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2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12.

13.

14. 15.

of Steel, Iron and Steel Institute 1971,p 78 T.G. Byrer, Ed., Forging Handbook, American Society for Metals, 1985 H. Gegel, S. Nadiv, and R. Raj, Dynamic Effects on Flow and Fracture During Isothermal Forging of a Titanium Alloy, Scr. Metall., Vol 14, 1980, p 241 R. Raj, Development of a Processing Map for Use in Warm-Forming and Hot-Forming Processes, Metall. Trans. A, Vol 12A, 1981, p 1089-1097 S.L. Semiatin and G.D. Lahoti, The Occurrence of Shear Bands in Isothermal Hot Forging, Metall. Trans. A, Vol 13A, 1982, p 275-288 M.C. Mataya and G. Krauss, A Test to Evaluate Flow Localization During Forging, J. Appl. Metalwork., Vol 2, 1981, p 28-37 F.N. Rhines and P.J. Wray, Investigation of the Intermediate Temperature Ductility Minimum in Metals, Trans. ASM, Vol 54, 1961, p 117 A.M. Sabroff, F.W. Boulger, and H.J. Henning, Forging Materials and Practices, Reinhold, 1968 V. Vujovic and A.H. Shabaik, A New Workability Criterion for Ductile Metals, J. Eng. Mater. Technol. (Trans. ASME), Vol 108, 1986, p 245-249 M.G. Cockcroft and K.J. Latham, Ductility and the Workability of Metals, J. Inst. Met., Vol 96, 1968, p 33-39 A.L. Hoffmanner, Technical Report AFML-TR-69-174, U.S. Air Force Materials Laboratory, June 1969 Y.V.R.K. Prasad, H.L. Gegel, S.M. Doraivelu, J.C. Malas, J.T. Morgan, K.A. Lark, and D.R. Barker, Modeling of Dynamic Material Behavior in Hot Deformation: Forging of Ti-6242, Metall. Trans. A, Vol 15A, 1984, p 1883-1892 H.L. Gegel, Synthesis of Atomistics and Continuum Modeling to Describe Microstructure, in Computer Simulation in Materials Processing, Proceedings of Materials Science Seminar, ASM INTERNATIONAL, 1987 P. Dadras and J.F. Thomas, Metall. Trans. A, Vol 12A, 1981, p 1867 S.L. Semiatin, J.F. Thomas, and P. Dadras, Metall. Trans. A, Vol 14A, 1983, p 2363

Workability Tests George E. Dieter, University of Maryland

Introduction WORKABILITY is a complex property of a material, as indicated in the article "Introduction to Workability" in this Section. It is difficult to isolate the intrinsic workability because this property is strongly influenced by stress state, which is in turn affected by friction and by the geometry of the tools and the workpiece. It has also been shown that the workability of a material is strongly influenced by metallurgical structure and that workability can be a complex function of temperature and strain rate. At the current state of development, the ability to model a forging process by calculating stress, strain, strain rate, and temperature throughout a deforming workpiece with a computer-based finite-element technique exceeds the ability to predict the workability of the material. A large number of tests are currently used to evaluate the workability of a material. The primary tests--tension, torsion, compression, and bend--will be discussed in this article. These are tests for which the state of stress is well defined and controlled. Of these four tests, the compression test has been the most highly developed as a workability test. The cold upset (compression) test will be described in detail in the article "Workability Theory and Application in Bulk Forming Processes" in this Section. Specialized workability tests that have been developed from the four primary tests will also be covered. Each of these tests provides information that is not readily available from the primary tests.

A number of workability tests will be discussed that are especially applicable to the forging process. These forgeability tests are used because they lend themselves to the form of material or the particular forging process. Although most workability tests seek to determine the extent of large-scale deformation that is possible before fracture, one class of test is concerned with the propensity for localized deformation and fracture. Finally, this article will conclude with a discussion of typical forging defects. Although not strictly related to workability tests, forging defects certainly represent a limitation to workability. Workability Tests George E. Dieter, University of Maryland

Primary Tests The primary tests for workability are those for which the stress state is well-known and controlled. Generally, these are small-scale laboratory simulation tests. The tension test is widely used to determine the mechanical properties of a material (Ref 1). Uniform elongation, total

elongation, and reduction in area at fracture are frequently used as indices of ductility. However, the extent of deformation possible in a tension test is limited by the formation of a necked region in the tension specimen. This introduces a triaxial tensile stress state and leads to fracture. For most metals, the uniform strain that precedes necking rarely exceeds a true strain of 0.5. For hot-working temperatures, this uniform strain is frequently less than 0.1. Although tension tests are easily performed, necking makes control of strain rate difficult and leads to uncertainties about the value of strain at fracture because of the complex stresses that result from necking. Therefore, the utility of the tension test is limited in workability testing. This test is primarily used under special high strain rate, hot tension test conditions to establish the range of hot-working temperatures. A description of this test method can be found later in this article. In the torsion test, deformation is caused by pure shear, and large strains can be achieved without the limitations

imposed by necking (Ref 2, 3). Because the strain rate is proportional to rotational speed, high strain rates are readily obtained (Table 1). Moreover, friction has no effect on the test, as it does in compression testing. The stress state in torsion may represent the typical stress in metalworking processes, but deformation in the torsion test is not an accurate simulation of metalworking processes, because of excessive material reorientation at large strains. Table 1 Torsional rotation rates corresponding to various metalworking operations Operation

von Mises effective (a)

-1

strain rate ( ) , s

Corresponding surface shear strain rate in torsion (

Rotation rate(b),rpm

), s-1

Isothermal forging

10-3

1.73 × 10-3

0.02

Hydraulic press forging

1

1.73

16.5

Extrusion

20

34.6

330.4

Mechanical press forging

50

86.6

827.0

Sheet rolling

200

346.4

3307.9

Wire drawing

(a)

(b)

500

866.0

8269.7

=

Assuming specimen geometry with r/L = 1.0

Because of the above advantages, the torsion test is frequently used to measure the flow stress and the stress-strain curve (flow curve) under hot-working conditions. Figure 1 shows typical flow curves as a function of temperature and strain rate. In the torsion test, measurements are made of the torque, M, to deform the specimen and the angle of twist or number of turns ( = 2 rad per turn). The shear stress on the outer surface of the specimen is given by:

where r is the specimen radius, m is the strain rate sensitivity found from plots of log M versus log at fixed values of , and n is the strain-hardening exponent obtained from the instantaneous slope of log M versus log .

Fig. 1 Flow curves for Waspaloy. (a) Effect of temperature at a fixed effective strain rate of 1 s-1. (b) Effect of strain rate at a fixed test temperature of 1038 °C (1900 °F). Flow softening at the higher temperature is a result of dynamic recrystallization. Source: Ref 4.

The engineering shear strain

and shear strain rate

are given by:

and

where r is the radius of the specimen and L is the gage length. These values of shear stress and shear strain are typically converted to effective stress and effective strain by means of the von Mises yielding criterion (see the article "Introduction to Workability" in this Section):

= and

Figure 2 shows agreement in plots of versus for stress-strain data determined in torsion, tension, and compression. The agreement becomes much better at hot-working temperatures.

Fig. 2 Comparison of effective stress-strain curves determined for type 304L stainless steel in compression, tension, and torsion. (a) Cold- and warm-working temperatures. (b) Hot-working temperatures. Source: Ref 2.

Fracture data from torsion tests are usually reported in terms of the number of twists to failure or the surface fracture strain to failure. Figure 3 shows the relative hot workability of a number of steels and nickel-base superalloys, as indicated by the torsion test. The test identifies the optimal hot-working temperature.

The compression test, in which a cylindrical

specimen is upset into a flat pancake, is usually considered to be a standard bulk workability test. The average stress state during testing is similar to that in many bulk deformation processes, without introducing the problems of necking (in tension) or material reorientation (in torsion). Therefore, a large amount of deformation can be achieved before fracture occurs. The stress state can be varied over wide limits by controlling the barreling of the specimen through variations in geometry and by reducing friction between the specimen ends and the anvil with lubricants. Compression testing has developed into a highly sophisticated test for workability in cold upset forging, and it is a common quality control test in hot-forging operations. Compression forging is a useful method of assessing the frictional conditions in hot working. The principal disadvantage of the compression test is that tests at a constant, true strain rate require special equipment. Compression Test Conditions. Unless the

lubrication at the ends of the specimen is very good, frictional restraint will retard the outward motion of the end face, and part of the end face will Fig. 3 Ductility determined in hot torsion tests. Source: Ref 2. be formed by a folding over of the sides of the original cylinder onto the end face in contact with the platens. The barreling that results introduces a complex stress state, which is beneficial in fracture testing but detrimental when the compression test is used to measure flow stress. The frictional restraint also causes internal inhomogeneity of plastic deformation. Slightly deforming zones develop adjacent to the platens, while severe deformation is concentrated in zones that occupy roughly diagonal positions between opposing edges of the specimen (see Fig. 7 in the article "Introduction to Workability" in this Section). Figure 4 shows the hot upsetting of a cylinder under conditions of poor lubrication in which the platens are cooler than the specimen. The cooling at the ends restricts the flow so that the deformation is concentrated in a central zone, with deadmetal zones forming adjacent to the platen surfaces (Fig. 4a).

Fig. 4 Deformation patterns in nonlubricated, nonisothermal hot forging. (a) Initial barreling. (b) Barreling and folding over. (c) Beginning of end face expansion. Source: Ref 5.

As deformation proceeds, severe inhomogeneity develops, and the growth of the end faces is attributed entirely to the folding over of the sides (Fig. 4b). When the diameter-to-height ratio, D/h, exceeds about 3, expansion of the end faces occurs (Fig. 4c). The conditions described above are extreme and should not be allowed to occur in hot compression testing unless the objective is to simulate cracking under forging conditions. Adequate lubrication cannot improve the situation so that homogeneous deformation occurs; however, with glass lubricants and isothermal conditions, it is possible to conduct hot compression testing without appreciable barreling (Ref 6). Isothermal test conditions can be achieved by using a heated subassembly, such as that shown in Fig. 5, or heated dies that provide isothermal conditions (Ref 8).

Fig. 5 Heated subassembly with specimen in position used to achieve isothermal test conditions. Thermocouple is removed prior to compression. Source: Ref 7.

The true strain rate in a compression test is:

where v is the velocity of the platen and h is the height of the specimen at time t. Because h decreases continuously with time, the velocity must decrease in proportion to (-h) if is to be held constant. In a normal test, if v is held constant, the engineering strain rate will remain constant:

The true strain rate, however, will not be constant. A machine called a cam plastometer can be used to cause the bottom platen to compress the specimen through cam action at a constant true strain rate to a strain limit of = 0.7 (Ref 9). The use of cam plastometers is limited; there probably are not more than ten in existence. However, an essentially constant true strain rate can be achieved on a standard closed-loop servo-controlled testing machine. Strain rates up to 20 s-1 have been achieved (Ref 6, 10). The history of the cam plastometer, the basic principles involved in the technique, and the equipment used are discussed in the article "High Strain Rate Compression Testing" in Mechanical Testing, Volume 8 of ASM Handbook, formerly 9th Edition Metals Handbook. When a constant true strain rate cannot be obtained, the mean strain rate may be adequate. The mean true strain rate, < >, for constant velocity vo, when the specimen is reduced in height from h0 to h, is given by:

Flow Stress in Compression. Ideally, the determination of flow stress in compression should be carried out under

isothermal conditions (no die chilling) at a constant strain rate and with a minimum of friction in order to minimize barreling. These conditions can be met with conventional servohydraulic testing machines. For an essentially homogeneous upsetting test, a cylinder of diameter D0 and initial height h0, will be compressed to height h and spread out to diameter D1 according to the law of constancy of volume:

h0 = D2h If friction can be neglected, the uniaxial compressive stress (flow stress) corresponding to a deformation force P is:

If substantial friction is present, the average pressure, , required to deform the cylinder is greater than the flow stress of the material, 0:

where a is the radius of the cylinder, and is the Coulomb coefficient of friction. The true compressive strain is given by:

The effects of friction and die chilling can be minimized through the use of a long, thin specimen. Therefore, most of the specimen volume is unaffected by the dead-metal zones at the platens. However, this approach is limited, because buckling of the specimen will occur if h/D exceeds about 2. An extrapolation method involves testing cylinders of equal diameters but varying heights so that the D0/h0 ratio ranges from about 0.5 to 3.0 (Ref 11). A specific load is applied to the specimen, the load is removed, and the new height is determined in order to calculate a true strain. Upon relubrication, the specimen is subjected to an increased load, unloaded, and measured. The cycle is then repeated. The same test procedure is followed with each specimen so that the particular load levels are duplicated. The results are illustrated in Fig. 6. For the same load, the actual strain (due to height reduction) is plotted against the D0/h0 ratio for each test cylinder. A line drawn through the points is extrapolated to a value of D0/h0 = 0. This would be the anticipated ratio for a specimen of infinite initial height for which the end effects would be restricted to a small region of the full test height. The true stress corresponding to each of these true strains is given by:

Fig. 6 Extrapolation method to correct for end effects in compressive loading. Source: Ref 11.

Ductility Testing. The basic hot ductility test consists of compressing a series of cylindrical or square specimens to

various thicknesses, or to the same thickness with varying specimen length-to-diameter (length-to-width) ratios. The limit for compression without failure by radial or peripheral cracking is considered to be a measure of workability. This type of test has been widely used in the forging industry. Longitudinal notches are sometimes machined into the specimens before compression, because the notches apparently cause more severe stress concentrations, thus providing a more reliable index of the workability to be expected in a complex forging operation. Plastic Instability in Compression. Several types of plastic instabilities can be developed in the compression test.

The first type is associated with a maximum in the true stress-strain curve. The second type concerns inhomogeneous deformation and shear band formation. Figure 7 shows the type of plastic instability that occurs in some materials in hot compression testing. At certain temperatures and strain rates, some of the typical strengthening mechanisms become unstable. Because the rate of flow softening exceeds the rate of area increase as the specimen is compressed, a maximum results in the flow stress curve.

Fig. 7 Example of compressive flow stress curve showing strain softening.

Analysis of the compression process indicates that the plastic deformation is stable (no maximum in the flow

curve) as long as ( + m) 1, where is the dimensionless work-hardening coefficient, and m is strain rate sensitivity. Both of these material parameters are defined below (Ref 12, 13). A material with a high strain rate sensitivity is more resistant to flow localization in the tension test (necking), but in compression testing, a higher rate sensitivity leads to earlier flow localization.

Flow softening or negative strain hardening can also produce flow localization effects in compression independently of the effects of die chilling or high friction. The constant strain rate, isothermal hot compression test is useful for detecting and predicting flow localization. Nonuniform flow in compression is likely if a flow parameter c exceeds a certain value:

where

and

Figure 8 illustrates the differences in deformation of titanium alloy samples. The specimens in Fig. 8(a) to (c) were deformed at a temperature at which c was high. In Fig. 8(a), = 10-3 s-1 and c = 2. In Fig. 8(b), = 10-1 s-1 and c = 5. In Fig. 8(c), = 10 s-1 and c = 5. However, the specimens in Fig. 8(d) to (f) were deformed at a temperature at which c was less than 0.

Fig. 8 Specimens of Ti-10V-2Fe-3Al from isothermal hot compression tests. (a) to (c) Tested at 704 °C (1300 °F). (d) to (f) Tested at 816 °C (1500 °F). Strain rates were 10-3 s-1 (a, d), 10-1 s-1 (b, e), and 10 s-1 (c, f). Before testing, the alloy had been annealed to yield an equiaxed starting microstructure. Source: Ref 14.

The bend test is useful for assessing the workability of thick sheet and plate. Generally, this test is most applicable to

cold-working operations. Figure 9 shows a plate deformed in three-point bending. The principal stress and strains developed during bending are defined in Fig. 10. The critical parameter is width-to-thickness ratio (w/t). If w/t > 8, bending occurs under plane-strain conditions ( 2 = 0) and 2/ 1 = 0.5. If w/t > 8, the bend ductility is independent of the exact w/t ratio. If w/t < 8, then stress state and bend ductility depend strongly on the width-to-thickness ratio.

Fig. 9 Three-point bend test.

Fig. 10 Schematic of the bend region defining direction of principal stresses and strains.

For pure plastic bending, in which elastic deformation can be ignored, the maximum tensile fiber strain is (Ref 15):

where Ro is the radius of curvature on the outer (tensile) surface and Ri is the radius of curvature on the inner (compressive) surface. When this strain is entered into the stress-strain equation or curve for the material, it gives the flow stress for the material

. Because of the plane-strain condition, the maximum fiber stress is 2 /

.

References cited in this section

1. G.E. Dieter, Mechanical Behavior of Materials Under Tension, in Mechanical Testing, Vol 8, 9th ed., Metals Handbook, American Society for Metals, 1985, p 20-27

2. S.L. Semiatin, G.D. Lahoti, and J.J. Jonas, Application of the Torsion Test to Determine Workability, in Mechanical Testing, Vol 8, 9th ed., Metals Handbook, American Society for Metals, 1985, p 154-184 3. M.J. Luton, Hot Torsion Testing, in Workability Testing Techniques, G.E. Dieter, Ed., American Society for Metals, 1984, p 95-133 4. S. Fulop, K.C. Cadien, M.J. Luton, and H.J. McQueen, J. Test Eval., Vol 5, 1977, p 419 5. J.A. Schey, T.R. Venner, and S.L. Takomana, Shape Changes in the Up-setting of Slender Cylinders, J. Eng. Ind. (Trans. ASME), Vol 104, 1982, p 79 6. G. Fitzsimmons, H.A. Kuhn, and R. Venkateshwar, Deformation and Fracture Testing for Hot Working Processes, J. Met., May 1981, p 11-17 7. J.F. Alder and V.A. Phillips, The Effect of Strain Rate and Temperature on the Resistance of Aluminum, Copper and Steel to Compression, J. Inst. Met., Vol 83, 1954-1955, p 80-86 8. F.J. Gurney and D.J. Abson, Heated Dies for Forging and Friction Studies on a Modified Hydraulic Forge Press, Met. Mat., Vol 7, 1973, p 535 9. J.E. Hockett, The Cam Plastometer, in Mechanical Testing, Vol 8, 9th ed., Metals Handbook, American Society for Metals, 1985, p 193-196 10. J.G. Lenard, Development of an Experimental Facility for Single and Multistage Constant Strain Rate Compression, J. Eng.Mater. Technol. (Trans. ASME), Vol 107, 1985, p 126-131 11. A.B. Watts and H. Ford, On the Basic Yield Stress Curve for a Metal, Proc. Inst. Mech. Eng., Vol 169, 1955, p 1141-1149 12. J.J. Jonas, R.A. Holt, and C.E. Coleman, Plastic Stability in Tension and Compression, Acta Metall., Vol 24, 1976, p 911 13. S.L. Semiatin and J.J. Jonas, Formability and Workability of Metals: Plastic Instability and Flow Localization, American Society for Metals, 1984 14. S.L. Semiatin, Workability in Forging, in Workability Testing, G.E. Dieter, Ed., American Society for Metals, 1984, p 197-247 15. P. Dadras, Stress-Strain Relationships in Bending, in Mechanical Testing, Vol 8, 9th ed., Metals Handbook, American Society for Metals, 1985, p 118-124 Workability Tests George E. Dieter, University of Maryland

Specialized Tests In the plane-strain compression test, the difficulties encountered with bulging and high friction at the platens in

the compression of cylinders can be minimized (Ref 11). As shown in Fig. 11, the specimen is a thin plate or sheet that is compressed across the width of the strip by narrow platens that are wider than the strip. The elastic constraints of the undeformed shoulders of material on each side of the platens prevent extension of the strip in the width dimension; hence the term plane strain.

Deformation occurs in the direction of platen motion and in the direction normal to the length of the platen. To ensure that lateral spread is negligible, the width of the strip should be at least six to ten times the breadth of the platens. To ensure that deformation beneath the platens is essentially homogeneous, the ratio of platen breadth to strip thickness (b/t) should be between 2 and 4 at all times. It may be necessary to change the platens during testing to maintain this condition. True strains of 2 can be achieved by carrying out the test in increments in order to provide good lubrication and to maintain the proper b/t ratio. Although the plane-strain compression test is primarily used to measure flow properties at room temperature, it can also be used for elevated-temperature tests (Ref 16, 17).However, because of its geometry, this test is more applicable to rolling operations than to forging. The true stress and true strain determined from the plane-strain compression test can be expressed as: Fig. 11 Plane-strain compression test.

Because of the stress state associated with plane-strain deformation, the mean pressure on the platens is 15.5% higher in the plane-strain compression test than in uniaxial compression testing. The true stress-strain curve in uniaxial compression ( 0 versus ) can be obtained from the corresponding plane-strain compression curve (p versus pc) by:

and

The partial-width indentation test is a new test for evaluating the workability of metals. It is similar to the plane-

strain compression test, but it does not subject the test specimen to true plane-strain conditions (Ref 18). In this test, a simple slab-shaped specimen is deformed over part of its width by two opposing rectangular anvils having widths smaller than that of the specimen.Upon penetrating the workpiece, the anvils longitudinally displace metal from the center, creating overhangs (ribs) that are subjected to secondary, nearly uniaxial tensile straining. The material ductility under these conditions is indicated by the reduction in the rib height at fracture. The test geometry has been standardized (Fig. 12).

One advantage of this test is that it uses a specimen of simple shape. In addition, ascast materials can be readily tested. One edge of the specimen can contain original surface defects. The test can be conducted hot or cold. Therefore, the partial-width indentation test is suitable not only for determining the intrinsic ductilities of materials but also for evaluating the inhomogeneous aspects of workability. This test has been used to establish the fracture-limit loci for ductile metals (Ref 19). The secondary-tension test, a modification of the partial-

Fig. 12 Partial-width indentation test. L = h/2 wa = 2L; l = 4L.

width indentation test, imposes more severe strain in the rib for testing highly ductile materials. In this test, a hole or a slot is machined in the slab-type specimen adjacent to where the anvils indent the specimen. Preferred dimensions of the hole and slot are given in Fig. 13. With this design, the ribs are sufficiently stretched to ensure fracture in even the most ductile materials. The fracture strain is based on reduction in area where the rib is cut out, h; b so that the fracture area can be photographed or traced on an optical comparator.

Fig. 13 Secondary-tension test showing the geometries of holes and slots L

h; wa

2h; b = h/4; D = h/2.

Ring Compression Test. When a flat ring-shaped specimen is upset in the axial direction, the resulting change in

shape depends only on the amount of compression in the thickness direction and the frictional conditions at the die/ring interfaces. If the interfacial friction were zero, the ring would deform in the same manner as a solid disk, with each element flowing outward radially at a rate proportional to its distance from the center. In the case of small, but finite, interfacial friction, the outside diameter is smaller than in the zero-friction case. If the friction exceeds a critical value, frictional resistance to outward flow becomes so high that some of the ring material flows inward to the center. Measurements of the inside diameters of compressed rings provide a particularly sensitive means of studying interfacial friction, because the inside diameter increases if the friction is low and decreases if the friction is higher (Fig. 14).

Fig. 14 Variation in shape of ring test specimens deformed the same amount under different frictional conditions. Left to right: undeformed specimen; deformed 50%, low friction; deformed 50%, medium friction; deformed 50%, high friction.

The ring test, then, is a compression test with a built-in frictional measurement. Therefore, it is possible to measure the ring dimensions and compute both the friction value and the basic flow stress of the ring material at the strain under the given deformation conditions. Analysis of Ring Compression. The mechanics of the compression of flat ring-shaped specimens between flat dies

have been analyzed using an upper bound plasticity technique (Ref 20, 21). Values of p/ 0 (where p is the average forging pressure on the ring, and 0 is the flow stress of the ring material) can be calculated in terms of ring geometry and the interfacial shear factor, m. In these calculations, neither 0 nor the interfacial shear stress, , appears in terms of independent absolute values, but only as the ratio m (see the article "Introduction to Workability" in this Section). The analysis assumes that this ratio remains constant for a given material and deformation conditions. If the analysis is carried out for a small increment of deformation, 0 and can be assumed to be approximately constant for this increment, and the solution is valid. Therefore, if the shear factor m is constant for the entire operation, the mathematical analysis can be continued in a series of small deformation increments, using the final ring geometry from one increment as the initial geometry for the subsequent increment. As long as the ratio of the interfacial shear stress, , to the material flow stress, 0, remains constant, strain hardening of the ring material during deformation has no effect if the increase in work hardening in any single deformation increment can be neglected. The progressive increase in interfacial shear stress accompanying strain hardening is also immaterial if it can be assumed to be constant over the entire die/ring interface during any one deformation increment. Therefore, the analysis can be justifiably applied to real materials even though it was initially assumed that the material would behave according to the von Mises stress-strain rate laws, provided the assumption of a constant interfacial shear factor, m, is correct. However, it has been shown that a highly strain rate sensitive material requires a different analysis (Ref 22). Based on these assumptions, the plasticity equations have been solved for several ring geometries over a complete range of m values from 0 to unity (Ref 23), as shown in Fig. 15. The friction factor can be determined by measuring the change in internal diameter of the ring.

The ring thickness is usually expressed in relation to the inside and outside diameters. The maximum thickness that can be used while still satisfying the mathematical assumption of thin-specimen conditions varies, depending on the actual friction conditions. Under conditions of maximum friction, the largest usable specimen height is obtained with rings of dimensions in the OD:ID:thickness ratio of 6:3:1. Under conditions of low friction, thicker specimens can be used while still satisfying the above assumption. For normal lubricated conditions, a geometry of 6:3:2 can be used to obtain results of sufficient accuracy for most applications. For experimental conditions in which specimen thicknesses are greater than those permitted by a geometry of 6:3:1 and/or the interface friction is relatively high, the resulting side barreling or bulging must be considered. Analytical treatment of this more complex situation is available in Ref 24. The ring compression test can be used to measure the flow stress under high-strain practical forming conditions. The only instrumentation required is that for measuring the force needed to produce the reduction in height. The change in diameter of the 6:3:1 ring is measured to obtain a value of the ratio p/ 0 by solving the analytical expression for the deformation of the ring or by using computer solutions for the ring (Ref 25). Measurement of the Fig. 15 Theoretical calibration curve for standard ring with area of the ring surface formerly in contact with the an OD:ID:thickness ratio of 6:3:2. die and knowledge of the deformation load facilitate calculation of p and therefore the value of the material flow stress, o, for a given amount of deformation. Repetition of this process with other ring specimens over a range of deformation allows the generation of a complete flow stress-strain curve for a given material under particular temperature and strain rate deformation conditions. Hot Tension Testing. Although necking is a fundamental limitation in tension testing, the tension test is nevertheless

useful for establishing the temperature limits for hot working. The principal advantage of this test for industrial applications is that it clearly establishes maximum and minimum hot-working temperatures (Ref 26). Most commercial hot tensile testing is done with a Gleeble unit, which is a high strain rate, high-temperature testing machine (Ref 27). A solid buttonhead specimen that has a reduced diameter of 6.4 mm (0.250 in.) and an overall length of 89 mm (3.5 in.) is held horizontally by water-cooled copper jaws (grips), through which electric power is introduced to resistance heat the test specimen (Fig. 16). Specimen temperature is monitored by a thermocouple welded to the specimen surface at its midlength. The thermocouple, with a function generator, controls the heat fed into the specimen according to a programmed cycle. Therefore, a specimen can be tested under time-temperature conditions that simulate hot-working sequences.

Fig. 16 The Gleeble test unit used for hot tension and compression testing. (a) Specimen in grips showing attached thermocouple wires and linear variable differential transformer for measuring strain. (b) Close-up of a compression test specimen. Courtesy of Duffers Scientific, Inc.

The specimen is loaded by a pneumatic-hydraulic system. The load can be applied at any desired time in the thermal cycle. Temperature, load, and crosshead displacement are measured as a function of time. In the Gleeble test, the crosshead speed can be maintained constant throughout the test, but the true strain rate decreases until necking occurs, according to the relationship:

When the specimen necks, the strain rate increases suddenly in the deforming region, because deformation is concentrated in a narrow zone. Although this variable strain rate history introduces some uncertainty into the determination of strength and ductility values, it does not negate the utility of the hot tension test. Moreover, a procedure has been developed that corrects for the change in strain rate with strain so that stress-strain curves can be constructed (Ref 28). The percent reduction in area is the primary result obtained from the hot tension test. This measure of ductility is used to assess the ability of the material to withstand crack propagation. Reduction in area adequately detects small ductility variations in materials caused by composition or processing when the material is of low-to-moderate ductility. It does not reveal small ductility variations in materials of very high ductility. A general qualitative rating scale between reduction in area and workability is given in Table 2. This correlation was originally based on superalloys. In addition to ductility measurement, the ultimate tensile strength can be determined with the Gleeble test. This gives a measure of the force required to deform the material. Table 2 Qualitative hot-workability ratings for specialty steels and superalloys Hot tensile reduction in area(a), %

Expected alloy behavior under normal hot reductions in open-die forging or rolling

Remarks regarding alloy hot-working practice

70

Superior hot workability, rare cracks. Ductile ruptures can occur if strength is too low.

Rolled or press forged with heavier reductions and higher strain rates than normal if alloy strength is sufficiently high to prevent ductile ruptures

Source: Ref 26 (a) Ratings apply for Gleeble tension testing of 6.4-mm (0.250 in.) diam specimens with 25-mm (1 in.) head separation.

Hot Tension Test Procedure Variations. Two variations of the hot tension test can be used to establish the

temperature limits of hot working: on-heating tests and on-cooling tests. The on-heating test method is used for a material for which little or no hot-working information is available. The specimens are resistance heated to the test temperature, held for 1 to 10 min, and pulled to fracture at a crosshead rate approximating the strain rate of plant practice. The reduction in area versus test temperature obtained by the on-heating testing of a heat-resistant alloy is shown in Fig. 17. The optimal reheat temperature for working lies between the peak ductility temperature and the zero-ductility temperature. The test clearly distinguishes between ingots prepared by electroslag remelting (ESR) and vacuum arc remelting (VAR) practices. The on-cooling test procedure is used to establish the optimal preheat temperature in this range. The objective is to determine which hot-working temperature provides the highest ductility over the broadest temperature range without risking permanent damage to the material from overheating. Unmachined specimen blanks are heat treated in a furnace at a given preheat temperature and duration to duplicate a furnace soak commensurate with the workpiece size and the hot-working operation. Samples are water quenched from the soak temperature to retain the hightemperature structure. After machining, tensile specimens are heated to the preheat temperature in the Gleeble unit and held for 1 to 10 min to dissolve any phases that may have precipitated during cooling. Specimens are then cooled to a series of temperatures below the preheat temperatures at 28 to 55 °C (50 to 100 °F) intervals, held 5 s at the test temperature, and pulled to fracture at the appropriate head speed. Data obtained from on-cooling tests conducted on three test specimens that were subjected to varying preheat temperatures are shown in Fig. 18. A preheat temperature of 1205 °C (2200 °F) was selected as optimal in this example, because it produced a slightly higher and rather broad band of high ductility. The minimum hot-working temperature was established as the temperature at which the reduction in area decreases to the 50% level for typical workpiece reductions. Fig. 17 Reduction in area versus test temperature obtained by hot tension testing on heating. Specimens were heated to the test temperature, held 5 min, and pulled to fracture.

Fig. 18 Reduction in area versus testing temperature for Unitemp HN (ESR) generated by testing on cooling. Specimen blanks were furnace soaked 2 h at the preheat temperatures. The specimens were then heated to the preheat temperatures in the Gleeble unit, held 5 min, cooled to the test temperature, held 5 s, and pulled to fracture.

On-cooling hot tension testing is useful, because the brief hold times for on-heating tests may not develop a grain size representative of that temperature, or they may be insufficient to dissolve or precipitate a phase that will occur during an actual furnace soak prior to hot working. In addition, most industrial hot-working operations are performed while the workpiece temperature cools slowly. On-cooling tests also indicate how closely the zero-ductility temperature can be approached before hot ductility is severely reduced.

References cited in this section

11. A.B. Watts and H. Ford, On the Basic Yield Stress Curve for a Metal, Proc. Inst. Mech. Eng., Vol 169, 1955, p 1141-1149 16. J.A. Bailey, The Plane Strain Forging of Aluminum at Low Strain Rates and Elevated Temperatures, Int. J. Mech. Sci., Vol 11, 1969, p 491 17. O. Pawelski, U. Rudiger, and R. Kaspar, The Hot Deformation Simulator, Stahl Eisen, Vol 98, 1978, p 181189 18. S.M. Woodall and J.A. Schey, Development of New Workability Test Techniques, J. Mech.Work. Technol., Vol 2, 1979, p 367-384 19. S.M. Woodall and J.A. Schey, Determination of Ductility for Bulk Deformation, in Formability Topics-Metallic Materials, STP 647, American Society for Testing and Materials, 1978, p 191-205 20. B. Avitzur, Metal Forming: Processes and Analysis, McGraw-Hill, 1968 21. B. Avitzur and C.J. Van Tyne, Ring Forming: An Upper Bound Approach, J.Eng. Ind. (Trans. ASME), Vol 104, 1982, p 231-252 22. G. Garmong, N.E. Paton, J.C. Chesnut, and L.F. Necarez, An Evaluation of the Ring Test for Strain-Rate Sensitive Materials, Metall. Trans. A, Vol 8A, 1977, p 2026, 2027 23. A.T. Male and V. DePierre, The Validity of Mathematical Solutions for Determining Friction From the Ring Compression Test, J. Lubr. Technol. (Trans. ASME), Vol 92, 1970, p 389-397 24. V. DePierre, F.J. Gurney, and A.T. Male, "Mathematical Calibration of the Ring Test With Bulge Formation," Technical Report AFML-TR-37, U.S. Air Force Materials Laboratory, March 1972

25. G. Saul, A.T. Male, and V. DePierre, "A New Method for the Determination of Material Flow Stress Values Under Metalworking Conditions," Technical Report AFML-TR-70-19, U.S. Air Force Materials Laboratory, Jan 1970 26. R.E. Bailey, R.R. Shiring, and H.L. Black, Hot Tension Testing, in Workability Testing Techniques, G.E. Dieter, Ed., American Society for Metals, 1984 27. E.F. Nippes, W.F. Savage, B.J. Bastian, H.F. Mason, and R.M. Curran, An Investigation of the Hot Ductility of High Temperature Alloys, Weld. J., Vol 34, April 1955, p 183-196 28. R.L. Plaut and C.M. Sellars, Analysis of Hot Tension Test Data to Obtain Stress-Strain Curves to High Strains, J. Test Eval., Vol 13, 1985, p 39-45 Workability Tests George E. Dieter, University of Maryland

Forgeability Tests Basically, all forging processes consist of the compressive deformation of a metal workpiece between a pair of dies (Ref 14). The two broad categories of forging processes are open-die and closed-die modes. The simplest open-die forging operation is the upsetting of a cylindrical billet between two flat dies. The compression test is a small-scale prototype of this process. As the metal flows laterally between the advancing die surfaces, there is less deformation at the die interfaces (because of the friction forces) than at the midheight plane. Therefore, barreling occurs on the sides of the upset cylinder. Generally, metal flows most easily toward the nearest free surface because this path presents the least friction. Closed-die forging is done in closed or impression dies that impart a well-defined shape to the workpiece. The degree of lateral constraint varies with the shape of the dies and the design of the peripheral areas where flash is formed, as well as with the same factors that influence metal flow in open-die forging (amount of reduction, frictional boundary conditions, and heat transfer between the dies and the workpiece). Because forging is a complex process, a single workability test cannot be relied on to determine forgeability. However, several testing techniques have been developed for predicting forgeability, depending on alloy type, microstructure, die geometry, and process variables. This section will summarize some of the common tests for determining workability in open-die and closed-die forging. Wedge-Forging Test. In this test, a wedge-shaped piece of metal is machined from a cast ingot or wrought billet and

forged between flat, parallel dies (Fig. 19). The dimensions of the wedge must be selected so that a representative structure of the ingot is tested. Coarse-grain materials require larger specimens than fine-grain materials. The wedgeforging test is a gradient test in which the degree of deformation varies from a large amount at the thick end (h2) to a small amount or no deformation at the thin end (h1). The specimen should be used on the actual forging equipment in which production will occur to allow for the effects of deformation velocity and die chill on workability.

Tests can be made at a series of preheat temperatures, beginning at about nine-tenths of the solidus temperature or the incipient melting temperature. After testing at each temperature, the deformation that causes cracking can be established. In addition, the extent of recrystallization as a function of strain and temperature can be determined by performing metallographic examination in the direction of the strain gradient. consists of compressing a cylindrical bar between flat, parallel dies where the axis of the cylinder is parallel to the surfaces of the dies. Because the cylinder is compressed on its side, this testing procedure is termed sidepressing. This test is sensitive to surface-related cracking and to the general unsoundness of the bar, because high tensile stresses are created at the center of the cylinder Fig. 19 Specimens for the wedge test. (a) As- (Fig. 20). The

sidepressing

test

machined specimen. (b) Specimen after forging.

Fig. 20 Effects of billet shape and degree of enclosure on stress state in forging with good lubrication and no chilling. Source: Ref 29.

For a cylindrical bar deformed against flat dies, the tensile stress is greatest at the start of deformation and decreases as the bar assumes more of a rectangular cross section. As shown in Fig. 20, the degree of tensile stress can be reduced at the outset of the tests by changing from flat dies to curved dies that support the bar around part of its circumference.

The typical sidepressing test is conducted with unconstrained ends. In this case, failure occurs by ductile fracture on the expanding end faces. If the bar is constrained to deform in plane strain by preventing the ends from expanding, deformation will be in pure shear, and cracking will be less likely. Plane-strain conditions can be achieved if the ends are blocked from longitudinal expansion by machining a channel or cavity into the lower die block. The notched-bar upset test is similar to the conventional upset test, except that axial notches are machined into the

test specimens (Ref 30). The notched-bar test is used with materials of marginal forgeability for which the standard upset test may indicate an erroneously high degree of workability. The introduction of notches produces high local stresses that induce fracture. The high levels of tensile stress in the test are believed to be more typical of those occurring in actual forging operations. Test specimens are prepared by longitudinally quartering a forging billet, thus exposing center material along one corner of each test specimen (Fig. 21). Notches with 1.0 or 0.25 mm (0.04 or 0.01 in.) radii are machined into the faces as shown. A weld button is frequently placed on one corner to identify the center and surface material of alloys that are difficult to forge because of segregation.

Fig. 21 Method of preparing specimens for notched-bar upset forgeability test. Source: Ref 30.

Specimens are heated to predetermined temperatures and upset about 75%. The specimen is oriented with the grooves (notches) in the vertical direction. Because of the stress concentration effect, ruptures are most likely to occur in the notched areas. These ruptures can be classified according to the rating system shown in Fig. 22. A rating of 0 indicates that no ruptures are observed, and higher numbers indicate an increasing frequency and depth of rupture.

Fig. 22 Suggested rating system for notched-bar upset test specimens that exhibit progressively poorer forgeability. A rating of 0 indicates freedom from ruptures in the notched area. Source: Ref 30.

Figure 23 shows roll-forged rings made from two heats of type 403 stainless steel. The ring shown in Fig. 23(a) came from a billet with a notched-bar forgeability rating of 0. The billet shown in Fig. 23(b) had a forgeability rating of 4.

Fig. 23 Rolled rings made from two heats of type 403 stainless steel exhibiting different forgeability ratings in notched-bar upsetting tests. (a) Forgeability rating is 0. (b) Forgeability rating is 4. Courtesy of Ladish Company.

Truncated Cone Indentation Test. This test involves the indentation of a cylindrical specimen by a conical tool

(Fig. 24). As a result of the indentation, cracking is made to occur beneath the surface of the testpiece at the tool/material interface. The reduction (measured at the specimen axis) at which cracking occurs can be used to compare the workability

of different materials. Alternatively, the reduction (stroke) at which a fixed crack width is produced or the width of the crack at a given reduction can be used as a measure of workability.

Fig. 24 Relationship between crack width and stroke in truncated cone indentation test for workability of various steels at cold-forging temperatures.

The truncated cone was developed as a test that minimizes the effects of surface flaws and the variability they produce in workability (Ref 31). This test has been primarily used in cold forging.

References cited in this section

14. S.L. Semiatin, Workability in Forging, in Workability Testing, G.E. Dieter, Ed., American Society for Metals, 1984, p 197-247 29. A.L. Hoffmanner, "Plasticity Theory as Applied to Forging of Titanium Alloys," Paper presented at the Symposium on the Thermal-Mechanical Treatment of Metals, London, May 1970 30. R.P. Daykin, Ladish Company, unpublished research, 1951 31. T. Okamoto, T. Fukuda, and H. Hagita, Material Fracture in Cold Forging--Systematic Classification of Working Methods and Types of Cracking in Cold Forging, Sumitomo Search, No. 9, May 1973, p 46; Source Book on Cold Forming, American Society for Metals, 1975, p 216-226

Workability Tests George E. Dieter, University of Maryland

Tests for Flow Localization Complex forgings frequently develop regions of highly localized deformation. Shear bands may span the entire cross section of a forging and, in extreme cases, produce shear cracking. Flow localization can arise from constrained deformation due to die chill or high friction. However, flow localization can also occur in the absence of these effects if the metal undergoes flow softening or negative strain hardening. The simplest workability test for detecting the influence of heat transfer (die chilling) on flow localization is the nonisothermal upset test, in which the dies are much colder than the workpiece. Figure 25 illustrates zones of flow localization made visible by sectioning and metallographic preparation.

Fig. 25 Axial cross sections of specimens of Ti-6Al-2Sn-4Zr-2Mo-0.1Si with an equiaxed starting microstructure. Specimens were nonisothermally upset at 954 °C (1749 °F) to 50% reduction in a mechanical press ( 30 s-1) between dies at 191 °C (376 °F). Dwell times on the dies prior to deformation were (a) 0 s and (b) 5 s. Source: Ref 32.

The sidepressing test conducted in a nonisothermal manner can also be used to detect flow localization. Several test specimens are sidepressed between flat dies at several workpiece temperatures, die temperatures, and working speeds. The formation of shear bands is determined by metallography (Fig. 26). Flow localization by shear band formation is more likely in the sidepressing test than in the upset test. This is due to the absence of a well-defined axisymmetric chill zone. In the sidepressing of round bars, the contact area starts out at zero and builds up slowly with deformation. In addition, because the deformation is basically plane strain, surfaces of zero extension are present, along which block shearing can initiate and propagate. These are natural surfaces along which shear strain can concentrate into shear bands.

Fig. 26 Transverse metallographic sections of specimens of Ti-6Al-2Sn-4Zr-2Mo-0.1Si with an equiaxed structure. Specimens were nonisothermally sidepressed with zero dwell time in a mechanical press ( 30 s-1) between dies heated to 191 °C (376 °F). Specimen preheat temperatures (Ts) and percent reductions (R), relative to the initial specimen diameter, were as follows: (a) Ts: 913 °C (1675 °F); R: 14%. (b) Ts: 913 °C (1675 °F); R: 54%. (c) Ts: 913 °C (1675 °F); R: 77%. (d) Ts: 982 °C (1800 °F); R: 21 %. (e) Ts: 982 °C (1800 °F); R: 57%. (f) Ts: 982 °C (1800 °F); R: 79%. Source: Ref 32.

Testing to evaluate material susceptibility to localized deformation can also involve the use of a cylindrical upset specimen with a reduced gage section (Ref 33), as shown in Fig. 27. The ability of the material to distribute deformation (Fig. 28) is measured by an empirical parameter--percent distributed gage volume (DGV). The larger the DGV percentage, the greater the penetration of the deformation into the heavy ends of the specimen and the greater the ability of the material to distribute deformation. Figure 29 shows the metallographic appearance of the condition with distributed flow (Fig. 29a) and concentrated deformation (Fig. 29b). Figure 30 illustrates that the DGV percentage is a sensitive parameter for detecting flow localization.

Fig. 27 Shape and dimensions of cylindrical compression specimen with a reduced gage section.

Fig. 28 Schematic of specimen cross sections showing the relative amount of gage volume penetration (DGV) into the specimen ends for two different deformation behaviors. (a) Distributed deformation. (b) Concentrated deformation. Source: Ref 33

Fig. 29 Light micrographs showing variations in flow line contours and gage penetration into the specimen ends. After press forging at 650 °C (1200 °F) (a), 815 °C (1500 °F) (b), and 870 °C (1600 °F) (c). Etched in oxalic acid

Fig. 30 Variation in DGV percentage for pressed specimens of Ti-6Al-4V as a function of upset temperature. Closed circles indicate starting microstructures of globular . Open circles indicate starting microstructure of acicular . Source: Ref 33

References cited in this section

32. S.L. Semiatin and G.D. Lahoti, The Occurrence of Shear Bands in Nonisothermal, Hot Forging of Ti-6Al2Sn-4Zr-2Mo-0.1Si, Metall.Trans. A, Vol 14A, 1983, p 105 33. M.C. Mataya and G. Krauss, A Test to Evaluate Flow Localization During Forging, J. Appl. Metalwork., Vol 2, 1981, p 28-37 Workability Tests George E. Dieter, University of Maryland

Forging Defects Cracking in Cold Forging. The types of cracks that develop in cold forging by upsetting-type processes are discussed

in Ref 31. The various geometric forms of cold forging are shown in Fig. 31. The classification of cracks is given in Fig. 32. Table 3 provides a detailed description of each type of crack. Table 3 Characteristics of cracks and crack growth mechanism Cracking type

No. of working method in the chart (see Fig. 31)

Characteristics of cracking and estimation of crack growth mechanism

00, 01, 11, 02, 21, 22, 08, 87, 07, 77, 09, 03, 33, 04, 44,

External cracking that appears at midheight of side surface of the specimen in the upsetting; two types of cracks, longitudinal and oblique, occur according to the degree of end constraint