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Sealing of Boreholes and Underground Excavations in Rock

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Sealing of Boreholes and Underground Excavations in Rock Edited by

K. Fuenkajorn Rock Engineering International, Tucson, USA

and

J. J. K. Daemen Mining Engineering Department, University of Nevada, USA

CHAPMAN & HALL London· Weinheim . New York· Tokyo Melbourne· Madras

Published by Chapman & Hall, 2-6 Boundary Row, London SEt 8HN, UK Chapman & Hall, 2-6 Boundary Row, London SEI 8HN, UK Chapman & Hall GmbH, Pappelallee 3, 69469 Weinheim, Germany Chapman & Hall USA, 115 Fifth Avenue, New York, NY 10003, USA Chapman & Hall Japan, ITP-Japan, Kyowa Building, 3F, 2-2-1 Hirakawacho, Chiyoda-ku, Tokyo 102, Japan Chapman & Hall Australia, 102 Dodds Street, South Melbourne, Victoria 3205, Australia Chapman & Hall India, R. Seshadri, 32 Second Main Road, CIT East, Madras 600 035, India First edition 1996

© 1996 Chapman & Hall Softcover reprint of the hardcover 1st edition 1996 Typeset in lOj12pt Times by Thomson Press (I) Ltd., Madras ISBN-l3: 978-94-010-7173-4 001: 10.1007/978-94-009-1505-3

e-ISBN-13: 978-94-009-1505-3

Apart from any fair dealing for the purposes of research or private study, or criticism or review, as permitted under the UK Copyright Designs and Patents Act, 1988, this publication may not be reproduced, stored, or transmitted, in any form or by any means, without the prior permission in writing of the publishers, or in the case of reprographic reproduction only in accordance with the terms of the licences issued by the Copyright Licensing Agency in the UK, or in accordance with the terms of licences issued by the appropriate Reproduction Rights Organization outside the UK. Enquiries concerning reproduction outside the terms stated here should be sent to the publishers at the London address printed on this page. The publisher makes no representation, express or implied, with regard to the accuracy of the information contained in this book and cannot accept any legal responsibility or liability for any errors or omissions that may be made. A catalogue record for this book is available from the British Library Library of Congress Catalog Card Number: 95-71862

~

Printed on acid-free text paper, manufactured in accordance with ANSljNISO Z39.48-1992 (Permanence of Paper).

Contents

List of Contributors

ix

Preface

xi

1. Introduction . . . . . . . . . . . . . . . . J. J. K. Daemen 1.1 Background.............. 1.2 Sealing Requirements-Rules and Regulations . 1.3 Current Practice . . . . . . . . . . . . . . . . . 1.4 Recent Borehole Sealing Research-An Introductory Overview . . . . . . . . . . . . . . . . . 1.5 Sealing of Shafts, Tunnels, Mine Adits, Portals and Drifts . 1.6 Book Summary 1. 7 The Future . . Acknowledgements . 2. Laboratory Performance of Cement Borehole Seals . D. L. South and K. Fuenkajorn 2.1 Introduction . . . . . . . 2.2 Conceptual Approach . 2.3 Experimental Apparatus . 2.4 Experimental Procedures 2.5 Analysis . . . . . . . . . 2.6 Summary of Test Results . 2.7 Design Implications Acknowledgements . . . . . . . 3. Strength Parameters of Cement Borehole Seals in Rock . H. Akgiln 3.1 Introduction . . . . . . . . . . . . . . . 3.2 Push-Out Test Experimental Procedure. 3.3 Push-Out Test Mechanical Interactions. 3.4 Finite Element Analysis and Discussion.

1 1 1 2 2 4 4 7 8 9 9 10 11 13 15 19 26 27

28 28

29 29 31

Contents

VI

3.5 Push-Out Test Results and Discussion . 3.6 Conclusions and Recommendations Acknowledgements . . . . . . . . . . . . . . 4. Dynamic Loading Impact on Cement Borehole Seals . G. S. Adisoma 4.1 Introduction...... 4.2 Experimental Methods 4.3 Flow Test Results . . . 4.4 Dye Injection Test Results 4.5 Dynamic Loading Test Results . 4.6 Design Considerations. 4.7 Discussion .. 4.8 Conclusions . Acknowledgements 5. Performance of Bentonite and Bentonite/Crushed Rock Borehole Seals . S. Ouyang and J. J. K. Daemen 5.1 Introduction . . . . . . . . . . . 5.2 Materials............. 5.3 Permeability Tests and Results. 5.4 Piping and Flow of Bentonite. . 5.5 Prediction of Bentonite Permeability . 5.6 Conclusions and Recommendations Acknowledgements . . . . . . . . . . . . .

6. In situ Performance of a Clay-Based Barrier. B. H. Kjartanson, N. A. Chandler and A. W L. Wan 6.1 6.2

Introduction............ The Canadian Nuclear Fuel Waste Management Program . . . . . . 6.3 Evaluation of Buffer Physical Performance. 6.4 In situ Experiments . . . . . 6.5 Buffer/Container Experiment 6.6 Isothermal Test 6.7 Conclusions . Acknowledgements .

36 39 39 40 40 42 44 50 52 56 60 62 64

65 65 66 70

82 86 93 95

96 96 97 99 100 102 117 123 125

7. In situ Hydraulic Performance Tests of Borehole Seals: Procedure and Analyses . W. B. Greer 7.1 Introduction....... 7.2 Proposed Test Schemes. 7.3 Seal Tests . . . . . . . .

126 126 127

131

Contents 7.4 Conclusions.. Acknowledgements.

8. In situ Hydraulic Performance of Cement Borehole Seals . W. B. Greer and D. R. Crouthamel 8.1 Introduction................ 8.2 Preliminary Survey for Seal Installation. 8.3 Seal Tests at Oracle Ridge Mine. . . . . 8.4 Discussion of Testing at Oracle Ridge Mine. 8.5 Seal Testing at Superior, Arizona. . . . . . . 8.6 Recommendations for Testing In situ Borehole Seals. Acknowledgements . . . . . . . 9. Sealing Boreholes in Rock Salt . J. C. Stormont and R. E. Finley 9.1 Introduction........ 9.2 General Practice in Sealing Boreholes in Rock Salt. 9.3 Rock Salt Properties Relevant to Borehole Sealing. 9.4 Materials for Sealing Boreholes in Rock Salt. 9.5 Seal System Performance . 9.6 Design Considerations 9.7 Conclusions . Acknowledgement . . . . . 10. Design of Underground Plugs. F. A. Auld 10.1 Introduction . . . . . . . 10.2 Types of Plugs . . . . . . 10.3 Factors to be Considered in Design of Plugs. 10.4 Design Calculations . . . . . . . . . . . . 10.5 Construction Aspects . . . . . . . . . . . 10.6 Plug Sealing and Resistance to Leakage. 10.7 Case Studies . . . . . . . . . . . . . 10.8 Conclusions and Recommendations . 11. Design of Borehole Seals: Process, Criteria and Considerations . . . . . . . . . J. J. K. Daemen and K. Fuenkajorn 11.1 Introduction . . . . . . . . . 11.2 Applications, Objectives and Requirements. 11.3 Considerations of Site Characteristics. . . . 11.4 General Design Criteria. . . . . . . . . . . 11.5 Material Selection and Placement Methods. 11.6 Summary.. Acknowledgements . . . . . . . . . . . . . . . . . .

vii 152 156 157 157 157 158 175 178 181 183 184 184 186 190 201 211 218 223 224 225 225 226 227 233 246 249 254 264 267 267 268 270 274 276 279 279

Contents

Vlll

12. Sodium Bentonite as a Borehole Sealant. 1. E. Papp

12.1 12.2 12.3 12.4 12.5 12.6 12.7 12.8 12.9 12.10

280

Introduction . . . .. Geographical Origin. Geological Origin . . Structure of Sodium Bentonite. Properties of Sodium Bentonite. Bentonite as a Borehole Sealant. Bentonite as an Annular Sealant. Types of Bentonite Sealants. . . . Placement of Bentonite . . . . . . Case History Featuring High-Solids Bentonite Grout 12.11 Conclusions

280 282 282 282 284 285 286 287

Bibliography .

298

Index

325

....

292

294 297

Contributors

G. S. Adisoma Independent Mining Consultants, Inc., 2700 East Executive Drive, Suite 140, Tucson, Arizona 85706, USA H. Akgful Department of Geological Engineering, Orta Dogu Teknik Universitesi, Middle East Technical University, Ankara, Turkey F. A. Auld I. W. Farmer & Partners, Ltd, Jilland House, 329 Bawtry Road, Doncaster, DN4 7PB, England N. A. Chandler AECL Research, Whites hell Laboratories, Pinawa, Manitoba, Canada ROE 1LO D. R. Crouthamel Stone & Webster Engineering Corporation, 245 Summer Street, Boston, Massachusetts 02210, USA J. J. K. Daemen Mining Engineering Department, Mackay School of Mines, University of Nevada, Reno, Nevada 89557-0139, USA R. E. Finley Sandia National Laboratories, Geotechnical Investigations Department, MS-1325, P.O. Box 5800, Albuquerque, New Mexico 87185, USA K. Fuenkajorn Rock Engineering International, 7226 West Rivulet Drive, Tucson, Arizona 85743, USA

W. B. Greer U. S. Bureau of Reclamation, Water and Power Resources Management Division, 2800 Cottage Way, Sacramento, California 95825, USA B. H. Kjartanson Department of Civil and Construction Engineering, Iowa State University, Ames, Iowa 50011-3232, USA

x

Contributors

S.Ouyang Industrial Technology Research Institute, Energy and Resources Laboratories, Building #24, 195-6 Chung Hsing Road, Section 4, Chutung. Hsinchu, Taiwan 310, R.O.C. J. E. Papp Colloid Environmental Technologies Company, 1350 West Shure Drive, Arlington Heights, Illinois 60004-7803, USA D. L. South Washington State Department of Ecology, Northwest Regional Office, State of Washington, 3190-160th Avenue, S. E., Bellevue, Washington 98008-5452, USA J. C. Stormont Department of Civil Engineering, University of New Mexico, Albuquerque, New Mexico 87131, USA A. W.L. Wan AECL Research, Whiteshell Laboratories, Pinawa, Manitoba, Canada ROE 1LO

Preface

Sealing of boreholes and underground excavations has not received much engineering attention until fairly recently. The growing awareness of and sensitivity to environmental concerns of the technical community as well as of the public at large has resulted in an increasing recognition of the fact that these geological penetrations may have an environmental impact. The issue of possible contamination resulting from migration along boreholes, adits, shafts or tunnels unquestionably has been raised most forcefully within the context of nuclear waste disposal. Several nuclear waste disposal programs, notably the Civilian and the Defence programs of the US Department of Energy, the US Nuclear Regulatory Commission and the Canadian and Swedish radioactive waste disposal programs have conducted major research efforts aimed at developing adequate seal designs for penetrations in host rock formations for high-level nuclear waste repositories. While a considerable data base has been gathered over the last two decades or so with regard to the performance of seals, most of the information is presented in research reports and widely scattered papers in journals and proceedings of conferences. Hence, the materials are not readily accessible to potential users such as designers, contractors or regulators who are not familiar with nuclear waste disposal programs. Although many government agencies have implemented regulations requiring that unused boreholes and underground excavations in rock formations be sealed, these regulations tend to be generic and broad, and rarely allow for taking into account site-specific conditions. As a result, it is probable that, for example, they are excessively conservative for some locations and inadequate for others. We organize and structure the available information on sealing boreholes and underground excavations in a format that makes it much more readily accessible. This book presents a comprehensive integrated summary of recent laboratory and in situ experiments conducted to assess the mechanical and hydraulic performance of the emplaced seals and host rocks. The contents are structured so as to highlight design considerations and recommendations under various aspects of seal and host rock characteristics, and installation environments. Uncertainties in terms of design, performance testing and performance predictions are recognized, and lead naturally to a discussion of remaining research needs. This book is immediately applicable as a compiled literature summary to senior college students and research graduates. It is intended to be a

xii

Preface

reference source for design professionals and researchers in many disciplines of engineering and sciences, including mining, geological, petroleum, nuclear, civil, sanitary and environmental engineers, as well as geologists, hydrologists and material scientists. The present testing techniques, approaches and recommendations will assist graduates and experimentalists in designing their test schemes to develop new sealants and installation methods for a specific application or to answer remaining questions on the sealing of underground space. The data and design considerations presented will allow engineers to plan and design seals with improved confidence. K. Fuenkajorn J. J. K. Daemen

CHAPTER ONE

Introduction J. J. K. Daemen

1.1

BACKGROUND

Sealing of underground openings has assumed a growing importance with the recognition that penetrations of geological formations can have a detrimental impact on the environment. Boreholes that penetrate aquitards may allow migration and mixing of groundwaters of different qualities, and may contaminate aquifers. Poorly sealed or open boreholes may allow premature and unnecessary depressurization of formations, and may result in wasting of natural resources. Inadequately sealed mine adits may allow unacceptable discharges of acid mine drainage. This book presents an overview of sealing methods that can be used to prevent or minimize the deterimental effects that may result from leaving geological pentrations open. The emphasis of this book is to present results of research investigations that address the difficult question of what level of sealing performance can be expected from various sealing methods. The work presented is oriented predominantly towards the sealing of boreholes in rock. Implications for seal design are given.

1.2

SEALING REQUIREMENTS-RULES AND REGULATIONS

Because of the recognition that leaving boreholes unsealed can have a variety of detrimental impacts, rules and regulations that govern the procedures to be followed to close boreholes have been implemented for many years in many jurisdictions. Among the earliest regulatory requirements were those imposed by the Texas Railroad Commission on sealing of oil and gas wells. A prime objective of these regulations was to prevent premature depressurization of oil and gas formations.

Sealing of Boreholes and Underground Excavations in Rock. Edited by K. Fuenkajorn and J. J. K. Daemen. Published in 1996 by Chapman & Hall. ISBN 0 412 57300 8.

2

Introduction

In recent years many countries and states have introduced or strengthened regulatory requirements for borehole sealing. An exhaustive set of US regulations and related documents has recently been complied and published by Baroid Drilling Fluids, Inc. and Environmental Planning Group, Inc. (1995), who also will provide annual updates of the regulations. The availability of this documentation should facilitate greatly the difficult task of trying to remain updated about the regulatory requirements. Regulatory requirements historically have been dominated by very practical prescriptions in terms of how borehole sealing is to be accomplished. For reasons that will become very clear later in this book, the obviously more attractive approach of making the regulations performance oriented or based remains exceedingly difficult to implement, because testing the performance of borehole seals remains difficult. 1.3

CURRENT PRACTICE

Current borehole sealing practice depends largely upon the industrial context within which boreholes are drilled. Within the petroleum industry the use of cementitious seals predominates. In the construction, geotechnical and water well industries clay, primarily bentonitic, seal materials share a substantial fraction of the market. While the casual 'sealing' of boreholes with telephone poles, beer cans and various types of debris may not yet have been abandoned completely, it certainly no longer is acceptable practice. Cementitious sealing of deep oil and gas wells requires sophisticated emplacement technology. The necessary equipment may be on site as part of the well completion equipment. Such technology rarely if ever is available for the sealing of geotechnical or ore exploration holes, or for water well sealing. Grout pumps may be available for the sealing of such holes. Careful downhole pumping of cement or bentonite grouts has been widely accepted as a borehole sealing practice. Bailer emplacement is an acceptable alternative for cement grouts. Bentonite chips can be dropped in holes containing standing water, and provide a low permeability seal. 1.4

RECENT BOREHOLE SEALING RESEARCH-AN INTRODUCTORY OVERVIEW

For nearly two decades borehole sealing research has received a major impetus, in several countries, within the context of sealing nuclear waste repositories. While the objectives of these programs may be rather unusual, these investigations have contributed substantially to the development of sealing technologies and to the determination of sealing performance that

Recent borehole sealing research

3

are of value in general sealing practice, i.e. in the far more numerous and mundane applications that dominate sealing requirements. Sandia National Laboratories (e.g. Hansen et al., 1995) has conducted and continues to conduct a major investigation of the sealing of penetrations, primarily boreholes and shafts, through rock salt and, more generally, evaporite formations. This work, some aspects of which are presented in Chapter 9, has included extensive laboratory and field experiments on the sealing performance of cementitious, salt and clay seals. The work can be considered as an extension and continuation of work initiated by Oak Ridge National Laboratory (e.g. McDaniel, 1980) and continued by the Office of Nuclear Waste Isolation (e.g. Roy, Grutzeck and Wakeley, 1985). Sealing of penetrations through salt and evaporites is particularly important because of the solubility of halite, and because of the numerous penetrations of such formations, whether for mining, waste disposal or oil and gas production purposes. Within the context of the Swedish nuclear waste disposal program Pusch and associates, e.g. Pusch (1983), have conducted extensive studies on the sealing performance of bentonite, especially of highly compacted bentonite. These studies have resulted in an exhaustive data basis on the sealing performance of highly compacted bentonite, and in a dramatically improved understanding of the performance characteristics of bentonite. The Office of Nuclear Waste Isolation has supported various borehole sealing studies, conducted, among the principal research groups, at the Materials Research Laboratory at Pennsylvania State University (e.g. Roy, Grutzeck and Wakeley, 1985) and at Waterways Experiment Station of the US Army Corps of Engineers (e.g. Buck, Burkes and Rhoderick, 1981). The main emphasis of this research has been on sealing of evaporite formations with cementitious materials. The research has continued for the Waste Isolation Pilot Plant (WIPP) (e.g. Wakeley, Harrington and Weiss, 1993). IT Corporation, formerly D' Appolonia Consulting engineers (e.g. D'Appolonia Consulting Engineers, 1981), also has participated extensively in the salt repository sealing studies (e.g. International Technology Corp., 1987). The Canadian nuclear waste disposal program has conducted investigations on borehole sealing, aspects of which are summarized in Chapter 6. The Department of Mining and Geological Engineering at the University of Arizona has been sponsored by the US Nuclear Regulatory Commission (NRC) from the late 1970s to the late 1980s to conduct investigations on borehole and shaft sealing. Several chapters in this book (2-5,7 and 8) are based on work sponsored by NRC. The Federal Highway Adiministration has initiated a research study on the sealing of exploratory boreholes, primarily geotechnical holes in soils, initial results of which are summarized by Lutenegger and DeGroot (1993).

4

Introduction 1.5

SEALING OF SHAFTS, TUNNELS, MINE ADITS, PORTALS AND DRIFTS

Sealing of excavations larger than boreholes, such as mine shafts or adits, or tunnels, may have a variety of purposes. The level of technology applied for particular cases tends to be highly variable, depending upon the objectives of the sealing operations. For many abandoned or to be abandoned mine entries the prime purpose of closure is to prevent human intrusion, with the ultimate objective of forestalling accidents that frequently result when people enter old mines. For many such mine closures the technology tends to be rather rudimentary, aimed at a minimal physical barrier that will satisfy the requirements to minimize the risk of penetrations. At the other extreme are situations where the objective of permanent closures is to provide a true seal, e.g. tight to both air and water. Historical examples of this type include primarily major case studies of mine flooding and of the design of plugs to allow controlling the flooding (e.g. Garrett and Campbell Pitt, 1961). Current examples tend to be oriented primarily towards reducing the environmental impact from mine discharges (e.g. Einarson and Abel, 1990). Design approaches for sealing large excavations, as well as recent examples, are given in Chapter 10. 1.6

BOOK SUMMARY

Chapter 2 presents the results of flow testing of cement grout seals emplaced in boreholes in rock. The study was intended primarily to look at the performance of seals in deep boreholes in strong intact rock. Flow tests were conducted in triaxial cells, with particular attention given to the influence of changing lateral stresses on the sealing performance. Detailed attention also was given in this study to the risk of interface flow, i.e. flow along the contact between rock and seal cement grout. It was found that for the materials tested, and under the conditions for which the tests were conducted, the interface did not usually form a preferential flow path. While drying of the cementitious seals tested here does induce shrinkage and does increase the hydraulic conductivity and hence allows increased flow, resaturation with water flow reverses the sealing performance noticeably. Chapter 2 includes a summary of the results of an extensive numerical study of the flow paths that can be expected around a borehole seal as a function of the hydraulic conductivity of the seal relative to that of the host rock. No significant reduction in flow through the seal system is obtained by reducing the hydraulic conductivity of the seal below about an order of magnitude less than that of the host rock. Because these results are summarized in dimensionless form, they are valid for seals of any diameter. They provide guidance for the selection of seal permeability goals, assuming that the permeability of the host rock is known or can be estimated.

Book summary

5

Chapter 3 briefly summarizes a study of bond strength of cement grout seals in rock. Bond strength usually is not an issue for cementitious borehole seals, whose length generally far exceeds the length necessary to provide adequate bond strength. Bond strength often is a critical design parameter for shaft and tunnel seals, at least for plugs at great depths that may be subject to high axial loading. Chapter 3 provides bond strength values determined experimentally on plugs emplaced in small-diameter boreholes, and guidance on how these results might be scaled up to larger diameter sealing plugs. Chapter 4 summarizes an investigation of dynamic effects on the performance of cement grout borehole seals. It is found that vibrations, even at considerable acceleration, do not significantly deteriorate sealing performance. Conversely, allowing cementitious borehole seals to dry can rapidly enhance their permeability and reduces their bond strength. Of considerable interest is the observation that these seals tend to recover significantly upon resaturation. The complex drying/resaturation sequence and effects deserve further study, because they almost certainly are significant with respect to the long-term performance of many seals. Chapter 5 presents the results of experimental and modeling investigations on the borehole sealing performance of bentonite and mixtures of bentonite and crushed welded tuff. The incentive for the investigation is the proposition that benefits may be derived under certain circumstances from loading bentonite with crushed rock. Such mixture sealants may maintain the attractive features of bentonite seals, notably their low permeability, self-healing, durability, swelling deformability and sorptive characteristics. The addition of crushed rock may have some potential benefits, e.g. decreasing the shrinkage potential during drying, increasing the strength or bearing capacity and increasing the achievable density. The influence of adding crushed rock on swelling potential remains rather uncertain. The investigation reported in Chapter 5 addresses primarily the questions of the influence of crushed tuff on hydraulic conductivity and of the conditions under which bentonite might be lost from a seal, e.g. as a result of bentonite flow into fractures in the host rock within which the seal is emplaced. The experimental investigations are conducted by means of permeability tests on reference samples of bentonite only, and mainly by permeability tests on mixtures of crushed tuff and bentonite. Bentonite weight percentages of 15, 25 and 35% have been tested. Several different gradations of crushed tuff size distributions have been tested. Piping tests are reported, aimed at determining the conditions under which bentonite may be lost from the seals. Modeling efforts consist of determining empirical correlations between various characteristics of the seals and the permeability or resistance to bentonite loss, as characterized by the yield stress. The authors conclude that crushed rock/bentonite mixtures preferably should contain at least 25% bentonite, in order to achieve permeabilities

6

Introduction

comparable to those of bentonite only. Compaction is a significant variable affecting sealing performance. The mixtures, as compacted here, show anisotropy and heterogeneity. Chapter 6 presents the results of extensive experimental investigations on the sealing performance of bentonite-sand-aggregate earthen barriers. The investigations were intended to identify an optimum sand/bentonite mix that would give the best overall combination of physical performance characteristics. The authors identify some serious difficulties with conventional quality control procedures of compaction, and for that reason implemented stringent controls on compaction procedures to be followed. The authors recognize and stress the need for large-scale field testing, even though many aspects of barrier performance can be studied adequately in laboratory testing. The experimental investigations have focused strongly on the consequences of moisture migration from, into and through the sandjbentonite barriers. The observations during the experiments confirm experience gained elsewhere (e.g. Gaudette and Daemen, 1988) that moisture migration through bentonitic barriers is an exceedingly complex phenomenon. Nevertheless, current observations on these continuing tests indicate that the barriers will provide a very low hydraulic conductivity. Chapter 7 presents methods to determine the in situ hydraulic performance of borehole seals. Experimental determination of the in situ hydraulic performance of borehole seals is difficult, at least when the seal performance is in the low hydraulic conductivity range desirable for sealing openings in intact low-permeability rock, because the conductivities are very low indeed. Depending on the hydraulic characteristics of the host rock, the flow paths may be influenced significantly by flow through the host rock. Chapter 7 outlines systematic experimental and analyses strategies for testing the performance of such low-conductivity seals. Chapter 8 describes field tests on cementitious seals emplaced in boreholes in rock. The test methods and interpretation procedures introduced in Chapter 7 are illustrated and applied here for several case studies. One of the case studies, with a cement grout seal in a nearly horizontal hole, demonstrated the satisfactory closure of an initial interface gap as a result of cement grout expansion. The other experiments, on seals in vertical holes, confirmed the excellent sealing performance of cementitious seals emplaced by relatively simple and conventional procedures. Chapter 9 deals with the sealing of boreholes in rock salt. Although a somewhat specialized geological environment, it is rather widespread. Evaporite aquitards are frequently penetrated by boreholes and shafts, and sealing failures in soluble evaporites can have and have had notoriously disastrous consequences. Some aspects of rock salt behavior, notably its potential for creep and for healing, are favorable with respect to sealing, while others, especially its solubility, are potentially detrimental. The authors stress the need to consider the entire seal system, i.e. including seal, seal-rock interface and host rock, when seal design and performance are

The future

7

considered. For that reason they include a discussion of rock salt behavior and characteristics. Materials discussed in Chapter 9 for the purpose of sealing boreholes in rock salt include granular rock salt, cement grouts, concrete and clays. Results of laboratory and field tests on the performance of various seals are presented. In their discussion of design implications the authors place considerable emphasis on the preference for truly permanent seals, a preference which may have major implications for the selection of sealing materials. Chapter 10 presents an approach to the design of plugs for shafts, ramps, drifts and tunnels. The chapter introduces a functional classification of the main types of plugs used in underground excavations, and identifies the main factors that need to be considered in the design of plugs. Detailed design calculations are included, as well as allowable stresses for rock, concrete and steel. Construction procedures for concrete plugs are given, with emphasis on the need for careful attention to many details. The author discusses procedures for reducing leakage parallel to plugs and gives several design and construction case studies. Chapter 11 outlines some design considerations for borehole seals. It is recognized that most regulatory borehole sealing requirements prescribe materials and emplacement procedures primarily. Chapter 11 gives a broad overview of the many factors that need to be accounted for when planning and designing borehole seals. Chapter 12 presents a discussion of the use of sodium bentonite for the sealing of the boreholes. The increasing popularity and acceptance of various forms of bentonite for the sealing of boreholes is based on the wide recognition of its desirable properties for sealing purposes: low hydraulic conductivity, geochemical compatibility in many natural environments, great capacity for self-sealing and ease of installation in a variety of forms. Bentonite, by itself or as an admixture, has been used extensively for sealing purposes in a variety of applications. In Chapter 12 some of the origin and geological aspects of bentonite are introduced. The use of bentonite as a borehole sealant is presented. Discussed in this context are various forms of bentonitic sealants, i.e. chips and tablets, grouts, drilling fluids and granular bentonite. Placement methods of bentonite are described, and illustrated with a case history.

1.7 THE FUTURE Borehole sealing technology is well developed in a number of areas, particularly petroleum engineering, water well sealing, and exploratory and monitoring hole sealing. The necessity to seal holes upon decommissioning and abandonment is widely recognized. However, some uncertainties remain. Among the high priorities that deserve attention are the improvement

8

Introduction

of methods to detect abandoned holes, the development of non-invasive methods to verify that seals have been emplaced adequately and will perform satisfactorily, and the development and validation of methods to predict the performance of seals over an extended period. ACKNOWLEDGEMENTS The editors wish to acknowledge the support of the US Nuclear Regulatory Commission for the research support provided over many years, which has allowed us to perform in-depth studies of borehole sealing. We particularly wish to acknowledge the valuable insights and guidance provided by contract monitors Dr F. Larry Doyle and Mr Jacob Philip.

CHAPTER TWO

Laboratory Performance of Cement Borehole Seals D. L. South and K. Fuenkajorn

2.1

INTRODUCTION

This chapter presents an experimental method to assess the performance of cement borehole seals (plugs) under laboratory conditions. One of the prime goals is to obtain experimental data regarding the effectiveness of sealing. Laboratory conditions represent a highly idealized approximation of field conditions. This permits the determination of the best possible sealing performance that may be expected because emplacement of a plug may be done under carefully controlled conditions. In the field, plugs are emplaced in deep holes, in highly inaccessible locations, frequently in a wellbore filled with mud. Many variables are introduced in such a process, few of which can be controlled; some may not even be recognized. Laboratory testing permits a systematic, controlled variation of the parameters that influence plug behavior. For the work presented here, the parameters considered are rock type (granite, basalt and tuft), cement plug, and mechanical plug-rock interaction under varying stress fields. Borehole plug performance under varying stress field is investigated; temperature is held constant and saturated conditions obtained. A significant part of the effort is the development of laboratory equipment for performing the experiments. Investigation of the cement borehole seals is concentrated on their mechanical and hydrological performance. Where possible, results are presented as ratios to facilitate application to other rock types and plug materials. Applications of test results to some engineering aspects of borehole seal design are discussed. Sealing of Boreholes and Underground Excavations in Rock. Edited by K. Fuenkajorn and J. 1. K. Daemen. Published in 1996 by Chapman & Hall. ISBN 0 412 57300 8.

10

Laboratory performance of cement borehole seals 2.2 CONCEPTUAL APPROACH

The conceptual approach used to evaluate borehole seals in this work is to compare flow through a sealed borehole in rock with flow through intact rock. Rock cores 15 cm in diameter and 30 cm long (nominal dimensions) have 2.54 cm diameter holes drilled from each end, leaving a rock bridge in the center of the specimen (Figure 2.1). Water pumped into the top hole flows through the specimen to the bottom hole. Steady-state flow rates are measured. The rock bridge is then drilled from the sample and replaced with a seal and the experiment repeated. This allows direct comparison of flow rate through intact rock with flow rate through the same rock after a small portion has been removed and replaced by a seal. One of the main areas of interest is the performance of the seal under varying stress conditions. The intact rock specimen is placed under axial and confining stresses approximating a near-litho static stress field at a depth of about 1000 m. The intact rock is tested, the rock bridge cored from the specimen, and a plug placed and tested while the specimen remains under this stress field. Axial and confining stresses are then lowered to simulate depths of about 600 and 300 m, and flow through the plug-rock

I

Nul

2 Washer 3 Top Plate 4 5

Loadmq Plalen Piston Plug

6 1 8 9

Locd Cell Piston Cell Cap Specimen

10 Pressure Cell II Boll

12 13 14 15 16

CENTIMETERS

Fig. 2.1

Specimen and permeameter configuration.

Cent.finC} Pin Aluminium P'oten Neoprene Gasket Bottom Plote Bottom Plug:

Experimental apparatus

11

system measured. Lowering the stress field is a more severe condition with respect to the plug-rock interface than raising it because radial stresses across the interface are lessened. The method of investigation is chosen to meet two criteria. First, and most generally, experimental data are desired, as opposed to computer simulation. This has the advantage of having a rock-plug interface, as opposed to measurement on individual rock cores and individual cement cores. Second, the method permits simple, straightforward data analysis with few assumptions necessary. The main disadvantage of the method chosen is that the experiments take a long time to perform. Intact rock is chosen, as opposed to fractured rock, because a borehole plug seals against the intact zone of a rock mass; that is, the area between joints and fissures is the area in which sealing occurs. Plug material adjacent to a fracture will result in the diversion of fluid reaching that point into the fracture and out of the borehole. In this work intact rock is considered as rock in which no fractures are visible upon careful examination by the unaided eye. 2.3

EXPERIMENTAL APPARATUS 2.3.1

Permeameter design

An assembly drawing of the permeameter design is shown in Figure 2.1. Small black rectangles indicate O-ring seals. The permeameter is designed to accept a 15 cm diameter, 30 cm long rock cylinder with 2.5 cm diameter holes drilled along its axis. Aluminum platens are usually closer to the stiffness properties of rock; stainless steel platens are used for most of this work because they are more chemically inert. A normal axial stress of up to 21 MPa may be applied to the rock cylinder by tightening the bolts. The load is measured using a load cell. Fluid may be pumped into the top hole, the bottom hole, and about the annulus between the rock cylinder and the pressure cell. Neoprene gaskets are used to seal the ends of the rock cylinder, isolating the annulus from the top hole and bottom hole. These gaskets are shown as heavy lines on the assembly drawing. Nominal maximum fluid pressure is 21 MPa. Access to the interior of the specimen is provided by removing the piston plug and the bottom plug. This may be done while the specimen is under an axial stress and, if pressure is maintained about the annulus, under a confining stress. A centering pin in the bottom plug is used to align the specimen when it is placed in the permeameter; the pin is removed during testing. The specimen is coated with epoxy on the outside to prevent fluid seepage from the annulus through the rock to the center holes. Confining stress

12

Laboratory performance of cement borehole seals

is applied by pressurizing water in the annulus between the specimen and the pressure cell. At this point pressure in the top and bottom holes of the specimen may be zero. The neoprene gaskets at the ends of the specimen maintain the confining pressure. Sealing by the gaskets depends upon an axial stress higher than the confining stress. It is not possible to maintain a higher confining stress than axial stress. Using an axial stress of 22.6 MPa it is possible to maintain a confining stress of 19.6 MPa within 2% over a period of 24 hours. The confining stress is applied with a manually operated pump. Pressure is either maintained or drops as water leaks through the seals. This equipment, with slight modifications, can be used to investigate both temperature and moisture variations. 2.3.2

Pump design

Water must be supplied to a specimen in the permeameter at a constant pressure and at a very slow rate. To compare experiments it is desirable to fix the pressure and let flow rates be controlled by the permeability of the rock-plug system. To achieve this a constant pressure pump is designed and four are constructed. Compressed nitrogen is supplied to a large diameter cylinder of the pump, forcing a piston downward. A smaller piston is thus forced into a cylinder containing water, forcing water out of the bottom of the small cylinder at constant pressure. The fitting connected to the bottom of the water cylinder is connected with tubing to a fitting on the permeameter. Pressure intensification is approximately 11.5. A nitrogen pressure of 1.38 MPa yields a water pressure of 15.86 MPa. The pumps are made of stainless steel and have performed well. The main limitation is that pressures below about 1 MPa cannot be well maintained at constant pressure because of O-ring friction, which is also responsible for pressure fluctuations of 0.1-0.2 MPa at higher pressures. 2.3.3

Data acquisition system

Data are collected using an automatic data-logging system. The volume of water pumped into the sample is measured by the pump piston displacement with linear encoder. The amount of water flowing out of the sample is collected in a flask which sits on a force transducer. Fluid pressure is measured using National semiconductor pressure transducers Model LX-1450 AF and Model LX-1460AF. The Model LX-1450 AF has a range of 0-14 MPa and the Model LX-1460 AF a range of 0-21 MPa. The lower pressure range transducer is used to measure pressure in the top hole. The higher range pressure transducers are analog devices and are connected to the analog-to-digital voltage converter card. Axial stress is monitored using 50 ton (45455 kg) capacity load cells.

Experimental procedures 2.4

13

EXPERIMENTAL PROCEDURES 2.4.1

Sample preparation Rock specimen

Five different rock types are tested: two granites (Oracle granite and Charcoal granite), basalt from the Sentinel Gap area on the Columbia Plateau and two tuffs (Topopah Spring tuff and Apache Leap tuft). South and Daemen (1986) describe in detail the petrographical and mechanical properties of the rocks. Fuenkajorn and Daemen (1991a, 1992a) describe the properties of Apache Leap tuff. Cylindrical rock specimens are obtained either by laboratory coring of boulders collected from the field or by field drilling. The cylinders are cut to length, 30 cm, with a diamond saw and the ends ground flat and parallel. Grinding is one of the most important steps as flat, parallel ends are necessary to provide good end seals and to provide a uniform stress distribution. Specimen ends are prepared to specifications set forth by the International Society for Rock Mechanics (1981) for preparing samples for uniaxial compressive strength testing. This specification states that the ends shall be flat to 0.02 mm and shall be parallel to within 0.10 mm in 50 mm. Specimen flatness and parallelism are checked with a dial gauge. Next, 2.54 cm diameter holes are drilled along the specimen axis from each end to a depth of one-third the total length. A blind bit is used to flatten the bottom of the hole. In order to prevent seepage of water into the sides and ends of the specimens, several coats of epoxy are applied. Next, the sample is placed in the permeameter and a small axial load is applied to keep the top plate secure. The permeameter is turned over, the bottom plug removed and the bottom hole filled with distilled water. Enough water is poured into the bottom hole so that when the bottom plug is replaced water is forced from the valve, ensuring that no air is entrapped. The permeameter is righted and connected to a pump, ready for saturation and testing. Cement borehole plug

The cement used is a proprietary formulation provided by Dowell. It is composed of Ideal Class A cement (Tijeras Canyon) with 50% water content and proprietary Dowell additives D53 (10%), an expansive agent, and D65 (1 %), a dispersant. All percentages are weight percent with respect to cement. The cement is mixed according to American Petroleum Institute specifications (American Petroleum Institute, 1986). Dowell indicates that the slurry has density of 1.88 glcc, a strength of 26.2 MPa after 14 days of curing at 43°C, and 0.18% expansion after 14 days. Radial expansive stresses of the cement tested here have been measured by various investigators (South

14

Laboratory performance of cement borehole seals

and Daemen, 1986; Fuenkajorn and Daemen, 1987; Akgun and Daemen, 1991c). They conclude that after 25 day of curing under water the radial expansive stress of cement can increase by up to 4 MPa. Measurements of the cement permeability by South and Daemen (1986) yield an estimate of 95 nanodarcy. 2.4.2 Testing

When the specimen is first tested it has a rock bridge in place. The testing procedure is to first apply an axial and confining stress and to then apply a vacuum to the top hole. At this time the bottom hole is filled with water. The vacuum draws the air from the top hole and the pore space of the specimen. Distilled water is then injected into the top hole through a manifold which allows the water to be injected without admitting air to the top hole. Hence, the air has been withdrawn from the specimen and water is now being injected to saturate it. Once the specimen has been saturated, as evidenced by water flowing from the bottom hole at the same rate as it is injected into the top hole, testing begins. Table 2.1 summarizes the normal test schedule. With the specimen under an axial stress of 23 MPa and a confining stress of 20 MPa, a fluid pressure of 10 MPa is applied to the top hole by the constant pressure pump. Flow occurs through the specimen to the bottom hole, which is at zero (atmospheric) pressure. Following the test at 10 MPa top hole pressure, tests are performed at 7.0 MPa and 3.5 MPa to provide data on the variation in flow rate with injection pressure. Following the tests at the three different top hole fluid pressures, the rock bridge is cored from the specimen. Axial and confining stress are maintained during this operation. A rubber stopper is placed at the location of the bottom of the rock bridge. A plug of cement is then placed. The cement is then covered with Table 2.1 Nominal schedule for rock bridge and plug flow testing Step No.

1 2 3 4 5 6 7 8 9

Description

Load permeameter, check seal Saturate, establish flow Test at axial stress = 23 MPa and confining stress = 20 MPa (Injection pressures = 10.0, 7.0 and 3.5 MPa) Core out specimen, place plug Cure plug Re-establish flow Test at axial stress = 23 MPa and confining stress = 20 MPa (Injection pressures = 10.0, 7.0 and 3.5 MPa) Reduce stress, axial stress = 15 MPa and confining stress = 13.5 MPa (Injection pressures = 10.0, 7.0 and 3.5 MPa) Reduce stress, axial stress = 8.5 MPa and confining stress = 7.0 MPa (Reduce injection pressures to 3.5 and 1.7 MPa)

Analysis

15

water and allowed to cure. Following curing the rubber stopper is removed and the same series of tests performed on the rock bridge is performed on the cement plug. Flow through the plug will thus be directly comparable with flow through the intact rock. Next, the axial and confining stresses to which the sample is subjected are reduced and the test series repeated. Axial stress is reduced to about 15 MPa and confining stress to about 13.5 MPa. Following this test series axial and confining stress are again reduced, to 8.5 and 7.0 MPa. respectively. To avoid inducing tensile stresses, and to maintain the end seals, injection pressures of 3.5 and 1.7 MPa are used. 2.5

ANALYSIS

2.5.1 Flow calculation Finite element analyses are performed using the program FREE SURF (Neuman and Witherspoon, 1970) to determine the permeabilities of the rock bridge and plug from the measured flow rates and to estimate the amount and path of water flow through the specimen. For this analysis, the model is designed for a hollow cylinder having length-to-diameter ratio (LID) = 2, outer-to-inner diameter ratio = 6, with the rock bridge (or plug) having length-to-diameter ratio = 1 at the mid-section of the hole. This analysis is made in axisymmetry, assuming that the specimen and plug are homogeneous, isotropic and fully saturated, and that Darcy's law is applicable. The side and ends of the specimen are no-flow (impermeable) boundaries. The boundaries of the top and bottom holes are constant-head boundaries, the top hole at the head matching the injection pressure, the bottom hole at zero (gage). Figure 2.2 shows the finite element mesh and boundary conditions used in the flow analysis. Due the two symmetry planes, only one-fourth of the specimen geometry is analyzed. Due to the direct proportionality between the flow rate (Q) and the head difference (Ilh- between the top and bottom holes) and between the flow rate and specimen diameter (D), the calculated flow rates are presente

le(

0: W

a.

30

::;

w

I- 20

10

0

DISTANCE FROM CENTER OF EMPLACEMENT HOLE (M)

Fig 6.7 Temperatures measured in the buffer and the rock in three different directions, illustrating axial symmetry.

110

In situ performance of a clay-based barrier 6.5.11

Rock hydrogeology and buffer moisture regime

After the start of the heating phase, pore pressures in the rock responded dramatically to changes in temperature (Figure 6.8). The thermal expansivity of water is an order of magnitude greater than that of the rock, hence expansion of the water will be sufficient to cause the measured increase. The migration of the water into cooler rock away from the experiment has resulted in a subsequent gradual pressure decrease. Figure 6.9 shows the pore pressure contours surrounding the Buffer/Container Experiment and demonstrates that the pore pressures and the direction of flow are affected significantly by increasing temperatures. The pore-water pressure gradients in the rock are directly related to the rate of moisture supply to the buffer. Piezometers installed in the rock at the mid-height of the heater were used to determine the hydraulic gradients near the experiment (Figure 6.10). The flow towards the emplacement borehole, as calculated using these gradients, is approximately 0.06 mL/min into the entire emplacement borehole. It can be inferred from the data that flow patterns are complicated (Figure 6.9). The flow of water indicated by the contours of pore pressure implies a more variable degree of water supply from the rock throughout the length of the emplacement hole. It may be generally stated that water uptake was initially evident in the buffer near the buffer-rock boundary (Figure 6.11). Furthermore, significant increases in measured earth pressures slightly above the top of the heater could be taken as an indication of a notable increase in moisture in this area. Although not evident in the temperature measurements, the

ELAPSED DAYS FROM THE START OF HEATING (Nov 20,1991)

o

200

400

600

........ _-----

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0

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Fig 6.8

1RW25

Buffer/Container Experiment • Borehole HG11

1992

J

DEC JAN FEB MAR APR MAY JUN JUl AUG SEP

1993

Measured pore pressure increases following heater activation.

111

Buffer/container experiment

Scale:

r""'1"iii"" Pore pressure contours in kPa and possible flow paths ( - ) in the Buffer/Container Experiment.

Fig 6.9

An interpretation of pore pressure contours (kPa) surrounding the Buffer/Container Experiment and potential flow paths.

250

Pressure = 62.6 + 223.3 Ln{r} kPa

Flow

Of200

!2

=143.1 k (m /s) 3

Hydraulic conductivity (k)

• RP5

=1.4 x Hi'2

mls

i150

too 50

or-~~-----------------------------------1 0.8

1.2

1.4

1.6

1.8

2

Radial Distance from Borehole Center (m)

Fig 6.10

Pore pressure distribution in the rock at the heater mid-height.

2.2

112

In situ performance of a clay-based barrier

Fig 6.11 Moisture content changes in the buffer after 350 days (from psychrometer and thermal needle data).

moisture sensors and earth pressure cell responses adjacent to the heater in the buffer annulus region indicate thermal drying in the buffer annulus. 6.5.12

Buffer mechanical response

In situ compaction of the buffer resulted in 'locked-in' pressures of several

hundred kPa being recorded on both the contact and free-field earth pressure cells. Most of the cells show a gradual increase in pressure prior to heater activation (Figure 6.12). Earth pressure cells adjacent to and below the heater responded strongly to both heater activation and the power output change to 1200 W. The cells adjacent to the heater subsequently showed pressure decrease after the initial sharp increase. The pressure increases result from thermal expansion of the heater, buffer and other components of the experiment, whereas the decrease in total pressure is interpreted as being associated with consolidation of a zone of saturated buffer adjacent to the borehole wall under fixed displacement boundary conditions. The measured total pressures below and immediately above the heater continue to increase with time. The cells immediately adjacent to the

113

Buffer/container experiment

.~

GEOHOR CELL OATA EARTH PRESSURE

VS DATE

NOT TEJ.ftRATlRE COARECTEO

Fig 6.12

Total earth pressure response at the buffer-rock interface vs. time.

top of the heater (BG4 in Figure 6.12) have shown a particularly strong increase with time, probably a result of thermally induced vapour transport and subsequent swelling pressure development, as described above. The total stresses in the annulus region continue to decrease with time, with BG6 reading near zero pressure. These readings tend to support buffer drying and shrinkage in the annulus region, as described above. 6.5.13

Geomechanical response

Instrumentation was installed in the rock to monitor the strains induced by heating and by mechanical interaction with the buffer. Vertical displacements in the rock surrounding the emplacement borehole are measured in two extensometer strings. The strings are anchored 15 m into the rock and each has eight displacement transducers measuring vertical movement of the rock. Radial strain cells measuring horizontal stress changes were installed in boreholes adjacent to the emplacement hole. These cells measure changes in borehole diameter which can be converted into stresses using elastic theory. The extensometers and radial strain cells have been responding largely as expected. The measured vertical and horizontal strains are more or less proportional to temperature changes in the rock. The influence of the buffer

114

In situ performance of a clay-based barrier

on the rock displacements appears to be negligible. Of interest, however, is the fact that the in situ thermal expansivity of the rock appears to be greater than the thermal expansivities measured on rock samples in the laboratory. 6.5.14

Analysis and interpretation

History-matching analyses have been carried out during the course of the experiment to • track the progress of the responses and ensure that the experiment is operating more or less as expected; • assist with experiment operation (e.g. provide quantitative data to change the power output, as was required); • develop a better understanding of the system response during the experiment; and • identify areas of continuing uncertainty to assist with the design of experiment decommissioning. The analyses, which examined the coupled heat-moisture and mechanical response of the buffer, used updated and revised boundary conditions and material properties, as appropriate. The temperature changes, moisture changes and mechanical effects were systematically evaluated with the results of one analysis providing constraints and input on performance for the other. For example, as the heat transfer and moisture conditions of the buffer are intrinsically linked, the thermal response of the buffer would provide constraints on how significantly the moisture conditions could change to correspond to that observed thermal response. Similarly, the mechanical response of the buffer depends intrinsically on the moisture regime in the buffer. The Ontario Hydro Research Division coupled heat and moisture flow code TRUCHAM (Radhakrishna and Lau, 1992) and deformation analysis code TISDA (Lau and Radhakrishna, 1992) were used in the analyses. Based on history matching, the in situ thermal conductivity of the sand infill material was increased from the initial value of 0.38 W /m;oC to 0.45W/m;oC. In addition, to more closely reflect reality, TRUCHAM was run isothermally to account for potential water uptake during the installation phase; the resulting moisture content distribution was used as the initial condition for the heating phase. The rock-buffer boundary was modeled as a permeable boundary; the buffer was given free access to water, and water could move across the boundary, depending on the driving forces and transport coefficients. Other details of the analyses are reported in Radhakrishna and Lau (1992). Calculated and measured temperature profiles through a cross-section of the buffer are shown in Figure 6.13. Temperatures in the system generally match to within 1°C. Calculated moisture content variations across the buffer annulus at the mid-height of the heater are shown in Figure 6.14. Moisture is seen to be driven radially outwards

115

Buffer/container experiment 90

Heater

Buffer

Rock

80

.......................

Calculated Actual

70

G

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~

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201 Day

I- 40

54 Day 30

-._-...... 26 Day

20

B Day 10

o

500

1000

1500

2000

3000

2500

Radial Distance from Borehole Center (m)

Comparison of measured and calculated thermal gradients.

Fig 6.13

0.38

.1

0.36

~ Rock

~.--.--.---.--.--.--.--.

C 0.34 (I)

C o () 0.32

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(I)

:;

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0.26 0.24

Heater

0.22

o

50

100

150

200

250

Elapsed Time Since Start 01 Heating (days)

Fig 6.14

Calculated moisture content variations across the buffer annulus.

300

116

In situ performance of a clay-based barrier

toward the rock. These results are consistent with the trends of the moisture content sensor readings and the observed thermal response. Using the calculated moisture distribution (Figure 6.14) and laboratoryderived relationships between moisture content, elastic parameters, dry density and volume change (Lau and Radhakrishna, 1992), the response of the contact earth pressure cells on the borehole wall at the heater mid-height was modeled. The results are plotted on Figure 6.15. Two analysis cases were carried out: one with the sand annulus assumed to be compressible, and the other with the sand annulus assumed to be rigid. The graph indicates that the model forecasts the pressure responses reasonably and, in addition, matches the observed pressure drop near the end of the observed time period. The results of history tracking of the experiment demonstrated the usefulness of the observational approach used in this study when dealing with a system with complex interactions, and uncertainties are involved in the required material properties. Although the above analyses are encouraging and experiment performance has been excellent, several uncertainties still exist. Whereas the trends of the moisture content changes and the evolution of the moisture regime is sensible, quantitative interpretation of moisture content changes is extremely difficult. Interpretation of the moisture content sensors is complicated by total stress effects from heater thermal expansion and swelling pressure development, and by thermal expansion-induced excess pore pressures in the buffer. For example, the psychrometer may

600 ;\ 500

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.... 0 ....

incompressible sand layer

- - • - -

compressble sand layer

- - Measured Pressure Response

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-100OL--L.--1-00L.----J--200...L...-...L.....--:3~00---L.--:-40~0:--......L..--'---' Time (number of days starting from May 1.1991)

Fig 6.15 Calculated and measured total earth pressure response.

Isothermal test

117

register a change in total potential (vapour pressure) which reflects thermally induced vapor pressure changes rather than a change in local moisture content. The uncertainty of the moisture content regime leads to difficulties in interpretation of the total stress responses. Are the changes in total stresses largely due to local water content increases or is there a mechanical rebound component (release of 'locked-in' stresses) as well? Moreover, uncertainty still exists regarding the thermal and mechanical performance of the sand annulus material since its in situ properties were inferred from history-matching back analyses. These uncertainties can only be clarified through a well-designed and executed test decommissioning program.

6.6 ISOTHERMAL TEST The Isothermal Buffer/Rock/Concrete Plug Interaction Test, as mentioned earlier, is being conducted at the URL to provide supporting information for the Buffer/Container Experiment. The primary objective of the Isothermal Test is to assess the rate at which the buffer takes up moisture from the rock under constant temperature conditions. Installation of the Isothermal Test was completed in November, 1992 and, at the time of writing, the test is still in progress. 6.6.1

Experiment design

Both the Isothermal Test and the Buffer/Container Experiment were installed in boreholes 5 m deep by 1.24 m diameter in granite. However, the Isothermal Test design differs from that of the Buffer/Container Experiment in a few respects. The lack of a heater in the Isothermal Test means there is no internal cavity or sand installed in the test. Also, only the bottom 2 m of the Isothermal Test borehole were filled with buffer material. This allowed room within the emplacement borehole for a 1.25 m thick concrete plug to resist vertical expansion of the buffer. The restraint relies primarily on the bond strength between the rock and the concrete. However, should this bond fail with time, restraining bars embedded in the concrete and extending outward into boreholes in the rock will provide vertical resistance, preventing further movement of the plug. Essentially the same types of instrumentation were installed in the Isothermal Test as were used in the Buffer/Container Experiment. Data from 220 sensors are being recorded regularly for the Isothermal Test. In addition, a method for remotely assessing the buffer moisture content was implemented, and 32 electrodes are installed around the circumference of the emplacement borehole. By varying the electrodes used as the anode and cathode and by measuring the potential at the remaining electrodes, a tomographic image of electrical impedance within the buffer can be created

118

In situ performance of a clay-based barrier

(Strobel, 1993). The relationship between the impedance and the moisture content of the buffer has been established through a series of laboratory and field tests. 6.6.2

Geology and hydrogeology

The granite in the vicinity of the Isothermal Test has essentially identical thermal, mechanical and hydraulic properties to those determined for the rock surrounding the Buffer/Container Experiment. However, the granite at the Isothermal Test location exhibits pink alteration (hematization of feldspars) caused by proximity to Fracture Zone 2 (approximately 8 m below the base of the emplacement borehole). Within the borehole the rock is predominately medium-grained pink granite with some lenses of coarsegrained pegmatitic granite occurring near the base of the test hole. There are no natural fractures intersecting the borehole or within the experiment room. Although the thickness of the zone of visible excavation damage below the concrete floor is 0.5 m, the top of the concrete plug, and hence the top of the experiment, is 1.5 m below the floor. The background pore-water pressures in the vicinity of the Isothermal Test is approximately 1600kPa, as shown in Figure 6.3. Inflows into the test hole were monitored immediately before placement of the buffer and the measured inflow rate was consistent with a rock hydraulic conductivity of between 10 - 13 and 10 - 12 m/s. 6.6.3 Concrete plug The primary purpose of the concrete plug is to act as a restraint to upward movement of buffer resulting from swelling. The principal component of resistance is the bond strength between the concrete and the rock. Furthermore, if required, secondary restraint is provided by 32 stainless steel bars embedded in the concrete and extending outward into horizontal boreholes in the rock. It is anticipated that, when saturated, at a dry density of 1.73 Mg/m 3 , the buffer will exert a maximum swelling pressure of 1.5 MPa on the plug. Therefore, using a s~fety factor of 3, the bars were designed to resist shear forces under this maximum swelling pressure. The final dimensions of the concrete plug are 1.25m long and 1.24m in diameter. The concrete was a mixture of sulphate-resisting cement, crushed granite aggregate, silica sand, silica fume, superplasticizer and potable water. Approximately 3800 kg of fresh material was used for constructing the plug. Preparation of the fresh concrete was carried out underground near the test hole. A portable gravity mixer driven by an electric motor was used for the mixing. Results from laboratory tests show that the hardened concrete has an unconfined compressive strength of about 100 MPa at 28 days and the maximum long-term shrinkage of the hardened concrete should be about 0.2%.

119

I sothermal test 6.6.4

Test results

Temperature

The temperatures of the buffer, rock and concrete plug vary between 11 De and 13°e (Figure 6.16). However, short-term thermal transients were noted in the buffer and the surrounding rock during the construction of the plug. The increase in temperature in the buffer and the rock was attributed to the heat of hydration effect resulting from the hardening process of the fresh concrete used in formation of the plug. In response to the thermal effects from the concrete, the temperature of the buffer immediately adjacent to the plug rapidly rose from its initial temperature of 13°e to a maximum temperature of 30°C. The excess temperature in the system dissipated within 20 days. At steady state, a temperature gradient of about OAoe/m is noted along the length of the buffer mass, increasing towards the opening of the test hole. Furthermore, the temperatures in the buffer and the rock appear to be tracking the ambient underground temperature, which varies from l20e in the winter to 16°e in the summer. Generally, differences in temperature between the buffer and the rock are within ± 1°C.

Tomp$f:/ut-$ .

g

...... N

Open hoi

--

Design considerations

221

confinement for it to remain effective and the performance of seals is dependent on cementitious seal performance, as presented in Figure 9.6a. Prior to the drilling of the borehole, the initial permeability of the system is that of the formation. For rock salt this is very low, of the order of 10 - 22 m 2 or less. The permeability increases to that of an open borehole upon drilling. The emplacement of the seal material causes an immediate reduction in the seal system permeability. The permeability of the cementitious seal system indicates an almost immediate decrease upon seal placement to a value of about 10- 17 to 1O- 18 m 2 . This is considered to be the minimum permeability of a carefully constructed cementitious seal system immediately after placement, based on experimental evidence of similar systems (Peterson, Lagus and Lie, 1987). The permeability of the system can be attributed to the seal-rock interface and the adjacent disturbed rock zone. Creep of the host rock will tend to reduce the permeability of the disturbed rock zone and to create a tight seal-rock interface. After some time the seal system attains its lowest permeability, which is equivalent to the permeability of the seal material itself. For example, high-quality cements have permeabilities as low as 1O- 2 °m 2 . After some period of time, cementitious seal materials may begin to degrade because of the chemical incompatibility between the seal, the host rock and formation water. Assuming that degradation proceeds until the seal is reduced to its constituents, the ultimate condition of the seal may be comparable to sand and have a permeability of the order of 10- 15 m 2 . Because the permeability of the cementitious seal system will never achieve a permeability as low as the host rock and the seal is likely to degrade eventually, cementitious seals are considered to be long-term but not permanent seals by themselves. The highly time-dependent behavior of a granular salt seal system is illustrated in Figure 9.6b. For low initial emplacement densities (say 0.8 fractional density), the initial permeability of the seal system is largely a function of the seal material itself. These materials will continue to consolidate as a consequence of the creep closure of the borehole until they achieve a permeability equivalent to that of the host rock. The time required to achieve this degree of consolidation is of the order of hundreds of years for typical creep rates. Emplacement of granular salt of higher initial density offers an attractive alternative to cementitious seals. An initial emplaced density of 90% can provide immediate performance comparable to the cementitious seals (about 10- 17 m 2 ). As shown in Figure 9.6b, higher initial densities would be result in even lower permeabilities. The time required for the permeability to decrease to levels comparable to the intact formation for 90% or greater initial fractional density may be less than 100 years. It is important to note that the granular salt seals are not expected to degrade with time, due to the natural compatibility of the seal materials with the host rock. Granular salt seals are therefore considered to be permanent seals.

222

Sealing boreholes in rock salt

A design approach which takes advantage of the attributes of both cementitious and granular salt seals is the use of composite seals comprised of both types of seals. Granular salt seals are placed in the substantial rock salt sections. Cementitious seals are located between water-bearing zones and rock salt formations, preferably sealing across the contact of the salt and adjacent units. This approach utilizes cementitious seals to limit fluid movement into the granular salt as it is consolidating. Once the granular salt has consolidated, the performance requirements for the cementitious seals diminish. Successful design of a permanent borehole seal requires an understanding of the properties and behavior of the host rock salt. Borehole closure due to creep of the surrounding rock will drive the consolidation of the granular salt and eventually reduce the permeability of the disturbed rock zone. Creep rates vary from site to site, and even within a borehole, due to stress, temperature, moisture and impurity variations. Bounds on the expected creep rates must be developed, preferably from field measurements. Another important consideration in the permeability of the granular salt seal system is the magnitude and extent of the disturbed rock zone. This region, which extends about 0.5-1 borehole radius into the rock, can be sufficiently permeable to serve as a bypass of the seal system. For example, for a seal system with a granular salt seal emplaced with an initial permeability of 10- 17 m 2 , a disturbed rock zone with a permeability of 10- 16 m 2 will be the principal flow path through the system. Fortunately, the permeability of the disturbed rock zone will decrease to intact values as the formation creeps in and consolidates the granular salt seal. The inflow of water into the borehole may compromise a granular salt seal system. It is possible that if sufficient water enters the pore spaces of the granular salt it could impede further consolidation. One source of water is brine from the rock salt formation itself. Because the brine inflow rates are low they are difficult to measure. The few measurements of brine inflow which have been made suggest that brine inflow can be estimated, based on measurements of the formation permeability and pore pressure. Water inflow from adjacent water-bearing zones will probably be the principal source of water. The seal system (probably cementitious) between the granular salt seal and the water-bearing zone will be relied upon to limit this quantity. The emplaced density of the granular salt seal is a principal factor in the performance of the seal system. Simply dumping granular salt into the borehole will result in a fairly low initial fractional density (60-80%). The time to achieve consolidation will increase, the performance requirements on other sealing components such as cementitious seals will increase, and the certainty of the overall seal system performance will be reduced. The greater the initial density of the granular salt, the better. As with most granular materials, tailoring the moisture content and the particle size distribution can increase the initial density. The greatest improvement in initial densities, however, is achieved with some type of downhole compaction.

Conclusions

223

The above considerations have been combined into a model to predict the performance of a granular salt seal system (Nowak and Stormont, 1987). This model could be adapted or modified for other sealing applications, or a separate model could readily be developed. A model of this type requires estimates and simplified representations to be made of closure, brine inflow, water inflow from adjacent water-bearing units, the disturbed rock zone, and granular salt consolidation. Because it is important to limit the amount of water which enters the consolidating granular salt seals, an understanding of the performance of the portion of the seal system which protects the consolidating granular salt is necessary. These seals will consist principally of cementitious materials and perhaps other materials such as clays. The flow of water through these seal systems will depend in part upon the hydraulic pressure of the waterbearing zones and the permeability of the rocks between the water-bearing zones and the rock salt section. The permeability of the seal material and the seal-formation interface are obviously also important. It is reasonable to assume that the seals will degrade with time, although degradation rates and ultimate conditions are difficult to estimate. A model which predicts flow through a seal system above a consolidating granular salt seal was developed by Stormont and Arguello (1988). The model accounts for the formation pressure, formation permeability, seal and seal-rock interface permeability and seal degradation. This approach is one method of developing an estimate of the performance of the portion of the overall seal system designed to protect the consolidating granular salt. 9.7

CONCLUSIONS

A borehole seal design should strive for a seal system which is a permanent solution. This approach is necessary to satisfy existing or future regulatory requirements and to obviate the need ever to perform remedial measures on a borehole seal system. Sealing boreholes in rock salt requires an understanding of the host rock's response to the drilling of the borehole. Rock salt dilates and becomes permeable in the region immediately adjacent to the borehole. This region can be the most permeable component of the seal system, and thus can dominate its performance. However, as the host rock salt creeps inward and contacts the seal materials, the resultant stress build up tends to 'heal' the rock to a condition comparable to that before excavation. Granular rock salt has the potential to be the ultimate seal material for sealing penetrations in rock salt. Granular rock salt consolidates in a borehole which is closing from creep, reducing its porosity and permeability to that comparable to intact rock salt. Granular rock salt is chemically compatible with the host rock salt and therefore should be effective indefinitely. It is readily available and inexpensive.

224

Sealing boreholes in rock salt

In order to design an effective granular salt seal, the following information is required: (1) borehole closure rate, (2) extent and magnitude of the disturbed rock zone, (3) water inflow rate and (4) emplaced density of the granular salt. Experience has shown that borehole closure estimates need to be derived on a case by case basis. It should be well established that the borehole is closing at a measurable rate before emplacing a granular salt seal. The disturbed rock zone is expected to extend 0.5-1 borehole radius into the formation and to have a permeability which can be as great as 1O-16 m 2. Water inflow into the borehole can come from overlying or underlying water bearing zones, or from the rock salt itself. Conventional techniques such as pumping tests or drill stem tests can be used to estimate water inflow rates of producing formations. Because water inflow from the host rock salt is typically not measured, it may be desirable to use the limited inflow rates available from the literature. The emplaced density of the granular salt is a function of the emplacement technique. Techniques being developed which employ downhole compaction offer the best chance of constructing seals with minimum porosity and consequently will achieve a low permeability with minimal borehole closure and time. High-quality cementitious seals are an important part of the design of an effective borehole seal system in rock salt. These seals will isolate the consolidating granular salt from water-bearing zones above and below the salt formations. These seals also provide redundancy and can be used to seal across interbed layers of anhydrite, clay or other materials, if desired or required during seal construction. Clays, asphalt and epoxy are other materials which can serve as effective components in a seal system. The longevity of these materials is difficult to estimate, but is probably of the order of hundreds of years. 9.8

ACKNOWLEDGEMENT

The research and experience of the authors was developed while working for Sandia National Laboratories under contract to the US Department of Energy.

CHAPTER TEN

Design of Underground Plugs F. A. Auld

10.1

INTRODUCTION

To sink mine shafts and drive inclined drifts, underground roadways or tunnels successfully, experience and skill is needed to maintain excavation stability and to deal with and control groundwater. The presence of the latter is possibly the most serious threat to working in the underground environment and the miner must always operate with care when approaching known zones of water-bearing strata. During development work in shafts and tunnels, techniques are available whereby strata water can be controlled temporarily prior to installing a water-tight lining. Such methods are pumping, where the amount is not excessive; pre-grouting of the strata for reducing water to within the available pumping capacity; and freezing, if excessive amounts are expected. Before commencing development work hydrogeological boreholes are normally drilled from the surface to locate the water-bearing zones approximately. Pressure recovery tests are carried out within the boreholes are to provide data for estimating water inflow quantities which could be expected during excavation. Subsequently, forward probe drilling is carried out prior to each section of excavation to locate the water exactly. Such procedures allow development to take place safely irrespective of the presence of water. However, it is not always possible, or economic, to provide fully watertight linings for shafts and tunnels and, throughout the life of the underground system, ground relaxation and stress readjustment may allow further ingress of groundwater. Accidental inrushes of large quantities of water are also a potential hazard if mining takes place too close to undetected sources and ground instability occurs, or if drilling interconnects with unexpected water-bearing zones. Therefore, it can be seen that, in many cases, water will be prevalent in underground workings, whether it is expected or unexpected, and the means must be provided for sealing off areas of the workings either for Sealing of Boreholes and Underground Excavations in Rock. Edited by K. Fuenkajorn and J. 1. K. Daemen. Published in 1996 by Chapman & Hall. ISBN 0 412 57300 8.

226

Design of underground plugs

temporary water control while pumping to disposal or on a permanent basis. Plugs of concrete with a designed specific length and which fill the shaft or tunnel cross-section are used for this type of sealing. The design of underground plugs is well documented for the gold mines of South Africa where reasonably hard rock and relatively high water pressures are experienced at deep levels (Garrett and Campbell Pitt, 1958, 1961; Lancaster, 1964). However, very little new information has been forthcoming since 1964 and published design data concerning other situations in softer rocks and with lower imposed hydrostatic pressures are virtually non-existent. This chapter therefore sets out to review underground plug design, with the object of bringing the subject to prominence and more up to date. An attempt has also been made to rationalize the design process in relation to current practice. The following five sections of this chapter consist of a description of various types of plug; a discussion of the factors to be considered in plug design; detailed design calculations; construction aspects; and plug sealing and resistance to leakage. To elucidate the contents of these five sections more fully, the sixth section comprises three case studies of actual plugs. Based on the overall concepts contained in this chapter, conclusions and recommendations for plug design are formulated at the end.

10.2 TYPES OF PLUGS Four different catagories of underground plugs can be defined: (1) precautionary plugs, (2) control plugs, (3) emergency plugs and (4) temporary or consolidation plugs. Basic descriptions follow outlining the functions of each type. 10.2.1

Precautionary plugs

The plugs are normally constructed in underground roadways to limit the area of flooding should water inrushes occur. Watertight doors are built into them which can be shut when any danger of flooding arises. Precautionary plugs are installed as a safety measure prior to development in areas known to be potential water-bearing zones and such plugs are designed to withstand full hydrostatic pressure from surface level. 10.2.2

Control plugs

Sealing off or controlling the inflow of water from abandoned mining areas involves the introduction of control plugs. Plugs constructed in boundary pillars between adjacent mines also fall into this category. They are referred

Factors to be considered in design of plugs

227

to as boundary plugs and serve to prevent water flowing from abandoned areas of one mine into the workings of an adjacent mine. No means of access to the sealed off areas is provided through control plugs but normally drain pipes, with valves, are cast into them. These plugs are designed to resist full hydrostatic pressure from surface level or the pressure imposed by the head of water to the highest overflow point. 10.2.3

Emergency plugs

Plugs of this type are constructed to seal off unexpected inrushes of water either temporarily or permanently. No means of access to the sealed off areas is provided in such plugs and they are usually designed to withstand full hydrostatic pressure from surface level. 10.2.4 Temporary or consolidation plugs

Plugs which allow inflow water to be controlled or stopped while simultaneously providing the resistance for high-pressure grouting and consolidation operations are known as temporary or consolidation plugs. They are normally removed after the water pressure zones are sealed. Full hydrostatic pressure from surface level may again be the dominant design parameter for these plugs. 10.3

F ACTORS TO BE CONSIDERED IN DESIGN OF PLUGS

When designing underground plugs the following factors need to be considered: (1) the purpose for which the plug is to be constructed; (2) the type of excavation in which the plug is to be installed (shaft or tunnel); (3) where the plug is to be sited in relation to the prevailing rock and working conditions; (4) plug shape; (5) head of water to be withstood by the plug; (6) the condition of, and the stress in, the rock surrounding the plug; (7) the strength of, and stresses in, the material of the plug; and (8) the method of plug construction. 10.3.1

Purpose

Each of the four categories of plug described above has a different specific function and the form of a particular plug will be dependent upon the prevailing situation. 10.3.2

Type of excavation

Undisturbed ground stress conditions alter locally in the areas surrounding an excavation. The adjusted stresses differ, depending upon whether the

228

Design of underground plugs

excavation is for a vertical shaft or a horizontal tunnel. A more uniformly distributed stress occurs around a shaft excavation whereas, for a tunnel, the vertical ground pressure may be different from the horizontal, causing stress variations around the perimeter. In highly stressed ground a fracture zone may surround the excavation and its extent will also depend upon whether it encompasses a shaft or a tunnel. Therefore, the installation of a plug in a shaft will require different design considerations than for construction in a tunnel. 10.3.3 Location of plug One of the most important factors in deciding where to place a plug is the condition of the surrounding rock. Preferably, the ground should be free from geological disturbances which could provide leakage paths for water. However, there could be limitations for the choice of site and the presence of faults or dykes in the immediate vicinity may have to be accommodated. It is not advisable to site plugs in or near the fracture zones of highly stressed ground resulting from mining excavations although it is probably impossible always to avoid such situations. Control plugs may have to be located near mined-out areas to restrict outflow of water and the distance of boundary plugs from the workings depends upon the width of the boundary pillars in which they are installed. Boundary plugs need careful inspection at all times as boundary pillars which are too thin will be pervious to water and the danger of plug failure could be present under high hydrostatic pressures. Plugs should be sited in ground which is not likely to be affected by subsequent gound movements resulting from mining operations. Damage to both the plug and the surrounding strata would annul the grout sealing integrity and introduce fresh leakage paths. The prevailing working conditions could also influence the choice of plug site. When there is an inrush of water, depending on the amount of water flowing into the workings, preference will be shown for a site which can be temporarily dammed upstream, providing relatively dry construction conditions for the plug. Ventilation would be another criteria to be considered, particularly for an underground environment where high temperatures prevail. However, in an emergency an adequately ventilated site may not necessarily be forthcoming. 10.3.4 Plug shape Three basic forms of solid plug can be considered (Figure 10.1). The first consists of a thin reinforced concrete wall (Figure 1O.la) or unreinforced arch (Figure 1O.lb) keyed into the excavation all around the perimeter in contact with the ground. Design of the slab involves calculation of bending moments, shear forces and axial forces, sufficient strength being incorpor-

229

Factors to be considered in design of plugs Possible water leakage paths through strata , ; - - .........

*-

* ~~..~...~~.

JllIYii~II!l.

".' ~~

. «!.

---

Possible water leakage paths through strata

....

..

..

Water

Water

pressure

pressure

", ........

Possible water leakage paths through strata

(a)

_-_/

Possible water leakage paths through strata

(b)

...Water -pressure \'

(e)

1l\t".~''1I\",,,.

(d)

Steel bulkhead door

Steel bulkhead door •

O. D

lIS

(e)

In\\\'~

,'o~' ~

(I

(I.

.0'

•

•

..

o'

•

0

...0' o.

•

(f)

Fig.lO.l Plug shapes. (a) Reinforced concrete slab in rectangular opening (adequate strength but insufficient leakage resistance). (b) Unreinforced concrete arch in rectangular opening (adequate strength but insufficient leakage resistance). (c) Unreinforced concrete tapered plug in rectangular opening (adequate strength and leakage resistance but uneconomical). (d) Unreinforced concrete parallel plug in rectangular opening (economical with adequate strength and leakage resistance). (e) Unreinforced concrete cylindrical parallel plug, with human access, in circular opening. (f) Unreinforced concrete cylindrical parallel plug, with roadway access, in circular opening.

230

Design of underground plugs

ated in the structure to resist the applied pressure. The amount of keyed-in area is related to the bearing resistance of the surrounding ground. A solid plug of the second type possesses a longer length, no reinforcement and incorporates a taper to provide the ground-bearing area (Figure 10.1c). Parallel plugs are the third type (Figure lO.1d) and resistance to the applied end hydrostatic pressure is achieved through mechanical interlock with the rough excavation face of the surrounding rock. Garrett and Campbell Pitt (1958, 1961) consider plug length to be governed more by leakage resistance around the sides and through the surrounding rock than by structural strength. The longer length required for leakage sealing also ensures low shearing or bearing stresses at the concrete-rock interface. Thin barriers, although economic on materials, have very short, unsealable leakage paths at their extremities and so are not suitable for underground plugs. Tapered plugs, when compared with parallel plugs, require more rock excavation, which introduces further rock destressing, extra construction time and added cost. Increased quantities of concrete are involved and tapered plugs are subjected to larger pressures, resulting from the greater projected end area at the maximum cross-section dimensions. Such factors are a disadvantage when plugs are required to be installed under emergency conditions. Although the leakage resistance paths are adequate with tapered plugs, the other factors are prohibitive. Garrett and Campbell Pit (1958, 1961) have reported the results of tests in South Africa on an experimental plug at West Driefontein, and on a Virginia/Merriespruit boundary plug, which show conclusively that parallel but rough-sided excavations will retain a plug without any sign of failure under very heavy load conditions. On this basis, all further discussion on plugs in this chapter is focused predominantly on parallel plugs. The section on plug length based on bearing strength of concrete or rock at the interface (discussed later) does however contain tapered plug design theory. Solid plugs installed in shafts or tunnels will have a cross-section of the excavation shape in which they are constructed. Shaft plugs will generally be circular in cross-section whereas for drifts, roadways or tunnels the shape may be square, rectangular, D-shaped, circular or otherwise. For precautionary plugs with access ways through them, either purely for human entry (Figure lO.le) or roadway access for materials transportation (Figure 1O.1f), a different concept is required. To resist high strata grouting pressures, which are applied in the transverse direction for leakage sealing purposes, only the circular shape provides adequate strength. Precautionary plugs with access through them therefore need to be in the form of concrete cylinders with sufficient length for leakage resistance, adequate mechanical interlock automatically being provided. In plugs incorporating access roadways the dimensions required for clearance govern the inner diameter while strength to resist radial grout pressure determines the wall thickness. These two criteria apply for part of the length in a plug which is only provided with human access, as a structural

F actors to be considered in design of plugs

231

concrete wall can be incorporated integrally with the concrete cylinder at the upstream face (Figure lO.le). In this case the strength of the wall is adequate, the concrete cylinder acting as a sufficiently long sleeve to provide leakage resistance and mechanical interlock with the surrounding ground. In addition to the concrete cylinder, two other steel components are necessary for the successful operation of a precautionary plug with an access way: the bulkhead door for sealing off the plug in an emergency and a load transfer cylinder (Figures lO.1e, lO.lf). The steel load transfer cylinder allows the bulkhead door pressure to be carried by the concrete cylindrical plug through bearing on the ring flanges. Enough flanges are provided to reduce the bearing stresses to permissible limits. Load-transfer cylinders must also be designed with sufficient wall thickness to permit their interface with the concrete to be grouted up to the required pressure for sealing off water penetration resulting from the hydrostatic pressure on the plug.

10.3.5 Head of water to be withstood For the majority of plugs the design head of water will be that from ground surface to the level of plug installation. This should be taken as normal for all designs unless very clearly defined lower overflow levels are shown to exist below ground surface which produce heads of water that cannot under any circumstances be exceeded.

10.3.6 Condition of, and the stress in, the surrounding rock The successful sealing of water flow by the introduction of a plug depends on the capacity of the surrounding rock to prevent leakage. Any discontinuities in the strata will make the task of sealing off more difficult. Fissures of geological origin or fracture planes resulting from high ground stress could endanger plug performance and installation of plugs in such areas should be avoided wherever possible, as indicated in section 10.3.3. The type of rock in which a plug is constructed is also a very important factor in governing how well leakage paths can be sealed or how efficient the shearing resistance or bearing capacity will be along the concrete - rock interface. The presence of weak beds of shale, clay, sandstones or conglomerates will increase leakage potential and reduce interface shearing resistance or bearing capacity. As already indicated, undisturbed ground stresses alter once excavation takes place and the magnitude and variation of such stresses around an opening differ for shafts and tunnels. High ground stresses, which cause rock fracture, depend upon the following factors: (1) the depth below the surface, (2) the size of the opening, (3) the proximity of other mining

232

Design of underground plugs

excavations and (4) the proximity of geological disturbances which may introduce tectonic stresses. The subject of stress evaluation around underground openings is a complex one and is too large a topic to be introduced into this chapter. Nevertheless, it is a subject which must be fully understood if a true evaluation of concrete plug-rock interaction is to be formulated and more research into this area is required. 10.3.7 Strength of, and stresses in, plug material Five points warrant consideration when evaluating the stresses in, and strength of, underground plugs: (1) concrete compressive strength, (2) the early-age development of strength, (3) the shear or bearing stress at the plug-rock interface, (4) the pore-water pressure in the concrete and (5) the possible end spalling of the plug due to high stresses set up by ground pressure. Provided the recommendations of current Codes of Practice for structural concrete (in the UK, British Standards Institution, 1969b, 1985, 1987) are followed, with Grade 25 concrete (characteristic strength 25 N mm - 2) as the minimum specified requirement, then dense, impermeable and durable concrete ought to be achieved easily. On this basis, in conjunction with the length required for sealing which ensures low stresses, no problems of strength should be encountered. Early-age strength development is important from two aspects. First, it is essential that plugs develop their specified strength without any detrimental effects occurring from shrinkage, thermal changes or ground pressures. Provided care is taken to overcome these factors, the integrity of the concrete mass will be protected and leakage paths through plugs minimized. The second aspect is how quickly a plug needs to be sealed. It is possible to use higher-strength concrete mixes than are required purely for design strength. This allows higher strengths to be achieved at earlier ages and hence the problem of sealing can be tackled more quickly. The factor of safety against shearing or bearing failure in the rock or concrete of the plug at their interface depends upon the magnitude of the induced stresses, which in turn is related to plug length. Since the length of a plug should be determined with leakage sealing in mind, which means providing a longer length than is necessary for structural strength purposes, relatively low interface stresses are inherent in good plug design. Knowledge of pore-water pressure behavior within a concrete plug is limited. A pressure gradient exists from the hydrostatic pressure at the face in contact with the impounded water to zero at the opposite end. How the pressure and induced stresses are dissipated throughout the system and into the surrounding ground is a matter for conjecture at the present time and this area, in conjunction with rock stress evaluation, needs further research. Theoretical stress distributions in shaft plugs have been determined by Sitz,

Design calculations

233

Koeckritz and Oellers (1989) using finite element analysis. However, it is unlikely that spalling of the free face of a plug will occur due to porewater pressure unless non-homogeneous irregularities occur in the concrete mass. It is conceivable that high localized ground stresses at the ends of plugs could cause spalling at these points, reducing the effective resistance to applied pressure and leakage. Careful choice of site related to a study of the induced rock stresses and rock strength can avoid or minimize this risk. Additional plug length would also contribute to solving this problem. 10.3.8

Method of plug construction

For precautionary, control and temporary or consolidation plugs, which can generally be constructed in phase with and under normal mine operating conditions in a relatively dry environment, the method of construction has little influence on design. However, in conditions of emergency, materials access problems and water inflow quantities may require consideration of different methods of construction. Normal concrete transportation, placing and compaction can be replaced by grouted concrete in which a mixture of cement, sand and water is introduced into pre-placed aggregate (Chamber of Mines of South Africa Code of Practice, 1983). This technique is particularly suitable for the construction of plugs in areas where access is difficult or for plugs installed under water in flooded shafts. Concrete can also be placed by tremie under water if necessary. Resulting from the chosen method of construction, different concrete-rock interface allowable shear or bearing stresses may have to be used, depending upon how dense and impermeable the plug mass is expected to be and how integral a contact can be achieved with the surrounding work.

10.4 10.4.1

DESIGN CALCULATIONS

Formulae for calculating plug length and strength

Plug length based on shear strength of concrete or rock at the interface

Garrett and Campbell Pitt (1961) quote the following formula which can be applied to parallel-sided plugs with rectangular cross-sections if interface shearing is accepted as the governing failure mechanism: pbh = 2(b + h) IPpe'

(1O.1a)

where p is the intensity of applied pressure, b is the width of the plug, h is the height of the plug, I is the length of the plug and Ppe is the permissible punching shear stress of the rock or concrete at the interface.

234

Design of underground plugs

By transposing Equation (1O.1a), the length of plug can be obtained:

1=

pbh

2(b

+ h) Ppe

(1O.1b)

For a square cross-section, Equation (1O.1b) becomes

1= pb/4p pe

(lO.1c)

The length of circular plugs, of radius r, can be obtained from pnr2 = 2nrlp pe '

(lO.2a)

giving (1O.2b)

Plug length based on bearing strength of concrete or rock at the interface

Although the shear strength concept of the previous section can be employed, Garrett and Campbell Pitt (1961) also considered that, alternatively, the mechanism of interaction between concrete plugs and the surrounding rock could be more in the form of direct bearing rather than shearing at the interface. Mechanical interlocking action is achieved at an excavation face through the various inclined planes of its surface. Orientation of these planes can be in any direction lying between the extremes parallel with or normal to the general direction of the excavation face. An assumption can be made that half the inclined planes resist movement by direct bearing (Figure 1O.2a) while the others are subjected to tensile stresses and therefore can be neglected. For a parallel plug, consider an element of the excavation face ABC (Figure 1O.2b) with a horizontal length AC = 1', which contributes to the plug bearing resistance over the element of length 1'/2. The permissible bearing stress in the concrete or the rock is Pbe and F~ represents the total bearing resistance over the element of plug bearing length BC, inclined at an angle of ()( to AC such that BC cos ()( = 1'/2.

(10.3)

In the triangle of forces (Figure 1O.2b) P' is the element of applied horizontal force which is resisted by the horizontally resolved component of F~. Therefore P'

= F~

sin ()(;

(10.4)

however (10.5)

235

Design calculations

L - _ - - ' - - _ - - L_ _.L-_.....L_Total

effective resistance length =

(a)

i

-Compression component ex t of force on plug

AII+,______

~_D

________

~

(b)

Fig. 10.2 Evaluation of parallel plug length based on bearing strength of concrete or rock at the interface. (a) Plug bearing resistance. (b) Element of plug bearing resistance. (Garrett and Campbell Pitt, 1961.)

From Equation (10.3)

Be =

(10.6)

[1/(2 cos IX)

and combining Equations (lOA), (10.5) and (10.6) gives (10.7)

pi = (PbJ/2) tan IX

Summing all the forces on the plug results in ['

L pi = P = pbh = L Pbe 2: tan

[

IX

= 2(b + h) 2: Pbe tan IX

(1O.8a)

from which pbh

[=

(b

+ h) Pbe tan IX

(10.8b)

236

Design of underground plugs

Since the surface planes will be inclined at angles of between 0° and 90 to the direction of thrust, Garrett and Campbell Pitt (1961) considered the assumption that the average inclination IX = 45° for a parallel-sided plug was justified. Equation (1O.8b) becomes 0

1=

pbh

(b

(1O.8c)

+ h)Pbe

For a square cross-section, Equation (10.8c) reduces to (1O.8d)

1= pb/2Pbe' The length of circular parallel plugs can be obtained from

(1O.9a) glvmg [=

(1O.9b)

pr/Pbe

for IX = 45°. Tapered plugs can also be considered if appropriate amendments are made to Equations (10.3)~(1O.9) (Figure 10.3). The element of bearing length BC (Figure 10.3b) is now inclined at an angle IX + {3 to the horizontal and BC cos

IX =

['/2 cos {3,

(10.10)

where (3 is the angle of plug taper. From the triangle of forces P'

= F~

sin(IX + (3);

(10.11)

however ['

F' = P BC = P b be be 2 cos IX cos (3 .

(10.12)

Combining Equations (10.11) and (10.12) gives P

,

[' [' sin(IX + (3) {3 = Pbe ~ (tan IX + tan (3). 2 cos IX cos 2

= Pbe ~

(10.13)

Summing all the forces on the plug results in

I

P'

l'

= P = pb max hmax = I Pbe"2 (tan IX + tan (3) =

1

2(bay + hay) "2 Pbe(tan IX

+ tan (3),

(10. 14a)

where bmax is the maximum plug width at the water face, hmax is the maximum plug height at the water face; bay is the average plug width along its length and hay is the average plug height along its length.

237

Design calculations

.. Water pressure x end area = total force (p)

(8) B Compression component of force on plug

f (1 - tana tan~) (b)

Fig. 10.3 Evaluation of tapered plug length based on bearing strength of concrete or rock at the interface (developed from Garrett and Campbell Pitt, 1961). (a) Plug bearing resistance. (b) Element of plug bearing resistance.

From Equation (1O.14a) (1O.14b) For

rJ.

=

45° Equation lO.14b becomes (1O.14c)

and for a square section (1O.14d)

238

Design of underground plugs

The length of circular tapered plugs can be obtained from

1)2 + (r max -

pnr!ax = n(r max + rmin) [( "2 X

Pbe(tan rx

rmin)2

J1/2

+ tan /3),

(1O.15a)

where rmax is the maximum plug radius at the water face and rmin is the minimum plug radius at the face remote from water. Equation (1O.15a) gives r!ax 2 1= 2 [ P ( 2 2 rmax + rmin) Pbe (1

+ tan

/3)2 - (r max - rmin)

2J1 /2

(1O.15b)

for rx = 45°. An alternative form of bearing calculation for tapered plugs is that for a smooth-faced wedge driven into an opening. On this basis the whole surface area acts in bearing and the element of bearing length becomes AC (Figure 10.3b), inclined at an angle /3 to the horizontal where AC cos/3 = l'

(10.16)

sin /3;

(10.17)

and P'

= F~

however F~ = Pbe AC = Pbe1' /cos /3.

(10.18)

Combining Equations (10.17) and (10.18) gives P'

=

Pbe/' tan /3.

(10.19)

Summing all the forces on the plug results in

Ip' =

P = pbmaxhmax =

I

Pbe 1' tan /3 = 2(b av + hay) IPbe tan /3 (1O.20a)

from which 1=

pbmaxhmax 2(b av + haJ Pbe tan /3

(1O.20b)

is derived. For a square section 1=

pb!ax 4b av Pbe tan /3

(1O.20c)

Comparing Equations (1O.20b) and (1O.20c) with Equations (1O.14c) and (1O.14d), respectively, if rx = 45° is replaced by rx = 0° in the latter two equations then compatibility is achieved, except for the anomaly of reducing the length by half in the case of the wedge theory due to use of the full bearing area.

Design calculations

239

The equivalent length of circular tapered plugs based on the smoothwedge principle can be obtained from (1O.21a) where (10.21b)

Cylindrical plug strength

The strength of cylindrical parallel plugs (Figures 1O.le and 10.10 can be determined using the standard Lame elastic design theory for thick cylinders (Auld, 1979, 1982a): (J

=

ri

2Pr(t + t(t + 2r i) ~ Pc'

(10.22)

where (J is the maximum tangential stress in the concrete cylinder wall, occurring at the inside face, Pc is the permissible concrete compression stress, Pr is the externally applied radial pressure, r i is the inside radius of the cylinder and t is the concrete cylinder wall thickness. Bearing strength of cylinder walls

Plugs of the type shown in Figure 10.le, which carry load from a circular face wall back through a cylindrical rear section and thence into the surrounding rock, must have sufficient strength at the interconnection between the wall and cylinder. The cylinder end area must be sufficiently large to reduce the bearing stress imposed by the end wall to a value within the permissible limit. Hence, the calculated concrete bearing stress, (1O.23a) where Pb is the permissible concrete bearing stress, ro is the outside radius of the cylinder and P is the horizontal applied force on the cylindrical rear section; i.e.

P= pnr;

(1O.23b)

- [the concrete or rock permissible surface resistance over length 1* of the front wall]. Combined stress at the interconnection between the face wall and the cylindrical rear section

The cylinder stress and the bearing stress determined from the above sections act together at the interconnection to produce a combined compres-

240

Design of underground plugs

sive stress situation. Care should be taken to ensure that the calculated combined compression and bearing stress, (1O.23c) Punching shear of front wall

The punching shear resistance of the front wall against the cylindrical rear section (Figure 1O.le) must also be adequate. Therefore, the calculated concrete punching shear stress, fp

=

(10.24)

pnr;;2 nrr ~ Pp'

where Pp is the permissible concrete punching shear stress. 10.4.2

Formulae for designing steel load transfer cylinders

Load transfer to concrete by flanges

The concrete bearing stress, (10.25)

where rof is the outside radius of the load transfer cylinder flange, ro is the outside radius of the steel cylinder wall and n is the number of flanges. Flange bending

The bending moment is given by f; (r of - r0) (r of - r0 + t).

(10.26)

The steel bending stress, f ms =

3fb

-2

tf

(rOf - rO)(rof - ro + t) ~

(10.27)

Pms'

where Pms is the permissible steel bending stress and ness.

tf

is the flange thick-

Compression resistance to radial grouting pressure

This aspect is important for the sealing in of the steel load transfer cylinder into the concrete plug to prevent leakage along the interface. If d is the distance between the centers of the flanges, the flange crosssectional area is (r of - r0) t f and the cross-sectional area of the steel cylinder wall between the flange centers becomes td.

Design calculations

241

Evaluating the effective membrane thickness for both the cylinder wall and the flange gives the effective membrane thickness per unit length, to = [n(rof - ro) tf + (n - 1) tdJ/I.

(10.28)

The hoop stress for the steel load transfer cylinder, fes' can be evaluated using Equation (10.28): (10.29) where Pes is the permissible steel compression stress. To enable the cylindrical wall and flanges to act as a composite structure, any welding of flanges to the cylinder must be capable of carrying the interaction stresses. Direct bearing of end load on cylinder wall

The bearing area of the cylinder wall must be capable of transferring the end load from the bulkhead door to the bearing flanges without overstressing, i.e. steel bearing stress, P1tr~f rf)

f bs = 1t(r~ _ where

Pbs

~

Pbs

(10.30)

is the permissible steel bearing stress. Resistance to axial compression

In addition to end bearing, the cylindrical wall must be capable of acting as a column between the flanges to allow flange bearing to be effective. Timoshenko and Gere (1961) give the critical load, Per' for a column with built-in ends, which is a conservative approach for a cylinder wall as it neglects additional strength due to curvature, as (10.31) where E is the modulus of elasticity for steel (2.1 x 1011 Pa) and I is the moment of inertia of the cylinder wall equal to t 3 /12 per unit length. For the cylinder, multiplying Per by the circumference and dividing by the total applied end pressure will produce the factor of safety of (10.32) Provided the flanges are not spaced too far apart, satisfying the criterion of the section on 'compression resistance to radial grouting pressure' will automatically produce a large factor of safety in Equation (10.32).

242

Design of underground plugs 10.4.3

Bulkhead door design

Generally, bulkhead door pressures will be relatively large and therefore the best shape to resist the load is a spherical segment. The shell thickness for such doors can be determined on the basis that the meridional and hoop forces per unit length of the shell are equal to pal2 (Fliigge, 1967), where a is the radius of curvature of the shell. Dividing this value by the shell thickness, t, gives the compressive stress in the steel, f es' as fes

=

pal2t ~

Pes

(10.33)

It should be noted that thin shell domes are prone to buckling, and stiffening for the door should be provided to avoid any possibility of instability under load. The subject of steel bulkhead door design lies in the specialist field of pressure vessels and is outside the scope of this chapter. Operation and sealing of such doors are prime parameters to be considered in design and recourse should be made to specialist design and fabrication manufacturers for the supply of such elements.

10.4.4

Rock, concrete and steel permissible stresses

Permissible shear and bearing stresses for rock and concrete at the plug interface

The proposed formulae for determining the length of plugs, either on the basis of shear strength or on one of the two bearing strength philosophies are a very simplified form of a much more complex stress system. Both the rock and the concrete are in a confined state along their interface. The compression strength of concrete in the UK is quoted on the basis of 150mm cubes tested at 28 days in an unconfined compression testing machine (British Standards Institution, 1983). It is known that concrete, when tested in a confined state, shows an increase in strength over the unconfined condition (Jaeger and Cook, 1979). The confining action of the surrounding rock against the concrete plug could modify the bearing force calculated by the formulae in the section entitled 'plug length based on bearing strength of concrete or rock at the interface'. However, the true resistance probably lies somewhere between that given by the bearing capacity and the resistance provided by shear. Shearing in this context would be of a punching nature, as opposed to the traditional structural engineering form of beam shear, and even this could be modified, depending upon the magnitude of the interface confining stresses. Hence, the ultimate validity of the permissible stress values for shearing and bearing will depend upon the effectiveness by which the concrete of the plug is confined by the surrounding rock. The plug concrete can be considered as a homogeneous material on the assumption that good construction practice has been observed. However,

Design calculations

243

the surrounding rock will be anything but homogeneous, being cracked and fissured before excavation takes place. Destressing also occurs during and subsequent to excavation and therefore, when grouting and hydrostatic pressures are applied to the rock, movement inevitably will occur. The direct strains will be accompanied by movement in the direction of cracks and bedding planes and the effectiveness of the confining action will be dependent on this movement. Irrespective of which theory is applied to define the stress conditions, the governing factor remains the stress in the rock. As indicated above, more research is needed to understand how the stresses in the surrounding rock are modified by confined plugs subjected to end pressure. Until this aspect is investigated in detail the validity of any formulae utilized in defining plug and rock stress conditions will be in question. At the present time, with the formulae available, it will be necessary to check the shear and bearing stresses for both the concrete and the rock and to base the design on the weaker material. Concrete permissible stresses are contained in Table 10.1, based on the current UK Codes of Practice (British Standards Institution, 1969b, 1985). The values are all related to the concrete characteristic strength, this being the lower limit below which not more than 5% of the cube test results would fall based on a statistical analysis of samples tested. Both of the Codes of Practice are specifically for reinforced concrete and neither treats the unreinforced concrete situation realistically, particularly with regard to punching shear philosophy. However, Manning (1961) quotes the safe punching shear stress to be about one-fifth of the safe compressive stress and this has been included in Table 10.1. The maximum allowable values for Pc, Pb' Pbe' Pp and Ppe are heavily outlined in Table 10.1 as these are the suggested values to be adopted in design. The reason for using a factor of safety equal to 4 for Pbe and Ppe is explained below. It is much more difficult to propose realistic permissible stresses for rock. The strengths of rocks are normally determined by testing cylindrical samples and, as a result of the non-homogeneity of the material, it is normally only the best pieces from which the samples are obtained. It must always be remembered that strengths of rocks which are determined from such testing will not be typical of the actual strength in situ and appropriate adjustments should be made to allow for this. Assuming that the grouting process for strata water sealing is carried out methodically and conscientiously, most of the rock bedding planes and fissures local to the plug interface should be filled and consolidated. This, allied with the confining action of the surrounding rock, could allow the lower-strength concrete permissible stresses to be taken as being representative of the rock also. For the purposes of design, this would be an alternative if no actual data were forthcoming. The practice in South Africa is to use a permissible shear stress value of 0.59 N mm - 2(85lb in - 2) for concrete placed in the normal manner and 0.83 N mm - 2(120 lb in - 2) for grouted

a

Based on a factor of safety = 4;

h

0.67fcul 0.4fcul1.4 (1.4 x 1.5 7.14 7.98 8.57 9.57 10.00 11.17 11.43 12.76 12.86 14.36 15.95 14.29 15.71 17.55

BS811O:Part 1: 1985 value c Grade 25 (characteristic 30 strength,fc) 35 40 45 50 55 4.19 5.03 5.86 6.70 7.54 8.38 9.21

0.67fc)4

4.69 5.63 6.56 7.50 8.44 9.48 10.31

0.75fcul4

Concrete-rock interface bearing a , Pbe( = 3.75 Ppe) (Nmm- 2)

Not representative of unreinforced concrete in plugs

0.77 0.87 0.90 0.90 0.90 0.90 0.90

C

(fcul50) + 0.27 (Maximum = 0.90)

Beam shear (Nmm-2)

Type of stress

British Standards Institution (1969b) (withdrawn);

6.87 8.24 9.62 10.99 12.36 13.74 15.11

9.16 10.99 12.82 14.65 16.48 18.32 20.15

Grade 25 (characteristic 30 strength,fc) 35 40 45 50 55

0.75fcul2.73

Direct compression or bearing, Pb (Nmm- 2)

fcul2.73

Bending compression Pc (Nmm- 2)

Concrete permissible stresses

CP114:1969 value b

Table 10.1

British Standards Institution (1985).

1.28 1.44 1.60 1.76

1.12

1.60 1.92 2.24 2.56 2.88 3.20 3.52

2.56 2.94 3.30 3.66 4.04 1.28 1.47 1.65 1.83 2.02

0.80 0.96

2.20

1.84

Punching shear (Manning, 1961) Pp = 0.2 Pc (Nmm- 2)

1.10

0.92

Beam shear (Manning, 1961) 0.1 Pc (Nmm- 2)

1.34 1.51 1.68 1.84

1.17

0.84 1.01

(0.2 x 0.67fc)/4

~

1.25 1.50 1.75 2.00 2.25 2.50

0.2fc)4

Concrete-rock interface punching sheara Ppe (Nmm-2)

245

Design calculations

concrete where positive contact between the concrete and the surrounding rock is assured by subsequent grouting (Lancaster, 1964; Chamber of Mines of South Africa Code of Practice, 1983). This is a general rule and is not related specifically to concrete or rock strength; neither does it take into account the rock condition. The values are therefore unrealistic, particularly with regard to the increased concrete strengths currently being achieved in underground construction, due to the improved workability and quality control procedures adopted in conjunction with better batching, transportation and placing techniques. Therefore, it is considered that the values in Table 10.1 are more appropriate. Permissible concrete stresses other than at the plug interface with the rock

These are also covered by Table 10.1. Permissible steel stresses for load transfer cylinders and bulkhead doors

Typical permissible steel stress values for steel (Grade 43) are contained in Table 10.2. These are taken from the current UK Code of Practice for the use of structural steel in building (British Standards Institution, 1969a). 10.4.5 Factors of safety The factors of safety for structural concrete quoted by the UK Codes of Practice (British Standards Institution, 1969b, 1985) are given in Table 10.1. CP 114:1969 (now withdrawn) introduced a value of 2.73 to relate the characteristic strength to the permissible compression stress in bending, pc' BS 8110: Part 1:1985 is more specific in its breakdown of safety factor. The actual compression strength of concrete in a structure is equivalent to 0.85 x characteristic cylinder strength (Comite Europeen du Beton, 1970), where for plug construction the 0.85 factor takes account of the difference between instantaneous loads on cylinders at an age of 28 days and loads

Table 10.2

Permissible stresses in steel (Grade 43)a Type of stress

Thickness (mm) Bending, p (Nmm- 2 ) illS ~40

>40

165

150

Axial compression, pcs(Nmm- 2 )

155 140

a(BS449: Part 2: 1969, British Standards Institution, 1969a).

Bearing, Pbs (Nmm- 2 )

190 190

246

Design of underground plugs

applied for a longer duration on specimens of the same age. Since British Standard practice (BS 1881: Part 116:1983) uses the 28 day cube test as a means of strength control, a correction factor of 0.8 is needed to convert the cube strength to the equivalent cylinder strength. Hence, in relation to the characteristic cube strength, feu' the actual in situ strength of concrete is represented by 0.68 x characteristic cube strength, the value of 0.68 being equal to 0.85 x 0.8. A figure of 0.67 is employed in BS 8110 : Part 1:1985. Partial safety factors for load, Yf' and strength, Ym' are used in the ultimate limit state approach to the design of concrete structures (BS 8110: Part1:1985). On the basis of it normally being of a long-term permanent nature, the value of Yf for hydrostatic loading can be taken as 1.4. For Ym , which is introduced to account for possible strength differences between test specimens and the actual structure caused by such aspects as insufficient compaction and differences in curing, the specified value is 1.5. The effective factor of safety in accordance with BS 8110: Part 1:1985 is therefore 1.4 x 1.5 = 2.1 when related to the actual strength of concrete in situ, or (1.4 x 1.5)/0.67 = 3.13 when compared against the characteristic strength as given by 28 day cube test results. For the steel stresses in Table 10.2, the factor of safety to yield will be approximately 1.5 with probably the same again to failure. This gives a probable minimum factor of safety to failure equal to 2.25. The factors of safety for concrete and steel which have been built into the permissible stresses, in the range 2-3, are acceptable because the performance of the material under load is well established and quality control ensures consistency. For the mechanism of resistance at the plug-rock interface, the true behaviour is not understood fully and the rock shear and bearing permissible stresses cannot be established realistically. On this basis, and because plugs are normally installed as a safety measure, it would be prudent to adopt a higher safety factor when determining plug length using the shear or bearing resistance criteria. A minimum factor of safety of 4 is recommended in line with South African practice (Lancaster, 1964) and this has been introduced into Table 10.1 for the Pbe and P pe values. 10.5 10.5.1

CONSTRUCTION ASPECTS

Batching, transporting and placing concrete

In any concrete construction work it is necessary to have the right batching plant, geared to the demand. This is particularly important for underground construction, where pours must be completed with minimum interference from external factors. Rates of pouring are governed by the physical restrictions in particular areas underground and by the problem of access for materials to those areas. A surface batching plant is preferred because aggregate and cement weighing, together with the metering and addition of

Construction aspects

247

water and admixtures, can be controlled in the most effective manner. However, such a system depends on being able to transport the premixed concrete underground and, in some circumstances, an underground bat ching plant may be necessary. This does not relieve the problem of having to transport the concrete mix constituents underground as separate items. For grouted concrete, again a surface grout mixing set-up would be preferred. Normally, with surface batching plants large quantities of concrete or grout can be mixed and transported underground rapidly to give a constant uninterrupted supply. This is particularly advantageous for the construction of emergency plugs. Current trends in UK mining development have favored the employment of established ready-mixed concrete suppliers. By adopting such suppliers, high quality is achieved through the utilization of their specialist expertise in the production of concrete. Quality control measures for the use of concrete underground are the same as those for surface works and are in accordance with the relevant Code of Practice (BS 5328:Part 4: 1990). Independent approved organizations are normally employed for cube testing, or any other testing of the hardened concrete which is required. The preferred method of transporting and placing concrete underground is pumping. With shaft access, concrete can be dropped down a vertical pipe for further transportation underground by pumps situated at the bottom of the shaft. Drift access allows pumping from the surface, down the incline and then further underground directly to the point of plug installation. Once pipelines are installed, minimum interference with mining and plug construction operations is achieved and large volumes of concrete can be delivered and placed rapidly. Structural concrete mixes (Grade 25 and above) can now be transported from the surface, down shafts or drifts, directly to the point of placement underground through pipelines with small diameters, typically in the order of 32-50 mm (Martin, 1989).

10.5.2

Concrete mix design

Similar to any concrete construction, care must be taken to provide the right balance of ingredients, first of all to suit the particular mode of transportation and placing being used, and secondly to make sure that the optimum design is achieved. In addition to strength, the most important factor of the mix design is to obtain the correct workability. For underground work, with its restricted placing environment, it is essential that high-workability mixes are used. In the author's opinion (Auld, 1982b, 1982c), successful construction with concrete underground is entirely dependent upon the inclusion of plasticizing admixtures for transporting and placing. Two mix designs which have been used successfully by Cementation Mining Limited are contained in Table 10.3, on page 258. Cement replacement

248

Design of underground plugs

materials were incorporated to minimize the thermal effects which are discussed in more detail below. 10.5.3

Thermal and shrinkage effects

If it can be achieved, it is preferable to pour a concrete plug in one operation to avoid construction joints which are potential leakage paths through the plug itself. Although normal structural concrete mixes (BS 5328: Part 1: 1990) could be used for plug construction, the large volumes required for mass filling can be subjected to detrimental thermal effects during setting. This is dependent upon the amount of cement included in the mix. Internal buildup of heat within the mass due to the cement hydration process could induce high thermal stresses. The strength integrity of the structure would be impaired and, on cooling, thermal cracking may result. By cooling the aggregates and mixing water, the ultimate temperature attained by normal mixes can be reduced but it is preferable, where possible, to use cement replacement materials to minimize heat of hydration gain. Table 10.3 contains two such mixes. An additional factor which assists concrete thermal control is the embedment of service, water control and grouting pipes in the plug mass. Heat will be dissipated through these pipes, particularly in the case of temporary or consolidation and emergency plugs if water is flowing through them. Shrinkage will not normally be a problem with the designed concrete mixes currently being employed in underground construction. The use of plasticizers enables the water/cement ratio of the basic mix to be kept to a minimum, therefore ensuring very low water loss during the curing stage, which prevents excessive shrinkage. Three other factors also contribute to shrinkage reduction. These are the underground environment, in which no rapid drying out conditions normally prevail, the limited facial exposure to drying elements in the environment and the relatively thick concrete sections used. 10.5.4

Construction points of detail Excavation

Care should be taken during excavation to minimize damage to the surrounding strata. Machine cutting and hand trimming is preferred to drilling and blasting. Plug installation

Two factors assist in reducing leakage paths at the concrete-rock interface. Before beginning to pour concrete for a plug, the floor should be thoroughly cleaned to remove any debris or construction dust. At the roof of the

Plug sealing and resistance to leakage

249

plug, to ensure a tight seal, concrete must be discharged as high up as possible and a crown feed pipe, which can be withdrawn as topping up takes place from one end of the plug to the other, should preferably be installed. Air bleed pipes, which subsequently can be used for contact zone grouting, are also beneficial at roof level. Grout seals

Where mass concrete is cast directly against rock it is necessary to grout up the contact zone to prevent leakage through any shrinkage gaps. It is very difficult to obtain full tight contact with the surrounding rock over a large surface area with grouting. Therefore, it is preferable also to provide one or more narrow chases, surrounding the plug cross-section completely, in which grout can be injected and pressurized to provide a tight ring seal. Temporary water control

For temporary or consolidation plugs and emergency plugs, control of shaft water is essential to allow good construction. In roadways, floodwater will need to be temporarily dammed upstream and the water led off through valved pipes cast into the plug. Debris grills will need to be installed at the upstream ends of pipes. Temporary consolidation plugs in shafts also need to include vertical steel rising mains through which shaft inflow water can be withdrawn during construction, to prevent pressurizing of the underside of the plug whilst hardening. Services

Pipes need to be installed in precautionary plugs to carry services. These pipes should be fitted with sealing glands at each end for plug water tightness when the bulkhead doors need to be closed. 10.6

PLUG SEALING AND RESISTANCE TO LEAKAGE 10.6.1

Grouting procedure

Design calculations can be carried out as shown previously to determine plug dimensions. However, an integral part of the successful installation of an underground plug is the means by which leakage past the plug is minimized or eliminated. Grouting is the process by which this is achieved. The science or 'art' of grouting depends very much upon the knowledge and experience of mining development contractors, and cannot be discussed in detail here. As a process, grouting consists of the pressurized injection of cement or chemical grouts into the strata to fill voids, fissures, bedding

250

Design of underground plugs

planes and any other anomalies in the rock surrounding a plug. Its purpose is to seal off all water paths and grouting of the plug itself may be needed, depending upon whether construction joints are incorporated and also upon the standard of workmanship. Injection of grout at the contact surfaces between the plug and the rock is also necessary to fill shrinkage gaps, porous zones due to placing difficulties and cracks in the rock adjacent to the plug due to destressing. Pressures of up to twice (Garrett and Campbell Pitt, 1958) and 2.5 times (Garrett and Campbell Pitt, 1961; Lancaster, 1964) the pressure which the plug has to resist have been recommended for this grouting. These pressures are used in the deep gold mines of South Africa where generally strong rocks and relatively high water pressures are encountered. Even with localized fracture zones around such excavations, opening up of the cracks under high pressures to accommodate the entry of grout is not detrimental. However, in softer rocks at shallower depths, as in the UK coal measures, such pressures would be damaging and are not to be recommended. Precautionary plugs of the cylindrical type should only be stressed to a maximum of 1.25 times the hydrostatic pressure, related to surface level, this being the value by which the normal structural concrete permissible stresses can be exceeded for short-term loading (CP 114: 1969). Hence, the post-stressing of the plug and rock, which is advocated for the South African conditions (Garrett and Campbell Pitt, 1958) will generally not be as effective in UK practice for enhancing the confining action. The radial Poisson's ratio effect, resulting from the end pressure, will also be less effective with regard to increasing the interlocking resistance. Leakage associated with plugs can occur at the following places: (1) through the plug concrete, (2) along the concrete-rock interface, (3) through the rock surrounding the plug and (4) along the interface between the plug concrete and the steel load transfer cylinder if access through the plug is provided. 10.6.2

Leakage through the plug concrete

Three possible reasons exist for leakage through the plug concrete: (1) porosity of the concrete, (2) construction joints and (3) the presence of cracks. The presence of highly porous concrete is very unlikely, due to the dense, impermeable and durable mixes currently used in underground construction. High workability, achieved with the use of plasticizers, ensures full compaction and with good quality control and careful placing techniques there should not be any excess porosity problems. Grouting will help to seal off the more porous zones if they do occur. Wherever possible, construction joints should be avoided but, with the current mix designs, if they are necessary, very little joint preparation is required. Good bonding should be easily achieved between consecutive pours.

Plug sealing and resistance to leakage

251

Cracks in the concrete can be caused by excessive water pressure behind the plug, thermal effects during setting and maturing or excessive stresses and strains transmitted to the concrete by the surrounding rock. Provided the design is processed correctly in relation to the applied water pressure, adequate measures are introduced in the mix design to minimize the thermal effects and the chosen plug site is competent, a homogeneous structure can be constructed free from defects. 10.6.3

Leakage along the concrete to rock interface

Interface leakage could result from (1) shrinkage gaps at the interface, (2) shear cracks at the interface due to plug movement under high water pressure, (3) cracks caused by excessive ground stress and (4) poor contact with the surrounding rock, caused by debris and construction dust not removed from the floor prior to casting and also as a result of air and water pockets trapped at the underside of the roof. As discussed in section 10.5.3, shrinkage in underground concrete should be minimal. The grouting process also enables gaps caused by shrinkage to be sealed up. The possibility of plug movement will only occur if the plug length is too short and, as mentioned previously, the capacity to seal leaks is the prime factor in determining plug length. A longer length is needed to provide leakage resistance than is required for structural purposes. This will ensure that the interface shear stresses are sufficiently small to avoid any plug movement under high hydrostatic pressure. The radial adhesion between the plug concrete and the rock at the interface will be small and highly stressed rock conditions could damage the intimate contact. Judicious choice of the plug site could avoid such failure. Good workmanship will prevent problems such as construction debris not being removed prior to casting the plug concrete. Provision of the correct concreting facilities should assist in attaining close contact with the roof. 10.6.4

Leakage through the rock surrounding the plug

Leakage through the strata can occur as a result of (1) geological fissures or other discontinuities in the rock and (2) cracks in the rock formed by ground stress or by strain from mining operations. If possible, plugs should be sited away from faults in the rock. However, if they are unavoidable, the grouting process will help to seal and stabilize conditions. The problem of rock failure under high stress is experienced when other mining excavations encroach too closely or at great depth where overburden pressures become excessive. Every effort should be made to avoid overstressed areas.

252

Design of underground plugs 10.6.5

Determination of plug length required for sealing

The problem of leakage associated with underground plugs has been discussed previously on the basis of where it occurs and how it can be minimized or stopped by grouting. It is relatively straightforward to determine the plug length which conforms to the permissible punching shear and bearing stress values at the concrete-rock interface. Quantifying exactly the length which is required for a leakage-free plug is much more difficult. Published data concerning the subject are scarce and what information is available is related to specific ground conditions which cannot be applied on a general basis. Garrett and Campbell Pitt (1958, 1961) published the results of tests on an experimental plug, 1.220 m (4 ft) square by 3.350 m (7 ft 8.5 in) long and situated in sound quartzite, at West Driefontein. The static water pressure was approximately 20.7N mm - 2 (3000 lb in - 2). An extensive system of tapping points and holes in the rock were incorporated for studying leakage at the steel load transfer cylinder interface with the concrete, at the concrete-rock interface and through the strata. Leakage quantities were observed before grouting and after various stages of pressure grouting were completed. From the test results, Garrett and Campbell Pitt (1958) proposed certain concepts which are given below. • The resistance of a plug to the passage of water either along its contact with rock or through the adjacent fractured rock depends on two factors: the length of the plug and the resistance of the rock to the passage of water. • The latter, being a condition of the rock which varies greatly with different types and mining conditons, can be regarded as the practical consideration for determining plug length. • The two factors can be interrelated using the pressure gradient through the rock as the linking medium. Results from the West Driefontein test plug form the basis for the graphs contained in Figure 10.4 which are reproduced from Garrett and Campbell Pitt (1958). These results refer only to the rock and pressure conditions described. The graphs are (A) the minimum length of plug that would be required if the contact between plug and rock was ungrouted [P/l=0.23Nmm- 2 m- 1 (20.8Ibin- 2 ft- 1 )]; (B) the minimum length when the contact is grouted but before the rock is grouted [P/l = 3.64 Nmm- 2 m- 1 (161lbin- 2 ft- 1 )]; (C) the minimum length when normal grouting of the rock was 41.4Nmm- 2 (6000Ibin- 2 ) [p/l=9.14 N mm - 2 m - 1 (4041b in - 2 ft -1)]. This is normal to South African practice, being twice the hydrostatic pressure, but is not normal to the UK. The graph (D) is similar to C but with the addition of chemicals to seal rock fissures. Graph C is applicable in South Africa to a normally grouted plug but has no safety margin.

253

Plug sealing and resistance to leakage OOr-------~-------r--_r------,_----c__,

70

B I I-

C)

Z

~

30

....... .

\\ 4 \\'f.~ ..... ;,;. ~ ~--a

...........

-

c

~

~~~=::::J.D 2

3

4

6

6

10

12

14

16

HEAD(XlO' It) (WATER TABLE TO PLUG)

Fig. 10.4 Required plug lengths to resist hydrostatic pressure based on leakage resistance and bearing (Garrett and Campbell Pitt. 1958). Results in imperial units are used as presented by these authors. Dotted lines are based on bearing (Equation (1O.8d)), all other lines relate to leakage resistance of the test plug.

Garrett and Campbell Pitt (1958) suggested from this that plug length should be such that a leakage factor of safety should not be less than 4 and may be as much as 10. The choice depends on an assessment of many factors which include fracture of rock during excavation and subsequent destressing, porosity of the rock and its acceptance of grout. In Figure lOA the graphs show plug lengths when factors of safety of 4, 6, 8 and 10 are applied. These depend on the plug-rock contact and the rock being grouted to at least the same pressure as that which the plug is designed to resist. Plug lengths for various square section sizes based on a Phe value of 4.14Nm- 2 (600Ibin- 2 ) are also included on the basis of Equation (1O.8d) (shown by dotted lines). The value of 600 lb in - 2 was used by Garrett and Campbell Pitt (1958). As far as the author is aware, this is the only published information which attempts to quantify directly leakage resistance in relation to plug length, apart from records of past plugs which have been successful. It has been emphasized throughout this section that the data put forward relates only to the particular test conditions. This leaves the plug designer very much to his own initiative and experience in determining the plug length which will provide adequate sealing. Further research is therefore necessary into this

254

Design of underground plugs

area. However, if a plug is constructed with sufficient length to resist movement but cannot prevent leakage through the surrounding rock, the length can always be increased. In an emergency this may be important, because the plug could be constructed to prevent flooding and subsequently lengthened to reduce leakage. 10.7

CASE STUDIES

10.7.1 British Gypsum Ltd, Sherburn Mine, England, 1980 (Emergency plug) In 1980 a pressure pad was constructed by British Gypsum Ltd, (section A-A of Figure 10.6) in an attempt to seal off water inflow into the area of the pump sump (1 East 5 South in Figure 10.5). The general dip of the strata was from left to right in Figure 10.5 (west to east) and it was considered that water was following the interface between the gypsum, in which the roads were driven, and the marl bed below and making its way down the strike. Excavation of the gypsum appeared to have encroached on the marl bed and allowed water to come up through the floor. The main access to the mine was via the 1 in 4 adit which was close to the inflow position. Upon failure of the pressure pad during grouting operations, Cementation Mining Ltd were asked on 21 October 1980 to design a new scheme for sealing off the water. The water inflow at that time was estimated to be 182 L s -1 (2400 gal min -1) (Figure 10.7). Various structural schemes for pressure pads and combinations of pads and plugs were considered and discarded in favor of the complete plug solution shown in Figure 10.6 for simplicity, speed of construction and permanency. Urgency was the main criteria as, within 6 days of Cementation Mining Ltd being called in (27 October), the water inflow had risen to 379 L S-l (5000galmin-1) and it was rapidly becoming obvious that there was a danger of losing the mine. The plug scheme adopted is detailed in Figure 10.6. It was deemed prudent not to disturb the remaining sections of the original pressure pad. Another gravel bed was laid over the top in which six more water control pipes were placed in addition to the two pipes (one 200mm diameter and one 100 mm diameter) previously installed below the original pressure pad. The additional pipes were four of 200 mm diameter and two of 300 mm diameter and carried the water to a new sump position adjacent to the proposed plug site. Additional rising mains were installed in the shaft to cope with the increasing inflow. The first concrete was poured on 28 October and Figure 10.6 shows the concreting stages. Because of the large mass of concrete involved, construction joints were necessary and a low heat of hydration mix, incorporating a cement replacement material, was used (Table 10.3). Concrete was pumped

255

Case studies

..

'

6>~~' . . . Rising main q~ .o~~r

borehole to surface

--------.~~~~----~~~~

Emergency plug

_JI;~L==IT~~

~

~

'il-

0I

5I 10 15 20 25 I I I I meters

\i

D[ ,---,

I

Fig.tO.S British Gypsum Ltd, Sherburn Mine, England. Underground layout showing position of emergency plug.

from the surface down the 1 in 4 adit, through a 100 mm pipe, directly into position in the plug. The 4 week time period for placing the concrete resulted from various equipment, labour and general construction problems but once concreting had commenced the water inflow was controlled at a peak level of 606 L s - 1 (8000 gal min - 1) (Figure 10.7). Minimal true design was required for the Sherburn Mine plug. The depth below ground level was 48 m, which resulted in a hydrostatic pressure of 0.47Nmm- 2 (68lbin- 2 ). This is not excessive and the length of the plug was extremely long. However, length in this case was governed by practical considerations to suit the particular situation. The pressure gradient from

l

)\'1111 •. \\\\, _

\\'11 __ .. 1)\)\\"

--\\',\,'))')\

)"1. .. '"1111\'

B~

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Fig 10.7 British Gypsum Ltd, Sherburn Mine, England. Water inflow quantities related to concrete poured in emergency plug, showing effectiveness of water stopping (water inflow quantities in imperial units as recorded).

one end to the other was only 0.47/35.3 = 0.013 N mm - 2 m- 1 (0.59Ibin- 1 ft- 1)and the 35.3m (116 ft) length (Figure 10.6a) was eventually extended out to the adjacent access roadway. Figure 10.7 indicates how effective the plug was in stopping water. On completion of the various concrete stages, the control pipe valves were closed and the inflow almost completely stopped. Final sealing by grouting commenced after the valves were shut off and involved a combination of grout pipe positions. Some were previously cast into the plug to reach places which would have been inaccessible by drilling from the two plug faces. These, in addition to injection at the inflow point through the water control pipes and to the contact zones through other holes drilled from both faces, enabled the water to be sealed off completely on a permanent basis. Only cement grout injection was necessary. 10.7.2 Proposed precautionary plug, 1981, for overseas contract by Cementation Mining Ltd Figure 10.8 contains a proposal for three precautionary plugs to be installed near the bottom of a shaft for access protection in the case of an inrush. Construction of the plugs was to be in a thin limestone bed, 15 m thick,

258

Design of underground plugs

Table 10.3 Cement replacement mixes previously used by Cementation Mining Ltd for underground plugs

Site Supplier Total cementitious content Sand Sand % of total aggregate Coarse aggregate Water Water: cement ratio Slump without plasticizer Plasticizer Slump with plasticizer

Emergency plug in roadway: Grade 30 (0 PC replacement with PFA); 30Nmm- 2

Temporary consolidation plug in shaft: Grade 55 (OPC replacement with Cemsave ground granulated blast furnace slag); 55Nmm- 2

British Gypsum Ltd, Sherburn Mine Topmix Ltd 400 kg m -3(250 kg m -3 OPC, 150kgm- 3PFA) 770 kg m - 3(Elvaston Zone 2) 42

National Coal Board, North Selby No.1 shaft Topmix Ltd 500 kg m - 3 (30% OPC, 70% Cemsave) 595 kg m - 3 (Blaxton Zone 3) 34

1050 kg m - 3 (Elvaston Gravel) 180Lm- 3 0.45 50mm

1150 kg m - 3 (Blaxton Gravel) 180Lm- 3 0.36 60mm

Flocrete N (Cementation Chemicals Ltd) 160mm

Flocrete N (Cementation Chemicals Ltd) 160mm

situated above and below weak, water-bearing strata zones at a depth of 542.5 m (1780ft) [5.43Nmm- 2 (787Ibin- 2 ) hydrostatic pressure]. Design of the plug, load transfer cylinder and bulkhead door was carried out in accordance with the design calculation section, Grade 35 concrete being specified. The concrete-rock interface calculated punching shear stress was 0.63Nmm- 2 (91lbin- 2 ) and the pressure gradient 5.43/12= 0.45Nmm- 2 m- i (19.9 Ibin- 2 ft-i). The proposed grouting scheme depended on the actual ground conditions at the level of the plug when the shaft was eventually sunk. However, care needed to be taken above and below the plug in order not to encroach too close to the water-bearing zones. 10.7.3

National Coal Board North Selby Mine, England, 1982 (temporary consolidation plug)

During December 1982 Cementation Mining Ltd were sinking two shafts at North Selby for the National Coal Board Selby project. Both shafts had reached the stage of sinking through the Ackworth Rock, which is a Coal Measures sandstone and an aquifer, with No.1 shaft sump (Figure 10.9) standing at - 540.2 m (1772 ft) below surface level [hydrostatic pressure

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Design of underground plugs

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Case studies

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Fig 10.9 National Coal Board, North Selby Mine, England. Section through shaft, showing (a) temporary consolidation plug and (b) plan at pump lodge level.

5.4 N mm - 2 (7831b in - 2)]. The previous sump level in No.1 shaft stood 13.8 m (45 ft) above the new sump level and strata cover grouting was carried out from the previous level.

262

Design of underground plugs

During the period of strata cover grouting, problems of grout standpipe installation were experienced due to the poor rock conditions and deterioration and heave of the sump took place. The length of cover grouting was also long (over 40m) whereas the preferred maximum length was approximately 30m. To enable the wall of the cover grouting cone to be less prone to leakage at the lower level of treatment and to guarantee satisfactory grout standpipe installation, it was decided to sink to the - 540.2 level and install a concrete plug. This would be closer to the zone of strata requiring the major grout treatment and, by casting the grout standpipes into the plug, a pressure pad for the next cover of strata grouting could be provided. Due to the potential water inflows for sinking below the plug, it was necessary to install a pump lodge (Figure 10.9) for stage pumping to surface. No choice of position was available for the pump lodge other than immediately below the last cast section of the shaft wall. At the time of placing the plug, shaft water inflow to the sump was approximately 11 L s -1 (150 gal min -1). Figure 10.10 shows the framework for supporting the grout pipes and the water control rising mains during casting of the plug. The concrete mix design for the plug is given in Table 10.3. Minimal heat of hydration existed in the concrete mass, due to use of the cement replacement material (Cemsave) and additional heat removal occurred through the rising mains and grout pipes. The Grade 55 concrete was the same as the shaft lining concrete. However, designing the plug on the basis of 7 day cube test results, (2/3) x 55 = 36.7 N mm - Z (5317lbin- Z), allowed pressurizing of the plug for water stopping at the earliest opportunity. The benefit of the 28 day strength was taken for the wall-bearing resistance. The recommendations given in the design calculation section for cylindrical plugs were followed for the plug design. Neglecting the bearing resistance of the tapered plug, the calculated punching shear stress for the concrete-rock interface was 0.89 N mm - Z (129 lbin -Z) and the pressure gradient was 5.4/17.3 = 0.31 N mm -z m- 1 (13.7Ibin- 2 ft-1). Grouting up of the plug started from the bottom through 50 mm grout pipes installed in the rising mains. These pipes were grouted in, leaving the bottom free for injection into the gravel bed, and also secured by highpressure flanges bolted together at the top of the rising mains. The bottom injection was phased to follow backwall injection of the shaft wall above the plug, and controlled using the standpipes as 'tell-tales' before closing off for final pressurizing. The shaft water make was reduced to approximately 0.45 L S-l (6 gal min -1) before final tightening up, this amount being predominantly from behind the shaft lining above the pump lodge. The pump lodge was restricted to a position close to the plug. To enable the plug to be subsequently broken out without damaging the shaft wall, the bottom surface of the wall was painted with a bond-breaking agent, Setcrete 11 (Don Construction Chemicals Ltd); the hanging rod ends

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