Spa, warwickshire in Sept, 1950. Wire Rope Users Manual, Third Edition, Wire Rope Technical Board, Woodstock, Maryland.
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APPENDIX 1 Case history of selection of hoisting ropes in Indian mines.
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Alternative Access, Mining and Hoisting for Underground Deposits K Biegaj1 ABSTRACT It is proposed to consider a new access, mining and hoisting system for underground deposits, as an alternative to the commonly used decline or vertical shaft access. The proposed system offers fundamental improvements in gathering of exploration data, reduction of capital expenditure and significant savings in mine operating costs. The cost reductions are achieved through: 1.
superior orebody definition by Supersucker winzing on ore prior to the commencement of capital development;
2.
considerable reduction of capital development in waste;
3.
full capitalisation on all benefits of a raisebored excavation;
4.
introduction of rubber tyred skip and rubber tyred Mine Rapid Response Vehicle;
5.
introduction of simplified ore handling system and reduction of underground mobile diesel fleet; and
6.
introduction of minimum maintenance shaft concept. The system is applicable to deposits ranging from shallow, low dip angle, narrow vein, high-grade to bigger, massive type vertical mineralisations accessed from the surface, preferably from the bottom of an open pit; it is also suitable as a down dip extension of an existing underground mine. The system is highly conducive to mine automation and emerging trend aimed to increase the ratio of mined metal/mineral to mined waste. It offers substantial reduction in diesel exhaust gas emissions to the mine atmosphere.
INTRODUCTION The prevailing trend over the last decades, in relatively small and shallow Australian metalliferous mines, has been to abandon vertical shaft capital access and extensively use decline for ‘flexibility’ and easy underground access of rubber tyred mobile diesel equipment. Generally, mining costs are governed by the ratio of excavated tonnes of ore (metal/mineral) to tonnes of excavated waste, including the waste resulting from a capital development. With respect to the recent rather liberal use of declines as mine access, the ratio of capital waste development tonnes to mined ore tonnes (metal/mineral) has been in general excessive, especially in cases of narrow vein, high-grade, smaller deposits. Unnecessary spending of capital in a climate of declining commodity prices and the infamous reputation of the mining industry for yielding poor returns on capital is not viewed favourably by the investment community, and therefore capital for new mining projects is hard to obtain. Costs benefits resulting from a reduction of capital development in waste and the use of an inclined raisebore(s) as mine access/haulage way, are the main focus of this paper.
Exploration Any justification for capital funds to get a new mining project off the ground is based on the accuracy of exploration data.
1.
MAusIMM, Manager – Mining Projects, Central Norseman Gold Corporation Limited, PO Box 56, Norseman WA 6443; Director of Ausvac Mining Pty Ltd, PO Box 271, Norseman WA 6443. E-mail: [email protected]
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There have been numerous examples of decisions made on an incorrect geological interpretation, which have resulted in subsequent substantial financial losses or conversely, lost opportunities. Three-dimensional aspects of an orebody like shape, mineralised zone distribution and undulation, have been widely and commonly neglected by geologists when producing resource models (and in the resource/ore reserves reconciliation process in production phase). Those parameters are almost solely responsible for unplanned waste development, poorly optimised level spacing, and subsequent ore losses and excessive dilution when stoping. Simply speaking, to reconcile only ore tonnes, grade and metal/mineral mined against the original resource/orebody model, is far from being adequate.
Mine capital access Decline access mines Mine decline access provides flexibility but it comes at a price, especially in view of the recent more stringent safety requirements. Quite often the quoted ‘flexibility’ of a decline access is used to cover for lack of or bad planning practices, in both feasibility and operational stages. Meshing of declines down to 3.5 m from the floor is now a normal practice, with an average cost of $A 3500/m in good ground conditions for a 5.3 m high × 5.3 m wide excavation accommodating a typical 30 tonne truck. In bad ground conditions the price per metre can be substantially higher, often double, where special support like shotcrete/fibrecrete is required. With the increasing depths of mining and further tightening on safety requirements, the price of a typical decline excavation in Australia will more than likely increase further, following the footsteps of Canadian mines as an example.
Ore (metal/mineral) to capital waste ratio As an extreme example, a typical, small Australian narrow vein gold mine producing 100 000 tonnes of ore per year and yielding 700 to 1400 tonnes of ore per vertical m, would have a ratio of ore tonnes to capital development waste tonnes between 0.6 0.9, whilst being accessed via a 5.3 m × 5.5 m decline. Monthly production of ore on a 12 hour continuous shift roster is typically around 8000 tonnes (or approximately 5 × 30 tonne trucks per shift), which is less than 0.5 of an average trucking shift, or speaking metaphorically, less than one lane of the ‘decline highway’ is being utilised. Another words, we are excavating a big decline to efficiently hoist capital development waste. A poignant illustration of this point is to observe a small 1.5 m3 capacity LHD (1.5 m wide by 1.85 m high) travelling up a 5.5 m wide by 5.3 m high decline. The situation described above is not uncommon for small tonnage mines in Australia. In response to safety concerns and in order to meet current safety requirements, a number of mines, even those relatively young (25 years of age or less), have undergone a highly
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expensive and disruptive decline rehabilitation program necessitated mainly by deterioration of ground support or/and less stringent ground support standards applied in the past.
Winder Supersucker
Vertical shaft access mines Modern mines utilising rubber tyred diesel equipment underground and having a sole access to the surface via a vertical shaft, face a major inconvenience in lowering the dismantled equipment through the shaft for assembly in an underground workshop. The operational inconvenience of shifting mining mobile equipment between the levels is normally addressed by an internal decline excavated in addition to the already existing shaft, which represents a doubling up on capital waste development in an often financially unjustifiable manner.
Winze on ore
Rubber tyred Skip/Mine Rapid Response Vehicle
15m
Shaft ore handling system
120m
4.5m dia raisebore
Rock breakers, grizzlies, ore passes and loading pockets are highly capital intensive and time consuming at their construction stage, and very expensive to maintain in operational order throughout the mine life. FIG 1 - Alternative access, mining and hoisting system.
Shaft internal infrastructure a second parallel pilot hole needs to be drilled; both pilot holes need to be over-drilled to a final depth (600 metres down dip in this example).
Maintenance of internal structures of shafts like steel sets, guides, brattice sheets and stages, especially in a highly corrosive wet environment is very expensive and highly disruptive to production cycle.
4.
Transportation of dismantled raisebore head down the winze.
Level ore development and stoping
5.
Raiseboring of an inclined shaft excavation and vacuum lifting of raisebore cuttings to the surface with a Supersucker.
6.
Installation of high tonnage capacity, low speed (2.0 m/s) transportable/semi-transportable winder (West, 2000; Jance, 2000).
7.
Introduction of a rubber tyred hoisting skip/Mine Rapid Response Vehicle to shift mining equipment from the surface and between the levels.
8.
Excavation of level plats and a sump with drill and blast techniques and vacuum lifting of waste to the surface.
9.
Excavation of shaft stockpiles/system surge stockpiles with waste hoisted in a rubber tyred skip directly loaded into it with a LHD.
Often due to lack of accurate geological information on all critical parameters describing the mineralised zone in three dimensions (referred to in paragraph above) and subsequent lack of proper determination on what constitutes ‘ore’, it has been a common occurrence to mismatch the ‘orebody’ with the mining method and selected mining equipment. Selection of mining method and vertical level spacing has often been based on intuition rather than on engineering determination on what produces the best Net Present Value (NPV) for a mining project from exploration to rehabilitation stage. As a result, either excessive dilution and/or ore losses have been encountered, or the mine has been over level-developed (less common situation). Equipment selection has been often based on availability at a particular time and not on careful planning aimed at matching the three dimensional distribution of mineralised zone along the strike and up/down dip, and producing the best NPV for the project.
PROPOSED SYSTEM The proposed system is shown in Figure 1.
Brief description of the proposed system Listed below is an outline of the proposed system in sequence of construction: 1.
Excavation of winze in ore with a Supersucker for superior three-dimensional (3D) definition of mineralised zone in conjunction with other geological data (down to 120 vertical metres).
2.
Determination of mining method and equipment selection for best NPV result.
3.
Drilling of a raisebore pilot hole centrally to the strike length and close to the orebody; in case of a deeper deposit,
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10. Mining of ore with a direct loading into a skip with a LHD in a semi-automated/automated mode. 11. Down dip repetition of the above steps in 150 m stages as in any other development campaign. 12. Bigger and deeper deposits will require a second, parallel raisebore to cater for higher tonnage of hoisted ore and down dip increase of haulage distance; adequate winder capacity and scheduling of down dip extension work will ensure continuous and undiminished level of ore production.
Main features Exploration - Return to winzing A re-instatement of an old concept, proven and widely used in the past known as winzing, is strongly advocated. Obviously the proposed ‘return to winzing’ has very little to do with a shovel, kibble, hard manual labour and all the bad connotations from the past. The winzing concept is outlined in Figure 2.
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ALTERNATIVE ACCESS, MINING AND HOISTING FOR UNDERGROUND DEPOSITS
Increase of ore(metal/mineral) to capital development waste ratio
Open pit Surface ore outcrop
Ore drive
Bottom of open pit Bottom of u/g mine Advanced knowledge by winzing
Sub-level Drill hole
Winze Ore body
By positioning the raisebore(s) underneath the orebody as close as geotechnically advisable, a dramatic reduction in volume of capital waste excavated is achieved, which results in an increase of ratio of ore tonnes to capital development tonnes from current 0.6 - 0.9 to 14 (again, typical small Australian gold mine as an extreme example). Although the ratio of ore (metal) tonnes to capital waste development tonnes is not that low for higher tonnage decline mines, capital waste development tonnes can be reduced up to 100 per cent (depending on a vertical level spacing) when compared to the current capital waste development required to access the orebody.
Utilisation of all benefits of a raisebored excavation From the author’s experience and to the best of his knowledge, the mining industry so far, has only taken advantage of one feature of an opening excavated with a raiseboring technique, ie smooth surface area and shape providing low resistance for ventilating mine air (probably the least important). It is proposed to fully capitalise on other aspects/advantages of a raisebored excavation:
Increased Confidence in Invested Capital FIG 2 - Winzing concept.
When used in combination with drilling and modern exploration techniques currently on the market, winzing is the best known method of gaining an advanced knowledge of all three-dimensional parameters of a mineralised zone and reducing the risk on investment capital on a mining project. Having many advantages, winzing as any other method has its limitations and those will be briefly discussed in this paper.
1.
Use an inclined raisebored excavation as a main haulage and service access for ore hoisting and transportation of mining mobile diesel equipment, materials and personnel from the surface and between the levels.
2.
Lack of damage to the excavation normally caused by blasting and therefore excellent stability of the main mine access(s).
3.
Geotechnically perfect and long term stable shape, requiring zero or close to zero ground support - it is assumed from experience in typical Western Australian gold mines, that a raisebore not exceeding 4.5 m in diameter will be stable over a long period of time without any support in good ground conditions. Periodical check scaling of the raisebore will be required.
4.
Excellent road surface requiring zero maintenance if excavated in good ground conditions (which is normally the case).
5.
Close to zero maintenance access shaft due to lack of infrastructure normally installed in a conventional shaft (only winder rope rollers are installed on the raisebore floor).
Supersucker winzing It is proposed to implement modern winzing technique with a Supersucker, a giant vacuum cleaner. Supersuckers have been widely in use in Australia in shallow shaft sinking applications over the last 25 years or so. The technique originated from South Australian opal fields, where they still remain the main ore haulage system. About 160 of them are currently in use in Coober Pedy alone, transporting millions of tonnes of material every year (Vanajek, 1999). Shaft sinking depths commonly achieved in Australia reach up to 180 m vertically, with a 200 mm to 300 mm diameter suction pipe (Maher, 1999; Simpson 1999). Winzing in ore with sub-levelling capabilities allows a superior bulk knowledge of mineralised zone in advance, and when used with drilling data and latest radar techniques, gives the best possible definition of all critical parameters of an orebody, namely: grade distribution and continuity, shape and undulation all in 3D. In addition, the benefits of bulk sampling obtained by winzing for metallurgical purposes are obvious.
Introduction of rubber tyred hoisting skip It is proposed to introduce a simple in construction, rubber tyred skip for ore hoisting activities. The proposed skip, apart from winder rope rollers on the floor, will not require any infrastructure to be installed in a raisebore, which is normally an expensive capital and maintenance item in a conventional vertical or inclined shaft.
Operational benefits of winzing Apart from exploration aspects of winzing, there are a number of operational advantages of this method. Costs of winzing, perceived by some as prohibitive, need to be viewed in a much wider context with all measurable benefits attached to it, namely: superior exploration data in advance, superior tool in stope design, as an excavation nearly paid for if excavated on ore (with an immediate return) and as an excavation of a multipurpose future use for either second means of egress in advance, primary ventilation in advance, longhole slot raise in advance, etc.
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Direct loading of ore into the skip It is proposed to totally eliminate shaft loading pockets and associated infrastructure (expensive and time consuming capital and maintenance items) by direct loading into a skip with a stope LHD. A simple in construction, easily re-positioned chute will be used to prevent spillage. It is proposed to operate the skip in a semi-automated (initially and later in a fully automated) mode, directly linked with stope LHDs operating on different levels to eliminate skip waiting
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time. Capacity surge stockpiles located near the raisebore on each level will ensure fast loading and selection of an optimum level for loading (depending on the positioning of a LHD within the stope mucking and tramming cycle in relation to the empty skip). Recent developments in Australia in the area of laser guided autonomous LHD tramming will be highly conducive to this application with automated stope mucking as a next step to be implemented in practical applications very soon.
Introduction of rubber tyred Mine Rapid Response Vehicle To accomplish swift movement of personnel, mobile diesel equipment and materials from/to the surface and between the levels (which normally presents a major inconvenience/cost in a traditional mine accessed via a shaft), it is proposed to use a simple in construction, low maintenance, rubber tyred vehicle capable of all those functions and called in this paper Mine Rapid Response Vehicle (MRRV). Since the mass of the biggest LHD to be shifted between the levels necessitates high winder capacity, the functions of the rubber tyred skip and MRRV will be combined in the shallow stage of mine life. When in service mode, it is proposed to operate this vehicle in a push-button, automated mode.
10. The proposed dismantling of raisebore head and transportation down the winze has not yet been performed in practice to the author’s knowledge. 11. Limited to good ground conditions only to fully capitalise on all benefits of a raisebored excavation. 12. Diameter of the raisebore is not to exceed 4.5 m for local geotechnical stability. 13. Cross sectional area of the 4.5 m diameter raisebore limits the size of the biggest LHD to a 6 m3/8 cubic yard capacity (approximately 2.9 m wide by 2.9 m high), which restricts the maximum tramming distance from the hoisting raisebore to approximately 400 metres. 14. For mines deeper than 100 vertical metres, a second parallel raisebore will need to be constructed (and extended down dip), to cater for uninterrupted ore hoist and increased haulage distance. 15. To the author’s knowledge, rubber tyred skip and rubber tyred Mine Rapid Response Vehicle have not yet been constructed and trialed in practice in a raisebore; stable travelling of those vehicles on the floor of a raisebore needs to be ensured.
Safety, health and environmental benefits 1.
Fourteen times reduction in exposure to rockfalls in capital development headings alone due to reduction of surface area of backs.
2.
Smaller dimensions of shaft and level drives are inherently more stable and rockfall hazards will be easily identified and rectified by barring down with no need for any additional equipment (due to the lack of personnel carriers, the mine will be walked through and inspected on each shift).
Utilisation of shaft cross sectional area
4.
Owing to lack of internal infrastructure/support, which normally occupies a lot of shaft space, the proposed system allows a superior utilisation of shaft cross sectional area for hoisting and equipment shifting.
At least ten times reduction in diesel equipment kW installed underground and diesel exhaust gas emissions to the mine atmosphere.
Costs savings compared to a decline access
Challenges and limitations of the system
The following assumptions have been made for the purpose of this comparison (Tables 1, 2 and 3):
1.
1.
A mine accessed from the bottom of an open pit was considered as an example.
2.
Cash flow analysis comparison was carried out for a typical medium size Australian gold mine producing 720 000 tonnes of ore at 6.0 g/t for 139 000 ounces per year, with 15 m vertical level spacing and utilising top-down longhole mining method. A mine with a life of eight years was considered with a year zero included for the initial capital development.
3.
Adopted mining and milling costs, type, number and size of equipment selected, practices, productivities and personnel levels are based on current experience in Australian metalliferous mines.
4.
Each 100 vertical metres of capital development will provide two years of mine production.
5.
Assumed size of mined orebody: strike length 800 m, true ore thickness 4.0 m, dip 45°, grade 6.0 g/t (in case of massive type orebodies, 400 m tramming distance to the haulage raisebore).
Drastic reduction of underground mobile diesel fleet 1.
Total elimination of trucking fleet.
2.
Elimination of other mobile equipment normally engaged in capital waste development activities in decline access mines.
3.
Total elimination of underground light vehicles fleet.
Reluctance of mine managers and mine planning engineers to accept a new concept departing from a quite comfortable, ‘flexible’ and now commonly adopted decline mine stereotype.
2.
Perceived lack of flexibility when compared with a decline access.
3.
Perceived complications caused by unexpected change in orebody direction, size or presence of additional orebodies.
4.
Personnel access when excavating 150 m deep winze.
5.
Supersucker lifting capacity – 180 m vertical lifting is the limit with Supersuckers currently available.
6.
Ensuring that raisebore pilot holes are straight, especially those drilled at lower angles.
7.
Minimum angle of raisebore of 42° to horizontal to ensure trouble free rilling of cuttings for vacuum lifting (Newnham, 1999).
8.
Supporting/centralising raiseboring rods in the top part of a raisebore while extending it down dip.
9.
Maximum angle of raisebore of 75° to horizontal to ensure proper traction of the skip and MRRV by gravity force.
914
A re-sale value of the winder and headframe installations has not been included in the cash flow analyses, which represents a conservative approach.
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ALTERNATIVE ACCESS, MINING AND HOISTING FOR UNDERGROUND DEPOSITS
TABLE 1 Summary of capital development costs for a typical decline access of 100 vertical m of orebody. Decline 1 in 8
800 m
@ $3500/m
Decline s/piles
6.7 × 25 m
@ $3500/m
586 000
Accesses*
6.7 × 70 m
@ $4500/m
2 110 000
Access s/piles
6.7 × 25 m
@ $3500/m
586 000
Vent exhaust
6.7 × 55 m
@ $3500/m
1 290 000 670 000
Second egress
40 m
@ $2500/m
Other
100 m
@ $2500/m
$2 800 000
CONCLUSIONS 1.
The proposed alternative access, mining and hoisting system offers significant safety, health and environmental benefits over a decline accessed mine due to a drastic reduction of exposure to rockfalls and reduction of underground mobile diesel fleet.
2.
Supersucker winzing (in conjunction with other geological data) will deliver an unsurpassed definition of the mineralised zone in three dimensions and will enable informed mine design and planning.
3.
Capitalisation on all benefits of a raisebored excavation and introduction of a rubber tyred ore hoisting skip and Mine Rapid Response Vehicle will ensure most of the essential flexibilities of a decline with a reduction of capital waste development. It will significantly reduce underground infrastructure, simplify ore handling system and eliminate all mobile diesel fleet involved in mining of capital development waste in a decline mine.
4.
The proposed system offers 40 per cent improvement of mine’s pre-tax cash flow and 69 per cent increase of after tax NPV.
5.
Drastic reduction of waste produced from the mine will significantly reduce disturbance to the environment and will enable mining of deposits located in populated or environmentally sensitive areas.
6.
To increase hoisting capacity of the proposed system, either longer skip will be required with a low speed winder (2 m/s) or the winder will need to be sped up. In both cases bigger winders and headframes will have to be installed and additional capital outlay will follow. Financial calculations for those cases extend beyond the scope of this paper.
7.
Financial comparison of the proposed system against a traditional vertical shaft access for deeper mines (where a decline access cannot be justified) is yet to be conducted - it is expected costs benefits will be substantial.
250 000
Total
$8 292 000
TABLE 2 Summary of capital development costs of the proposed system. Supersucker winze
150 m
@ $6000/m
900 000
4.5 m inclined risebore
150 m
@ $6500/m
975 000
Vacuum lifting of cuttings
A conservative cut of 15 per cent has been applied to the mine operating costs in the proposed system. The savings are based on automation of winding, automation of LHD tramming and skip loading activities, reduction of maintenance cost of mobile fleet, reduced primary ventilation power consumption and reduction of number of mine personnel.
@ $750/m
112 500
Plats*
6.7 × 25 m
@ $6000/m
1 005 000
Plat s/piles
6.7 × 25 m
@ $3500/m
586 000
Raisebore head access
25 m
@ $6000/m
150 000
Other
100 m
@ $2500/m
250 000
Total
$3 978 500
* Same mine parameters as listed under Table 1.
TABLE 3 Summary of cash flow for the two mine accesses considered (in million of Australian dollars). Proposed System
Decline
$ Diff
%
Winder
10.7
0
Surface inst
1.0
2.6
1.6
-60
Capital dev U/g mobile eq
23.7
35.1
11.4
-32
11.2
26.0
14.8
-57
Winzing profit
1.9
0
Total capital
44.8
63.7
-18.9
-30
Total mining
34
40
-6
-15
Milling and cart
16
14
2
14
Non cash costs/year:
6.1
7.3
-2.3
-30
Capital:
Operating costs ($/t of ore):
Total cash flow (undiscounted): Before tax
142.6
101.7
40.8
40
After tax @ 30 %
101.7
71.9
72.6
40
Before tax
76.5
46.9
29.6
63
After tax
53.6
31.7
21.9
69
REFERENCES Personal communications: Jance, J, 1999. Principal Consulting Engineer, Western Australian Department of Minerals and Energy, Perth, Western Australia. Jordan, A, 1999. Director of Jordan Mining Pty Ltd, Perth, Western Australia. Maher, P, 1999. Director of Ausvac Mining Pty Ltd and Maher Mining Contractors Pty Ltd, Kalgoorlie, Western Australia. Newnham, L, 1999. Operations Manager of Byrnecut-Ruc, Kalgoorlie, Western Australia. Simpson, B, 1999. Director of Vacuum Mining Pty Ltd, Perth, Western Australia. Vanajek, I, 1999. Blower Manufacturing. Coober Pedy, South Australia. West, R, 2000. Manager of WA Operations, Combined Resources Engineering Pty Ltd, Perth, Western Australia.
NPV @ 12 % discount rate:
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The Influence of Surface Geometry on the Load Transfer Mechanisms of Grouted Bolts — A Laboratory Study N I Aziz1, B Indraratna2 and A Dey3 ABSTRACT The surface geometry of a rock bolt plays a very important role in the rock/resin/bolt load transfer mechanism. Very little work has so far been reported on the influence of the bolt surface geometry on the bolt/resin interface failure mechanism. This paper examines the behaviour of bolt surface roughness under constant normal stiffness condition in the laboratory. To study the shear behaviour of the bolt/resin interface a series of shear tests with an initial normal stress between 0.1 to 7.5 MPa has been conducted on the flattened bolt surface of two most popular bolt types currently being used in Australia. Bolts with deeper ribs offered better shear resistance at low initial normal stress conditions whereas bolts with shallow and closer ribs offered better shear resistance at high normal stress conditions. The maximum dilation occurred at a shear displacement of 60 per cent of the rib spacing irrespective of the bolt type and depth of the rib.
INTRODUCTION Every year a large number of roof bolts are being used to support the mine gate roadways. The load transfer characteristics of the bolt plays a very important role in the design of an effective support system for stabilising rock mass in various types of excavation. The load transfer characteristics of a bolt in the field depends largely on its surface profile configuration among other various parameters. The study of the load transfer mechanism would remain incomplete without the study of bolt surface roughness and the shear behaviour of the bolt-resin interface. In the past, a number of researchers (Fuller and Cox, 1975; Gale, 1986; Fabjanczyk and Tarrant, 1992; Signer, Cox and Johnston, 1997) calculated the shear stress developed at the bolt-resin interface by using strain gauged bolts. The shear stress developed at any point was calculated using:
∆τ =
F1 − F2 πd1
(1)
where,
When used to model the behaviour of a rock bolt, the major shortcoming of the above method is that, it does not consider the effect of the confining pressure on the shear stress at the bolt/resin interface. As the confining pressures or the horizontal stresses around the opening play a very important role in the failure mechanism of grouted rock bolts, incorporation of the confining pressure in the above formula would result in a better approximation of the in situ condition. After the installation of a bolt in the field the relative movement, though very minute, between the rock and the bolt causes an imbalance in the forces acting on the bolt/resin interface which is a function of axial bolt displacement, the radial stress acting at the bolt/resin interface and the stiffness of the host medium. In this process the normal stress at the resin/rock interface is modified by the ribs of the bolt. The effect is dependent on the relative displacement, shape of the rib and the composite stiffness of the bolt/resin/rock interface. The stiffness is a material property which is a constant parameter and cannot vary without any change in material and it is the normal stress which varies during the load transfer process. The Constant Normal Stiffness condition thus represents a better approximation of the deformation behaviour in the field as compared to conventional Constant Normal Load condition. The above hypothesis has been indicated by many researchers (Benmokrane and Ballivy, 1989; Ohnishi and Dharamaratne, 1990; Skinas, Bandis and Demiris, 1990; Indraratna, Aziz and Haque, 1997; Indraratna, Haque and Aziz, 1998). A novel approach was, therefore, adopted to study the shear behaviour of the bolt/resin interface under Constant Normal Stiffness (CNS) condition. In this method, the confining pressures were modelled by applying predefined normal stresses on the bolt/resin interface and the CNS condition was achieved by an assembly of four springs on top of the upper shear box. The applicability of laboratory stiffness to simulate the field condition is described later. The CNS testing machine developed at the University of Wollongong and reported earlier (Indraratna, Haque and Aziz, 1998) was modified to incorporate the surface feature of the bolt for purpose of present study.
BOLT SURFACE PREPARATION
∆τ =
average shear stress at the bolt-resin interface,
F1 =
axial force acting in the bolt at strain gauge position 1, calculated from strain gauge reading,
F2 =
axial force acting in the bolt at strain gauge position 2, calculated from strain gauge reading,
d=
bolt diameter, and
l=
distance between strain gauge position 1 and strain gauge position 2.
1.
MAusIMM, Associate Professor, Faculty of Engineering, University of Wollongong, Northfields Ave, Wollongong NSW 2522. E-mail: [email protected]
2.
MAusIMM, Professor, Faculty of Engineering, University of Wollongong, Northfields Ave, Wollongong NSW 2522.
3.
Doctoral Candidate, Faculty of Engineering, University of Wollongong, Northfields Ave, Wollongong NSW 2522.
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A 100 mm length of a bolt was selected for the surface preparation for CNS shear testing. The specified length of bolt was cut and then drilled through as shown in Figure 1. The hollow bolt segment was then cut along the bolt axis from one side and preheated to open up into a flat surface as shown in Figure 2. The surface features of the bolt were carefully
FIG 1 - Photograph showing hollow bolt segment.
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FIG 2 - Photograph showing flattened bolt surface.
FIG 4 - Photograph showing arrangement of casting resin samples.
protected while opening up the bolt surface. The flattened surface of the bolt was then welded on the bottom plate of the top shear box of the CNS testing machine as shown in Figure 3. Although these flattened bolt surfaces may not ideally represent the complex behaviour of circular shaped bolt surface observed in the field, nevertheless, they still provide a simplified basis for evaluating the impact of the bolt surface geometry on the shear resistance offered by a bolt. Table 1 shows the detailed specification of two types of bolt used for the study.
compressive strength (σc) of about 20 MPa, tensile strength (σt) of about 6 MPa and a Young’s modulus (E) of 7.3 GPa. Such model materials are suitable to simulate the behaviour of a number of jointed or weak soft rocks, such as coal, friable limestone, clay shale and mudstone, based on the ratios of σc/σt and σc/E applied in similitude analysis (Indraratna, 1990). Figure 5 shows a typical cast sample and Figure 6 shows a cast sample in the shear box of the CNS testing machine. The resin samples prepared in this way fitted exactly into the bolt surface during testing and thus allowed a close representation of the bolt-resin interface in practice.
FIG 3 - Photograph showing bolt surface welded on the top shear box plate.
TABLE 1 Bolt
FIG 5 - Photograph of cast resin block.
Core Dia (mm)
Finished Dia (mm)
Rib Spacing Rib Height (mm) (mm)
Type I
21.7
24.4
28.5
1.35
Type II
21.7
23.2
12.5
0.75
SAMPLE CASTING The welded bolt surface on the bottom plate of the top shear box was used to print the image of bolt surface on cast resin samples. The arrangement for casting the sample is shown in Figure 4. For obvious economic reason the samples were cast in two parts. Nearly three-fourths of the mould was cast with high strength casting plaster and the remaining one-fourth was topped up with chemical resin commonly used for bolt installation in mines. A uniform curing time of two weeks was allowed for all specimens before testing was carried out. The properties of the hardened resin after two weeks were, σc = 76.5 MPa, σt = 13.5 MPa and E = 11.7 GPa. The cured plaster showed a consistent uniaxial
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FIG 6 - Photograph showing cast resin samples inside the bottom shear box.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
CNS TESTING MACHINE Figure 7 shows the general view of the Constant Normal Stiffness (CNS) testing apparatus (Indraratna, Haque and Aziz, 1998), built at the University of Wollongong. It consists of two steel boxes, one of size 250 x 75 x 150 mm at the top, and the other of size 250 x 75 x 100 mm at the bottom. A set of four springs are used to simulate the normal stiffness (kn) of the surrounding rock mass. The top box can only move in the vertical direction along which the spring stiffness is constant (8.5 kN/mm). The bottom box is fixed to a rigid base through bearings, and it can move only in the shear (horizontal) direction. The desired initial normal stress (σno) is applied by a hydraulic jack, where the applied load is measured by a calibrated load cell. The shear load is applied via a transverse hydraulic jack which is connected to a strain controlled unit. The applied shear load can be recorded via strain meters fitted to a load cell. The rate of horizontal displacement can be varied between 0.35 and 1.70 mm/min using an attached gear mechanism. The dilation and the shear displacement of the joint are recorded by two LVDT’s, one mounted on top of the top shear box and the other is attached to the side of the bottom shear box.
FIG 7 - Photograph showing CNS testing machine.
EXPERIMENT AND DATA ANALYSIS A total of 12 samples were tested for two different types of bolt surface at initial normal stress (σno) levels ranging from 0.1 to 7.5 MPa. Each sample for bolt type I was subjected to five cycles of loading in order to observe the effect of repeated loading on the bolt/resin interface. Samples for bolt type II was subjected to only three cycles of loading as it was found that the stress profile
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did not vary significantly after third cycle of loading. The stress profile, as described above, is defined as the variation of shear (or normal) stress with shear displacement for various cycles of loading. The σno applied to the samples represented typical confining pressures which might be expected in the field. A constant normal stiffness of 8.5 kN/mm (or 1.2 GPa/m when applied to a flattened bolt surface of 100 mm length) was applied via an assembly of four springs mounted on top of the top shear box. The simulated stiffness was found to be representative of the soft coal measure rock. Figures 8 through 13 show the stress profile and dilation behaviour of the bolt/resin interface for various normal stress conditions for the bolt type I. The difference in stress profiles for various loading cycles were negligible at low values of σno. This was gradually increased with increasing value of σno reaching maximum between 3 and 4.5 MPa. The increased gap between peak shear stress level becomes more prominent between loading cycles I and II as shown in Figures 10 and 11. Beyond 4.5 MPa confining pressure the stress profile gap between I and II cycles decreased again. Similar trend was also observed for the type II surface. At low σno values, the relative movement of the bolt/resin surface caused an insignificant shearing and slickensiding of the resin surface, and for each additional cycle of loading the shear behaviour remained almost constant. However, as the value of σno was increased, the shearing of the resin surface was also increased and the difference in stress profiles for various cycles of loading became significant. At the medium level of normal stress application, the shearing and slickensiding of the contact surface was not sufficient to smoothen the resin surface excessively in the first cycle, and further cyclic loading caused a significant difference in the stress profile. As the initial normal stress was increased beyond this level, the difference in stress profile for various cycles of loading was decreased due to excessive shearing and smoothening of resin surface after the first cycle of loading. The difference in stress profile after third cycle was negligible at all normal stress levels thus indicating no change in surface properties after third cycle of loading as shown in Figures 8 through 13. The dilation of the bolt/resin surface for both the bolt types was characterised by the initial contraction and then subsequent increase in dilation for all the loading cycles. This initial contraction may be due to the early settlement of very fine irregularities of the resin surface. The dilation was reduced with increasing value of σno, correspondingly the increase in normal stress was also reduced. This gives an indication that the behaviour of rock/resin interface at CNS condition may approach to CNL condition at very high σno values. The characteristics of normal stress profiles were similar to dilation behaviour as shown in Figures 8 through 13 for the type I bolt surface. Figures 14 and 15 show the variation of shear stress, normal stress and dilation with shear displacement at various normal stresses for the first cycle of loading for type I and type II bolts respectively. The maximum dilation occurred at a shear displacement between 17 to 18 mm and 7 to 8 mm for type I and type II bolts respectively at different σno values and at different loading cycles. The distance between the ribs for both the bolt types are shown in Table 1. Therefore, it may be concluded that the maximum dilation occurred at a shear displacement of about 60 per cent of the bolt rib spacing. The shear displacement for peak shear stresses increased with increasing value of σno for both the bolt types. This was due to increased amount of shearing of resin surfaces with the increasing value of σno. However, there was a gradual reduction in the gap between the peak shear stresses with increasing value of σno. The shear displacement required to reach the peak shear strength is a function of normal stress being applied and the surface properties of the resin assuming that the geometry of the bolt surface remains constant for a particular type of bolt. This is clearly evident from Figures 14 and 15.
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FIG 8 - Stress profile and dilation of type I bolt at an initial normal stress of 0.1 MPa.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
FIG 9 - Stress profile and dilation of type I bolt at an initial normal stress of 1.5 MPa.
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FIG 10 - Stress profile and dilation of type I bolt at an initial normal stress of 3.0 MPa.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
FIG 11 - Stress profile and dilation of type I bolt at an initial normal stress of 4.5 MPa.
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FIG 12 - Stress profile and dilation of type I bolt at an initial normal stress of 6.0 MPa.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
FIG 13 - Stress profile and dilation of type I bolt at an initial normal stress of 7.5 MPa.
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FIG 14 - Stress profile and dilation of type I bolt for first cycle of loading.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
FIG 15 - Stress profile and dilation of type II bolt for first cycle of loading.
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7
Initial normal stress Normal stress at peak shear
Peak shear stress (MPa)
6 Cycle I
5
Cycle I Cycle II
4
Cycle II Cycle III
3
Cycle III Cycle IV
2
Cycle IV Cycle V
1
Cycle V
0 0
1
2
3
4
5
6
7
8
Normal stress (MPa)
FIG 16 - Variation of peak shear stress with normal stress for type I bolt.
7
Initial normal stress Normal stress at peak shear
Peak shear stress (MPa)
6 5 4
Cycle I
3
Cycle I Cycle II
2
Cycle II Cycle III
1
Cycle III
0 0
1
2
3
4
5
6
7
8
Normal stress (MPa)
FIG 17 - Variation of peak shear stress with normal stress for type II bolt.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
Figures 16 and 17 show the variation of peak shear stress with normal stress applied for various loading cycles for type I and type II bolts respectively. For the type I bolt surface, the graphs of cycle I through cycle III show a bi-linear trend whereas the graphs representing both cycle IV and V show only a linear trend. For the type II bolt surface, only cycle I shows a bi-linear trend and cycle II and III show a single trend. At low initial normal stress, the shearing of resin surface is negligible and hence the rate of increase of peak shear stress with respect to normal stress is high. At higher normal stress, the degree of shearing of resin surface is greater and some of the energy is thus utilised to shear off the resin surfaces. As a result, there is a lower increasing rate of peak shear stress with respect to the normal stress. As the samples are loaded repeatedly, the resin surface become smoothened reducing the surface roughness and as a result the rate of increase of peak shear stress is likely to remain constant with respect to normal stress. Figure 18 shows the stress profile of both the type I and type II bolts plotted together for the first cycle of loading. The following were observed from Figure 18.
• The shear stress profiles around peak were nearly similar for both the bolt types. However, slightly higher stress values were recorded for the bolt type I at low normal stress levels whereas slightly higher stress values were observed for the bolt type II at high normal stress levels in most of the cases.
• Post-peak shear stress values are higher for the bolt type I indicating better performance in post-peak region.
• Shear displacements at peak shear are higher for the bolt type I indicating the safe allowance of more roof convergence before instability stage is reached.
• Dilation is more for the bolt type I The implications of the above observation need detail explanation. The laboratory experiments were carried out with spring assembly with an effective stiffness of 8.5 kN/mm. In practice the stiffness of resin/rock system will be higher than the laboratory simulated stiffness. As the stiffness increases the effective normal stress on the bolt/resin interface at any point of time will also increase as per the following equation:
is usually restricted to the bolt rib and its surrounding area only. At low normal stress condition, the stress concentration on the resin surface around the bolt ribs is not sufficient to cause resin failure, however, at high normal stress condition, the stress concentration on the resin surface due to the projected ribs of the bolt surface reaches a value which is higher than the compressive strength of the resin causing it to crush around that zone. The stress concentration is more evenly distributed on the resin surface for the bolt type II with a rib spacing of 12.5 mm as compared to the bolt type I with a rib spacing of 28.5. At low normal stress values, the stress concentration on resin surface is not high enough to crush it and thus its integrity remains intact. As a result, the bolt with deeper profile will offer higher resistance due to higher dilation. However, at high normal stress condition as the concentrated stress around ribs are high enough to crush the resin surface, the bolt which provide more evenly distributed stress on the resin surface, in other words, bolts with lower rib spacing would offer higher resistance.
CONCLUSION It can be inferred from this study that:
• the shear behaviour of the bolt surface at various confining pressures directly affect the load transfer mechanism from the rock to the bolt;
• the type I bolt offered higher shear resistance at low confining pressure (below 6.0 MPa), whereas, the type II bolt offered better shear resistance at high normal stress conditions exceeding 6.0 Mpa;
• the impact of repeated loading on the effective shear resistance of the bolt/resin interface was influenced by the level of applied normal stress, the number of loading cycles and the surface geometry of the bolt;
• the maximum dilation occurred at a shear displacement of nearly 60 per cent of the rib spacing; and
• the bolt type I showed better performance than that of bolt type II in many respects.
ACKNOWLEDGEMENT σ n = σ no +
K .δv A
(2)
Technical support provided by Alan Grant, Senior Technical Officer at the University of Wollongong, is gratefully acknowledged.
REFERENCES
where: σn =
effective normal stress,
σno =
initial normal stress,
k =
system stiffness,
δv =
dilation, and
A =
area of the bolt surface.
Therefore, in general higher values of effective normal stresses should be observed for the type I bolt as compared to the type II bolt as long as the confining pressure remains low. Higher values of effective normal stress will have a direct positive impact on the peak shear stress values and, therefore, when installed in the field the type I bolt would outperform the type II bolt, particularly at low confining pressure conditions. In a relative movement situation, however minute the movement is between resin and bolt surface, the complete bolt surface is no longer in contact with the whole resin surface. The contact zone between the resin and the bolt surface at that point
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Benmokrane, B and Ballivy, G, 1989. Laboratory study of shear behaviour of rock joints under constant normal stiffness conditions, Proc of 30th US Symp on Rock Mechanics, Tuscon, pp 157-164. Fabjanczyk, M W and Tarrant, G C, 1992. Load transfer mechanism in reinforcing tendons, 11th International Conf on Ground Control in Mining, Wollongong, pp 212-219. Fuller, P G and Cox, R H T, 1975. Mechanics of load transfer from steel tendons to cement based grout, Fifth Australasian Conf on the Mechanics of Structures and Materials, Melbourne, pp 189-203. Gale, W J, 1986. Design considerations for reinforcement of coal mine roadways in the Illawarra coal measures, AusIMM Symposium on Ground Movement and Control Related to Coal Mining, Wollongong, pp 82-92. Indraratna, B, 1990. Development and applications of synthetic material to simulate soft sedimentary rocks, Geotechnique, 40(2):189-200. Indraratna, B, Aziz, N and Haque, A, 1997. Shear strength characteristics of soft rock joints based on constant normal stiffness testing, 16th Conf on Ground Control in Mining, Morgantown, pp 267-273. Indraratna, B, Haque, A and Aziz, N, 1998, Laboratory modelling of shear behaviour under constant normal stiffness condition, Journal of Geotechnical and Geological Engineering, 16:17-44.
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FIG 18 - Comparison of stress profile and dilation of type I and type II bolt for first cycle of loading.
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THE INFLUENCE OF SURFACE GEOMETRY ON THE LOAD TRANSFER MECHANISMS OF GROUTED BOLTS
Ohnishi, Y and Dharamaratne, P G R, 1990. Shear behaviour of physical models of rock joints under constant normal stiffness conditions, Proc Int Symp on Rock Joints, (Eds: Barton and Stephanson), Leon, pp 267-273. Signer, S P, Cox, D and Johnston, J, 1997. A method for the selection of rock support based on bolt loading measurements, 16th Conf on Ground Control in Mining, Morgantown, pp183-190.
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Skinas, C A, Bandis, S C and Demiris, C A, 1990, Experimental investigations and modeling of rock joint behaviour under constant normal stiffness, Proc Int Symp Rock Joints, (Eds: Barton and Stephanson), Leon, pp 301-308.
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The Load-Bearing Process of Fully Coupled Rock Bolts Chunlin Li1 ABSTRACT The paper focuses on how rock bolts are loaded in different rock masses, as well as in pullout tests. Three analytical models for rock bolts are presented in brief, one for bolts in continuously deformed rock masses, one for bolts in blocky rock masses, and one for bolts in pullout tests. The bolt models are used to interpret measurement data for rock bolts from three sources and to help specify the modes of axial load in the rock bolts. It is found that in continuously deformed rock formations fully coupled rock bolts have only one axial stress peak. The position of the axial stress peak moves towards the far end of the bolts with increasing rock deformation. In blocky rock masses several axial stress peaks may appear in the rock bolt. The position of this kind of axial stress peak does not change during the entire loading process. For rock bolts in pullout tests, decoupling of the interface always occurs first at the loading point. The decoupling front moves towards the far end of the bolt with an increase in the applied pull load.
INTRODUCTION Bolts are probably the most common devices used today for rock reinforcement in civil and mining engineering. The performance of rock bolts is understood quite well in a general sense, but it is not so clear how rock bolts behave in different loading conditions. Such knowledge is important in bolt-reinforcement design and in diagnosing the quality of reinforcement. Pullout tests are widely used to examine the installation quality of rock bolts. In practice one cannot pull out the entire bolt if it is appropriately installed. Knowledge of the behaviour of rock bolts under a concentrated pull load is needed in interpreting the data of pullout tests. Li and Stillborg (1999) recently developed three analytical models for rock bolts: one for bolts installed in continuously deformed rock masses, one for bolts installed in blocky rock masses, and one for bolts in pullout tests. These models can simulate the behaviour of rock bolts under different loading conditions. This paper focuses on the modes of axial load in rock bolts under different external loading conditions.
MODES OF AXIAL LOAD IN ROCK BOLTS Bolts in continuously deformed rock masses In relatively homogenous soft rocks or in extremely fractured rock masses, the deformation of the rock around an underground opening is continuous, with the maximum occurring at the wall of the opening (see the top sketch in Figure 1). The radial differential displacement induces a load in the bolts. In the case of a fully coupled bolt, the decoupling of the interface between the bolt and the rock starts at the wall and propagates towards the far end of the bolt with increasing rock deformation. The axial load in the bolt and the shear stresses acting on the bolt surfaces may be as shown in Figure 1. Freeman (1978) divided the length of a bolt into two parts, the pick-up length and the anchor length. The position between the pick-up length and the anchor length is called the neutral point, where the shear stress on the bolt is zero, while the axial load in the bolt is maximum. The mode of axial load in the fully coupled bolt in continuously deformed rock masses is characterised by the shift of the neutral point towards the far end of the bolt when rock deformation increases. There 1.
Division of Rock Mechanics, Luleå University of Technology, S-971 87 Luleå, Sweden. e-mail: [email protected]
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FIG 1 - The mode of axial load in fully coupled rock bolts in continuously deformed rock masses.
exists only one axial stress peak in the bolt in continuously deformed rock masses.
Bolts in blocky rock masses The deformation of a blocky rock mass is predominantly composed of the deformation of individual discontinuities in the mass. In the case where an underground opening is excavated, the rock joints near the opening will tend to open, which will result in a tensile axial stress in the bolt traversing the joint. Figure 2 shows the axial stress in a fully coupled bolt which traverses three actively opening joints. It is seen in the figure that three stress peaks occur at the positions of the rock joints. The position of this type of stress peak does not change, but its magnitude increases when the joint opens further. Thus, the mode of axial stress in a fully coupled bolt in blocky rock masses is characterised by a stationary position for the stress peak. There may exist several stress peaks in the entire length of a rock bolt installed in blocky rock masses, depending on how many of the traversed joints are activated to open. Hyett et al (1996) obtained similar results through numerical simulations.
Bolts in pullout tests For a grouted rock bolt subjected to a concentrated pull load, P0, the interface between the bolt and the grout (or the rock for frictional bolts) is decoupled first at the loading point. The decoupling front moves towards the far end of the bolt with an
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Tensile axial load in bolt (kN)
150
A c
100
b
50
a
P0
0
0.0
0.5
1.0
1.5
Distance from borehole collar (m)
Shear stress
B
FIG 2 - An example illustrating the mode of axial stress in a bolt subjected to the opening of three joints. The opening displacement in this example: 50 µm at joint a, 20 µm at joint b and 5 µm at joint c.
b
c
a 0
P0
Shear stress
FIG 4 - Sketches illustrating the distribution of stresses along a fully grouted bolt, corresponding to three levels of the applied pull load. (a) the axial load and (b) the shear stress.
FIG 3 - Distribution of the shear stress on a fully grouted bolt subjected to an applied pull load, P0.
increasing applied load. The shear stress is at the level of the shear strength at the decoupling front, while behind the decoupling front the shear stress becomes smaller, since the strength of the interface has been partially lost due to decoupling. The distribution of the shear stress on a fully grouted bolt is illustrated in Figure 3. The axial load in the bolt is illustrated in Figure 4a at three different applied loads. Figure 4b shows the distributions of the shear stress on the bolt at the three applied loads.
CASE STUDIES A fully bonded rock bolt in a soft rock formation The behaviour of fully bonded rock bolts in the Kielder experimental tunnel was studied by Freeman (1978). The type of
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rock was a soft mudstone (with a uniaxial compressive strength of about 37 MPa). Both the radial displacement of the rock around the tunnel and the axial load in the rock bolts were monitored. Figure 5 shows the measurement results for one of the monitored bolts. The curves of the rock movement demonstrated that the radial rock movement was continuous, with the maximum occurring at the tunnel wall. The axial load in the bolt varied with increasing rock movement. After 0.7 days the axial load in the bolt was at its maximum at the tunnel wall and decreased with the distance to the wall. After 1.5 days the peak of the axial load appeared at a distance of about 0.8 m to the tunnel wall. After 98 days the position of the load peak moved further into the rock, with a distance of about 1 m to the tunnel wall. The variation of the axial load in this bolt agrees with the model illustrated in Figure 1, ie the position of the load peak moves into the rock with increasing rock deformation.
Two fully bonded rock bolts in a blocky rock mass Björnfot and Stephansson (1984) monitored the loading process of fully bonded bolts in the iron mine at Kiruna, Sweden, where the rock is hard and strong (with a uniaxial compressive strength of over 200 MPa) and the rock mass contains a number of rock joints. Figure 6 shows the development of the axial load in two of the monitored rock bolts. In bolt A, the axial load had one peak after eight months of installation, while after 14 months two load peaks appeared. Although the magnitudes of the two load peaks increased with time, their positions did not change. It can be concluded according to the model illustrated in Figure 2 that the two load peaks may have been caused by the opening of two active rock joints which bolt A traversed. In bolt B there was only one load peak. It was also caused by the opening of a rock
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1.5 days 1.0
2.0
4 2
0.3
0
Peak load 98 days
30 20
1.5 days
10
0.7 days
0
Load in bolt (kN)
40
FIG 5 - In situ measurement of the rock movement and the axial load in a grouted bolt in the Four Fathom mudstone formation. The symbols are the measurement points and the curves are the best fitting lines. After Freeman (1978).
100 kN
= Yield load, 115 kN 1m Scale 100 kN
A B 8 months after installation 14 months after 17 months after 20 months after
Axial stress on steel bar (MPa)
99.1 days
Rock movements (mm)
THE LOAD-BEARING PROCESS OF OFUKLLY COUPLED ROCK BOLTS
15
10
Decoupled section
3 2
5
1 0 0
5
10
15
20
Distance to borehole collar (cm)
25
FIG 7 - The axial load in a fully grouted bolt at different levels of applied load. After Hawkes and Evans (1951).
say that curve 1 corresponds to the state of a compatible deformation across the interface in the entire length of the bar; curve 2 to the state of a partial decoupling of the interface in a section of the bar close to the loading point; and curve 3 to the state of a complete decoupling of the interface in a section close to the loading point.
CONCLUSIONS In continuously deformed rock formations, fully coupled rock bolts have one axial load peak. The position of the load peak moves towards the far end of the bolts with increasing rock deformation. In blocky rock masses several load peaks may appear in the rock bolt, depending on how many of the joints that the bolt intersects are actuated to open. The position of this kind of load peak is fixed at the actively opening joint, no matter how the magnitude of the peak load changes. In high stress/displacement environments associated with mass mining methods, bolts can be heavily loaded in a short period of time due to rapid changes in stresses and displacement in the rock mass. In blocky rock masses, the load on the face plate of a bolt can not be always treated as an indicator of the loading condition of the bolt. A bolt with a little-loaded face plate may be heavily loaded in the inside of rock due to joint opening. For rock bolts in pullout tests, decoupling of the interface always occurs first at the loading point. The decoupling front moves towards the far end of the bolt with an increase in the applied pull load.
REFERENCES FIG 6 - Development of the axial load in grouted bolts in a blocky rock mass. After Björnfot and Stephansson (1984).
joint, since the position of the load peak did not change with time.
Bolts subjected to pull loads Hawkes and Evans (1951) carried out a series of pullout tests using steel bars. The results for one of the bars are presented in Figure 7. It is seen that curves 1, 2 and 3 in this figure are similar to curves a, b and c in Figure 4a, respectively. Therefore we can
MassMin 2000
Björnfot, F and Stephansson, O, 1984. Interaction of grouted rock bolts and hard rock masses at variable loading in a test drift of the Kiirunavaara Mine, Sweden in Proc of Int Symp on Rock Bolting (Ed: O Stephansson) pp377-395 (Balkema: Rotterdam). Freeman, T J, 1978. The behaviour of fully-bonded rock bolts in the Kielder experimental tunnel, Tunnels & Tunnelling, June:37-40. Hawkes, J M and Evans, R H, 1951. Bond stresses in reinforced concrete columns and beams, J of the Institute of Structural Engineers, Vol XXIX, No X: 323-327. Hyett, A J, Mossavi, M and Bawden, W F, 1996. Load distribution along fully grouted bolts, with emphasis on cable bolt reinforcement, Int J for Numerical and Analytical Methods in Geomechanics, 20:517-544. Li, C and Stillborg, B, 1999. Analytical models for rock bolts, Int J Rock Mech Min Sci, 36: 1013-1029.
Brisbane, Qld, 29 October - 2 November 2000
935
Publications Published by The AusIMM
MONOGRAPH SERIES 1
• Detrital Heavy Minerals in Natural Accumulates
George Baker
1962
2.
• Research in Chemical and Extraction Metallurgy
Ed: J T Woodcock, A E Jenkins and G M Willis
1967
3.
• Broken Hill Mines - 1968
Ed: M Kadmanovich and J T Woodcock
1968
4.
• Economic Geology of New Zealand
Ed: G J Williams
1974
5.
• Economic Geology of Australia and Papua New Guinea - 1 Metals
Ed: C L Knight
1975
6
• Economic Geology of Australia and Papua New Guinea - 2 Coal
Ed: D M Traves and D King
1975
7.
• Economic Geology of Australia and Papua New Guinea - 3 Petroleum
Ed: R B Leslie H J Evans and C L Knight
1976
8.
• Economic Geology of Australia and Papua New Guinea 4 Industrial Minerals and Rocks
Ed: C L Knight
1976
9.
Field Geologists’ Manual 1st Edition 2nd Edition 3rd Edition
Ed: D A Berkman and W Ryall Ed: D A Berkman
1976 1982 1989
10
Mining and Metallurgical Practices in Australasia (the Sir Maurice Mawby Memorial Volume)
Ed: J T Woodcock
1980
11.
• Victoria’s Brown Coal - A Huge Fortune in Chancery (the Sir Willis Connolly Memorial Volume)
Ed: J T Woodcock
1984
12.
Australasian Coal Mining Practice 1st Edition 2nd Edition
Ed: C H Martin Ed: C H Martin and A J Hargraves
1986 1993
13.
Mineral Deposits of New Zealand
Ed: Dr D Kear
1989
14.
Geology of the Mineral Deposits of Australia and Papua New Guinea
Ed: F E Hughes
1990
15.
The Rocks Speak
H King
1989
16.
• Hidden Gold - The Central Norseman Story
J D Campbell
1990
17.
Geological Aspects of the Discovery of Some Important Mineral Deposits in Australia
K R Glasson and J H Rattigan
1990
18.
• Down Under - Mineral Heritage in Australasia
Sir Arvi Parbo
1992
19.
Australasian Mining and Metallurgy (the Sir Maurice Mawby Memorial Volume)
Ed: J T Woodcock and K Hamilton
1993
20.
Cost Estimation Handbook for the Australian Mining Industry
Ed: M Noakes and T Lanz
1993
21.
History of Coal Mining in Australia (the Con Martin Memorial Volume)
Ed: A J Hargraves, R J Kininmonth, C H Martin and S M C Saywell
1993
22.
Geology of Australian and Papua New Guinean Mineral Deposits
Ed: D Berkman and D Mackenzie
1998
Copies of all books currently in print can be obtained from The Institute office - Tel (03) 9662 3166 Key:
•
Out of print
PUBLICATIONS OF THE AUSTRALASIAN INSTITUTE OF MINING AND METALLURGY CONFERENCE, SYMPOSIUM AND MISCELLANEOUS 1972
1/72
SI
2/72 1973
*
Subsidence, Illawarra
S3
*
Mine Filling, North West Queensland
3/73
S4
*
Transportation, Sydney
4/73
S8
*
Mine Fires, Southern Queensland
*
Annual Conference, Western Australia
1/74
S5
*
Support in Pillar Extraction, Illawarra
2/74
S6
*
Recent Technical and Social Advances in the North Australian Minerals Industry, North West Queensland
3/74
S7
*
Pellets and Granules, Newcastle and District
*
Annual Conference, Southern and Central Queensland
1/75
S9
*
People and the Mining Industry - The Future, Broken Hill
2/75
S 10
*
Occupational Safety in Mines, Southern Queensland
3/75
S 11
*
Australian Black Coal, lllawarra
*
Annual Conference, South Australia
*
Landscaping and Land Use Planning as Related to Mining Operations, Adelaide
*
Design and Construction of Tunnels and Shafts, Melbourne - 2nd Australian Tunnelling Conference
1/76
S13
2/76 3/76
S 14
*
Thick Seam Mining by Underground Methods, Central Queensland
4/76
S 15
*
Sampling Practices in the Mineral Industries, Melbourne
*
AnnualConference, lllawara
5/76 1977
1 /77
S 16
*
Apcom ‘77, Brisbane
2/77
S 18
*
Coal Borehole Evaluation, Southern Queensland
3/77
S 17
*
Underground Operators’ Conference, Broken Hill
*
Annual Conference, Hobart
4/77 1978
1979
1 /78
S 19
*
Mill Operators’ Conference, North West Queensland
2/78
S20
*
Rock Breaking Equipment and Techniques, Melbourne
3/78
*
International Resource Management, Canberra
4/78
*
Annual Conference, Townsville
1/79
S21
*
Utilisation of Steelplant Slags, lllawarra
2/79
S22
*
Estimation and Statement of Mineral Reserves, Sydney
*
Annual Conference, Perth
3/79 1980
Annual Conference, Newcastle
S2
4/75 1976
*
2/73
4/74 1975
Project Evaluation and Management, Melbourne
1/73
5/73 1974
*
1/80
*
Annual Conference, New Zealand
2/80
S23
*
Australia/Japan Extractive Metallurgy Symposium, Sydney
3/80
S/24
*
Occurrence, Prediction and Control of Outbursts in Coal Mines, Southern Queensland
4/80
S25
*
Management in the Mining Industry, Melbourne
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662 * = Out of print The ‘S’ numbers in the third column refer to an older identifying number for Symposia, the numbers preceeding the ‘S’ number signify the new publication ordering number.
1981
I/81
S26
2/81 3/81
S27
4/81 1982
Annual Conference, Sydney Strip Mining 45 Metres and Beyond, Central Queensland
1/82
S29
*
Off Highway Truck Haulage Conference, Newman
2/82
S30
*
Mill Operators’ Conference, North West Queensland
3/82
S31
*
Underground Operators’ Conference, West Coast Tasmania
*
Annual Conference, Melbourne
5/82
S32
*
Carbon-ln-Pulp Technology for the Extraction of Gold, Perth and Kalgoorlie, (Reprinted 1988)
6/82
S33
*
Seam Gas Drainage with Particular Reference to the Working Seam, Wollongong
1/83
S34
*
Computers in Mining, Southern Queensland
*
Annual Conference, Broken Hill
3/83
S35
*
Project Development Symposium, Sydney
4/83
S37
*
Ventilation of Coal Mines, Wollongong
5/83
S40
*
Principles of Mineral Flotation (The Wark Symposium), Adelaide
1/84
S36
*
Metallurgy Symposium, Melbourne
2/84
S38
*
Coal and Mineral Sizing, Wollongong
*
Annual Conference, Darwin
4/84
S39
1/85
S41
*
Smelting and Refining Operators’ Symposium, North Queensland
2/85
S42
*
Underground Operators’ Conference, Kalgoorlie
*
Annual Conference, Brisbane
*
Scientific and Technological Developments in Extractive Metallurgy (G K Williams Memorial Volume), Melbourne
*
l3th Congress The Council of Mining and Metallurgical Institutions, Singapore, 6 Volumes
4/85
1987
Ignitions, Explosions and Fires, Wollongong
*
3/85
1986
Fourth Australian Tunnelling Conference
* *
3/84 1985
*
S28
2/83
1984
International Blast Furnace Hearth and Raceway Symposium, Newcastle
5/81
4/82
1983
*
S43
1/86
Gold Mining, Metallurgy and Geology, Kalgoorlie
2/86
S44
*
Selective, Open Pit Gold Mining Seminar, Perth
3/86
S45
*
Ground Movement and Control Related to Coal Mining, Wollongong
4/86
S46
*
Australia: A World Source of Illmenite, Rutile, Monazite and Zircon Conference, Perth
5/86
S47
*
Second Project Development Symposium, Sydney
6/86
S48
*
Large Open Pit Mining Conference, Newman
7/86
S49
*
Education and Research for the Mineral Industry for the Future, Melbourne
8/86
*
The AuslMM 10 Year lndex
1/87
*
Vl Australian Tunnelling Conference: Bore or Blast, Melbourne
*
Risk and Survival Seminar, Canberra
*
Annual Conference, Newcastle: Coal Power ‘87
*
Research and Development in Extractive Metallurgy, Adelaide
*
Leslie Bradford Golden Jubilee Oration
S52
*
Mining and Environment: A Professional Approach, Brisbane
*
Pacrim ‘87, Gold Coast, Queensland
S53
*
Dense Medium Operators’ Conference, Brisbane
2/87
S50
3/87 4/87
S51
5/87 6/87 7/87 8/87
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662 * = Out of print The ‘S’ numbers in the third column refer to an older identifying number for Symposia, the numbers preceeding the ‘S’ number signify the new publication ordering number.
9/87
S54
*
Equipment in the Minerals Industry: Exploration Mining and Processing Conference, Kalgoorlie
10/87
S55
*
Resources and Reserves, Sydney
*
South Australia’s Mining Heritage
*
21st Century Higher Production Coal Mining Systems Symposium, Wollongong
*
The Second International Conference on Prospecting in Arid Terrain, Perth
11/87 1988
1/88
S56
2/88 3/88
S57
*
Third Mill Operators’ Conference, Cobar
4/88
S58
*
Underground Operators’ Conference, Mount Isa
5/88
*
Fourth International Mine Ventilation Congress, Brisbane, (Proceedings and Addendum volume)
6/88
*
Annual Conference, Sydney: Minerals and Exploration at the Crossroads: The International Outreach
7/88
S59
*
Second AuslMM Mineral Heritage Seminar, Sydney
8/88
S60
*
Economics and Practice of Heap Leaching in Gold Mining Workshop, Cairns
*
Third International Mine Water Congress, Melbourne
9/88 10/88 1989
*
Explosives in Mining Workshop, Melbourne
1/89
S61
*
Mineralogy and Petrology, Sydney, February
2/89
*
Second Large Open Pit Mining Conference, Latrobe Valley Vic
3/89
*
NQ Gold ‘89 Conference, Townsville Qld
4/89
*
Annual Conference, Perth-Kalgoorlie: Education, Training and Professional Development; Industrial Minerals; Project Development/Processing
5/89
*
Mineral Fuel Alternatives and the Greenhouse Effect, July 1989
6/89 7/89
Non-ferrous Smelting Symposium: 100 Years of Smelting and Refining Operations in Port Pirie, SA September 1989 *
89 1990
1991
1 /90
Dewatering Technology and Practice Conference, Brisbane October 1989 MINVAL ‘89, Mining and Petroleum Valuation 1989, Sydney September 1989
*
Ore Reserve Estimates - The Impact on Miners and Financiers, Melbourne, March 1990
2/90
Annual Conference, The Mineral Industry in New Zealand, Rotorua New Zealand, March 1990
3/90
Pacific Rim Congress, Gold Coast Qld, May 1990
4/90
*
Mining Industry Capital and Operating Cost Estimation Conference, Sydney, June 1990
5/90
*
Third International Symposium on Rock Fragmentation by Blasting, Brisbane, August 1990
6/90
*
Sir Edgeworth David Memorial Oration, May 1990
7/90
*
Mine Geologists’ Conference, Mount Isa, October 1990
1/91
*
Fourth Mill Operators’ Conference, Burnie Tas, March 1991
2/91
*
World Gold ‘91, Cairns Qld, April 1991
3/91 4/91
Mining Industry Optimisation Conference, Sydney, June l991 *
5/91
PNG Geology, Exploration and Mining Conference, Rabaul, June 1991 Qld Coal Symposium, Brisbane, August 1991
6/91
*
Reliability Production and Control in Coal Mines, Wollongong, September 1991
7/91
*
Fifth AuslMM Extractive Metallurgy Conference, Perth, October 1991
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662 * = Out of print The ‘S’ numbers in the third column refer to an older identifying number for Symposia, the numbers preceeding the ‘S’ number signify the new publication ordering number.
1992
1/92
Enviromine Australia, Sydney NSW, March 1992
2/92
The AuslMM Annual Conference, ‘The State-of-the-Art - A Product of 100 Years of Learning’, Broken Hill NSW, May 1992
3/92
‘Energy, Economics and Environment’ Gippsland Basin Symposium, Melbourne, June 1992
4/92
Arnold Black Mineral Heritage Oration
5/92
The Man from ASARCO: a life and times of Julius Kruttschnitt
6/92
5th Underground Operators’ Conference, Ballarat, July 1992
7/92
*
11th International Conference on Ground Control in Mining, Wollongong, July 1992
8/92
*
Third Large Open Pit Mining Conference, Mackay, August 1992
9/92
Extractive Metallurgy of Gold and Base Metals Conference, Kalgoorlie, October 1992
10/92 11/92 1993
1994
1995
Sampling Practices in the Minerals Industry, Mount Isa, November 1992 *
Rehabilitate Victoria, Latrobe Valley, November 1992
1/93
Mining People - A Century
2/93
The AuslMM Centenary Conference, Adelaide, March 1993
3/93
XVIII International Mineral Processing Congress, Sydney, May 1993
4/93
Narrow Vein Mining Seminar, Bendigo, June 1993
5/93
International Mining Geology Conference, Kalgoorlie, July 1993
6/93
Vlll Australian Tunnelling Conference, Sydney, August 1993
7/93
World Zinc ‘93 - International Symposium, Hobart, October 1993
1/94
1994 AuslMM Student Conference, Brisbane, April 1994
2/94
PNG Geology, Exploration and Mining Conference, Lae, PNG, lune 1994
3/94
No Two The Same by Bert Mason
4/94
Sixth Extractive Metallurgy Conference, Brisbane July 1994
5/94
1994 AuslMM Annual Conference, Darwin, August 1994
6/94
4th Large Open Pit Mining Conference, Perth, September 1994
7/94
Recent Trends in Heap Leaching, Bendigo, September 1994
8/94
Maintenance in the Mining and Metallurgical Industries,Wollongong, October 1994
9/94
Fifth Mill Operators’ Conference, Roxby Downs, October 1994
10/94
Mineral Valuation Methodologies 1994, Sydney, October 1994
11/94
Victorian Mining Week Conference, Melbourne, October 1994
1/95
1995 AuslMM Annual Conference, Newcastle, March 1995
2/95
Sir Maurice Mawby Memorial Oration
3/95
World’s Best Practice in Mining and Processing Conference, Sydney, May 1995
4/95
APCOM XXV 1995 Conference, Brisbane, July 1995
5/95
Mineral Valuation Methodologies 1994, Sydney, October 1994 (revised)
6/95
EXPLO 95 Conference, Brisbane, September 1995
7/95
Underground Operators’ Conference, Kalgoorlie, November 1995
8/95
Young Professionals’ Conference, Mt Isa, October 1995
9/95 10/95
*
PACRIM ‘95 Congress, Auckland, New Zealand, November 1995 Ethics, Liability and the Technical Expert, Sydney, December 1995
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662 * = Out of print The ‘S’ numbers in the third column refer to an older identifying number for Symposia, the numbers preceeding the ‘S’ number signify the new publication ordering number.
1996
1/96
*
1996 AusIMM Annual Conference, Perth, March 1996
1a/96
*
1996 AusIMM Annual Conference Supplementary Volume, Perth, March 1996
2/96
Ethics, Liability and the Technical Expert, Sydney, March 1996
3/96
Entrepreneurs and Partners, Sydney, July 1996
4/96
Contract Operators’ Conference, Kalgoorlie, October 1996
5/96
Asia/Pacific Mining Communications Summit, Singapore, November 1996 Withdrawn
6/96 1997
1998
*
1/97
1997 AusIMM Annual Conference, Ballarat, March 1997
2/97
World Gold ‘97 Conference, Singapore, September 1997
3/97
Sixth Mill Operators’ Conference, Madang, PNG, October 1997
4/97
Gem 97, Madang, PNG, October 1997
5/97
Contract Operators’ Conference, Brisbane, Qld, October 1997
6/97
Third International Mining Geology Conference, Launceston, Tas, November 1997
7/97
Mindev 97 - The International Conference on Mine Project Development, Sydney, November 1997
8/97
1997 AusIMM Travelling Technology Forum, Singleton, NSW, March 1997
1/98
MINEFILL ‘98 - The Sixth International Symposium on Mining with Backfill, Brisbane, Qld, April 1998
2/98
1999
2000
Nickel ‘96, Kalgoorlie, November 1996
*
AusIMM’98 - The Mining Cycle, Mount Isa, Qld, April 1998
3/98
Seventh Underground Operators’ Conference, Townsville, Qld, June/July 1998
4/98
Mine to Mill Conference, Brisbane, Qld, October 1998
5/98
Third Regional APCOM - Computer Applications in the Minerals Industries International Symposium, Kalgoorlie, WA, December 1998
1/99
10th Australian Tunnelling Conference, Melbourne, Vic, March 1999
1a/99
10th Australian Tunnelling Conference Keynote Addresses and Asia–Pacific Forum, Melbourne, Vic, March 1999
2/99
Students and Young Professionals Conference, Perth, WA, July 1999
3/99
ICARISM ’99 Conference, Perth, WA, September 1999
4/99
PACRIM ’99 Congress, Bali, October 1999
5/99
Explo ’99 Conference, Kalgoorlie, WA, November 1999
1/2000
Southern Africa - Australia Mineral Sector Synergies Symposium, Canberra, ACT, March 2000
2/2000
After 2000 - The Future of Mining, Sydney, NSW, April 2000
3/2000
4th International Mining Geology Conference, Coolum, Qld, May 2000
4/2000
Young Leaders 2000, Sydney, NSW, July 2000
5/2000
MINPREX 2000, Melbourne, Vic, September 2000
6/2000
Seventh Mill Operators’ Conference, Kalgoorlie, WA, October 2000
7/2000
MassMin 2000, Brisbane, Qld, October - November 2000
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662 * = Out of print The ‘S’ numbers in the third column refer to an older identifying number for Symposia, the numbers preceeding the ‘S’ number signify the new publication ordering number.
SPECTRUM SERIES 1.
Making the Mount Isa Mine, 1923 - 1933
Don Berkman
1996
2.
History of Drilling
Graham McGogggan
1996
3.
The Cobar Mineral Field - A 1996 Perspective
Warren Cook Andrew Ford Julian McDermott Peter Standish Craig Stegman and Therese Stegman
1996
4.
Towards 2000 - Resource to Reserve Inputs Seminar - Melbourne, Vic
1997
5.
Towards 2000 - National Conference on Ironmaking Resources and Reserves Estimation, Perth, WA
1997
6.
Towards 2000 - The Resource Database Towards 2000 - Wollongong, NSW
1997
7.
Towards 2000 - Ore Reserves and Finance - Sydney, NSW
1998
8.
Towards 2000 - Assessment of Reserves in Low Rank Coals - Morwell, Vic
1997
9.
Towards 2000 - Ore Reserve Reconciliation Workshop - Darwin, NT
1997
10.
Towards 2000 - Gold and Nickel Ore Reserve Estimation Practice Seminar
1997
11.
Towards 2000 - Resource/Reserves Estimation Practice in the Central West New South Wales Mining Industry, Cobar, NSW
1998
Copies of all books currently in print can be obtained from The Institute office Tel (03) 9662 3166 or Fax (03) 9662 3662