ASM Metals HandBook Vol. 8 - Mechanical Testing and Evaluation

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Drop Tower Compression Test

The drop tower compression test, as the name implies, uses a falling weight to provide a compressive load to

the specimen. The test technique has the capability to generate high loads at medium strain rates, which cannot

be readily obtained by servohydraulic load frames or cam plastometers. Load capacities of 900 kN (100 tons)

have been demonstrated for this technique, with test durations from 0.1 to 20 μs resulting in loading rates as

high as 1 GN/s (115 × 10

tons/s).

The drop tower compression test has been used to measure compressive fracture strengths, to determine the

compressive stress-strain behavior of material at medium strain rates, and to evaluate the dimensional stability

of components subjected to impact compressive loads. Two general approaches can be used for drop tower

compression testing. Fracture or compressive stress-strain tests require that the strain rate does not vary

significantly during the test. This is accomplished in the drop tower test by ensuring that the available energy of

the crosshead (weight times height) is no less than three times (preferably an order of magnitude) greater than

the energy absorbed by the specimen and the test system.

The second test approach is used to evaluate the effect of subcritical, or low blow, loadings on the test specimen

or components. In this case, the available energy of the crosshead is selected to generate a specific load. During

the test, the crosshead comes to a stop and then rebounds. To avoid secondary loadings, a device for removing

the specimen or catching the crosshead on the rebound must be used.

Limitations. The drop tower compression test is neither a constant-displacement rate nor a constant-loading rate

test. The rate and form of the compressive loading depend on the specimen and test system compliances, as

well as the impact velocity and available energy of the falling weight. Test conditions must be determined by

trial and error or from empirically derived parameters. The test velocities and, therefore, the loading and strain

rates are limited by both the response time of the instrumentation and the inertia loading of the system. In many

cases, it is the latter that limits the test velocity. In general, it is better to drop a larger weight from a lower

height than the converse when energy requirements are a consideration. The lower velocity will reduce the

inertia loading and still generate comparable maximum loads.

Specimen displacement and energy calculations, which are commonly made for instrumented impact (i.e.,

Charpy) tests, are based on measuring the momentum impulse acting on the falling weight and calculating the

resultant temporal impact velocity. These calculated values can be highly inaccurate for the drop tower

compression test because deformations in the systems can be greater than the specimen deflections.

Equipment. A typical drop tower compression system is shown in Fig. 3. The test system consists of a drop

tower with a massive foundation, a dynamic compression fixture, and stop blocks. The 900 kN (100 ton)

capacity system utilizes a commercially available drop tower with an adjustable crosshead weight of 225 to

1000 kg (500–2200 lb) and a maximum drop height of 1.5 m (5 ft). A 12,000 kg (26,500 lb) steel-reinforced

concrete foundation is used to provide maximum rigidity to the drop tower.

Fig. 3 Drop tower with compression fixture and stop blocks in place

The dynamic compression fixture (Fig. 4) transfers the impact load from the crosshead to the specimen through

a shaft. Proper alignment of the shaft with the specimen axis is maintained throughout the test by linear ball

bearings. Small nonparallelisms in the specimen ends are accommodated by a spherical seat beneath the lower

platen. The impact load is sensed by a load cell located directly below the spherical seat. Stop blocks arrest the

falling crosshead to avoid continued loading of the dynamic compression fixture after the shaft has reached its

maximum travel. The development of the dynamic compression fixture is described in Ref 16.

Fig. 4 Dynamic compression test fixture

A principal consideration in the design of the dynamic compression system is that the entire load train

compliance must be kept quite low to achieve high loads. Gaps, mismatches, and foreign material at any of the

interfaces in the load train adversely alter the rate and form of the impact loading. A second consideration is to

minimize the mass of the fixture between the specimen and the centerline of the load cell. Acceleration of this

mass to the test velocity causes inertia loads (Ref 17) that can obscure the actual mechanical loading of the

specimen. The control of inertia effects is described in Ref 18.

Instrumentation. Equipment designed specifically for instrumented impact testing is used to acquire the load-

versus-time data for the drop tower compression test. Minimum instrumentation requirements are a transducer

to sense the specimen loading, the associated signal conditioning equipment, and recording instruments to

provide a permanent record of the load-sensing transducer. The integrated system must be capable of sampling

rates as fast as 35 kHz. The specific requirements for the load instrumentation, as well as the effect of limiting

the frequency response, are described in Ref 17 and 19.

Specimen Requirements. Specific specimen requirements have not been developed for the drop tower

compression test. However, several specimen configurations based on right-circular cylinders and lengths

greater than two times their diameter have been used. As with all compression tests, the end constraint of the

specimen due to friction is a primary concern. Hardened steel and modulus-matched inserts have been used

between the platens and the specimen to minimize the constraint.

The Hopkinson Bar

Development of test techniques based on the Hopkinson bar has led to significant advances in high strain rate

testing capabilities. These techniques yield the highest possible strain rates in a uniaxial compression test under

uniform deformation conditions. In addition, the determination of stress within the deforming specimen is made

without use of a load cell, and the measurement of strain is made without directly monitoring the specimen

length.

The Hopkinson bar derives its name from its developer, who in 1914 used a long elastic bar to study the

pressures produced by the impact of a bullet or by the detonation of an explosive (Ref 20). In designing this

experiment, Hopkinson recognized that as long as the pressure bar remains elastic, the displacements in the

pressure bar are directly related to the stresses and that the length of the wave in the bar is related to the

duration of the impact through the velocity of sound in the bar, a quantity that is well known. A significant

feature of this work is that the pressures were estimated by measuring the momentum (with a ballistic

pendulum) acquired by a small section of the bar placed in contact with the bar at the far end. Further

developments in the experimental techniques occurred a few decades after these original experiments, when

Davies (Ref 21) and Kolsky (Ref 22) designed condensers to measure displacements in the pressure bars.

Kolsky also introduced the split-Hopkinson pressure bar technique, in which the specimen is sandwiched

between two pressure bars. He demonstrated how stress and strain within the deforming specimen are related to

displacements in the pressure bars. Because of these contributions, the split-Hopkinson pressure bar is often

referred as the Kolsky bar.

There are two basic configurations for Hopkinson pressure bar testing: the split-Hopkinson pressure bar (Fig.

5a) and the single pressure bar configuration (Fig. 5b). The basic principles of the two techniques are similar,

and several investigators have used the single pressure bar configuration. The most widely used method is the

split-Hopkinson pressure bar method, which consists of two elastic pressure bars that sandwich the specimen

between them. Upon impact of a striker bar on an incident bar, an elastic compressive wave is generated within

the incident bar, and the time-dependent strain in the pressure bar is measured at strain gage A at the midpoint

of the incident bar (Fig. 5a). At the incident bar/specimen interface, the wave is partially reflected and partially

transmitted into the specimen. A portion of the incident wave is reflected back along the incident bar as a

tensile wave. This reflected strain is measured by strain gage A. Strain measurements are also taken on the

output bar with strain gage B. These strain measurements on the pressure bars are used to determine the stress-

strain behavior of the specimen if two basic conditions are met. First, the wave propagation within the pressure

bars must be one-dimensional. Secondly, the specimen must deform uniformly. Under these two conditions,

stress-strain behavior of the specimen can be determined as described below. More details are also in the

article“Classic Split-Hopkinson Pressure Bar Testing” in this Volume.

Fig. 5 Typical configurations of the Hopkinson bar. (a) Split-Hopkinson pressure bar with specimen

sandwiched between two long elastic pressure bars, each of which is instrumented at its midpoint with a

strain gage. (b) Single pressure bar with striker bar impacting the specimen directly. With this

configuration, the motion of the striker bar or displacement within the specimen must be monitored

independently.

Relating Hopkinson Bar Strain Gage Measurements to Specimen Stress-Strain Behavior. The strain rate in the

deforming specimen is:

(Eq 8)

where V

and V

are the velocities at the incident bar/specimen and specimen/output bar interfaces, respectively.

The velocity V

is the product of the longitudinal sound velocity, C

, in the pressure bar and the total strain at

the incident bar/specimen interface, which is ε

-ε

. Similarly, the velocity V

is equal to C

. In this

development the incident and transmitted strains (ε

, ε

) and ε are all compressive strains, but are considered

positive, whereas the reflected strain (ε

) represents a tensile strain and is negative. Replacing V

and V

in Eq 8

with these expressions yields:

(Eq 9)

The average stress on the specimen is:

(Eq 10)

where P

and P

are the forces at the incident bar/specimen and specimen/output bar interfaces, respectively,

and A is the instantaneous cross-sectional area of the specimen. At the incident bar/specimen interface, the force

is:

(t) = E[ε

(t) + ε

(t)]A

(Eq 11)

where E is Young's modulus and A

is the cross-sectional area of the bar. Likewise, the force at the

specimen/output bar interface is:

(t) = Eε

(t)A

(Eq 12)

Combining Eq 11 and 12 with Eq 10 yields:

(Eq 13)

When the specimen is deforming uniformly, the stress at the incident bar/specimen interface equals that at the

specimen/output bar interface and from Eq 11 and 12.

(t) + ε

(t) = ε

(t)

(Eq 14)

from which Eq 9 and 13 may be simplified:

(Eq 15)

(Eq 16)

Thus, the stress-strain behavior of the specimen is determined simply by measurements made on the elastic

pressure bars in a split-Hopkinson pressure bar test. The analysis applies as well to the single bar configuration

shown in Fig. 5(b). However, because the striker bar impacts directly with the specimen in this test, the strain

(t), which now represents the wave generated in the striker bar, cannot be measured with a strain gage, and

another measurement technique is required, such as with high-speed photography.

As previously noted, the above equations relating strain gage measurement to stress-strain behavior in the

deforming specimen require that two important conditions be met. The first is that wave propagation within the

pressure bars must be one-dimensional if surface displacement measurements are used to represent the axial

displacement over the entire cross-sectional area of the pressure bar. The second condition is that the specimen

must deform uniformly; this is opposed by both radial and longitudinal inertia and by frictional constraint at the

specimen/pressure bar interfaces. These two conditions are discussed in more detail in the article “Classic Split-

Hopkinson Pressure Bar Testing” in this Volume.

Rod Impact (Taylor) Testing

The rod impact test is based on the Taylor test for measuring the dynamic yield strength (Ref 23). The

technique has been improved so that the entire stress-strain flow curve can be determined at high strain rates

(around 10

or 10

-1

) and large plastic strains (50–150%) for materials at ambient or elevated temperatures.

The availability of this technique has been useful in determining constitutive behavior at high strain rates and

the evaluation of dynamic compressive fracture.

The classic Taylor test (Ref 23, 24) involves a cylindrical specimen rod that impacts a “rigid” plate (Fig. 6a).

The plastic deformation at the impact end shortens the rod, and the fractional change in rod length can, by one-

dimensional rigid-plastic analysis, be related to the dynamic yield strength. This relationship was shown to be

independent of both the rod aspect ratio and the impact velocity for a wide variety of materials, including

copper, lead, paraffin wax, and various steels.

Fig. 6 Two-rod impact configurations. (a) Classic Taylor test. (b) Symmetric rod impact test

Although this method is appealing in its simplicity, the Taylor test received only moderate interest. Twenty-five

years after the original Taylor tests, the use of two-dimensional wave propagation codes enabled a better

understanding of and renewed interest in this technique. In 1972, Wilkins and Guinan (Ref 25), using a two-

dimensional finite difference code and an elastic-plastic model with work hardening, were able to correctly

simulate the final shapes and final lengths of Taylor test specimens of several metallic alloys at ambient

temperatures. Their results showed good correlation between the dynamic yield strength and the fractional

change in rod length for a wide range of impact velocities and rod aspect ratios, thus confirming many of the

Taylor/Whiffin conclusions (Ref 23, 24).

Further improvements have lead to two types of tests: symmetric rod impact tests and asymmetric rod impact

tests. The advantage of the asymmetric rod impact test is that elevated temperature tests can be performed

without heating a moving projectile, as required with the symmetric rod impact test. These types of tests are

used in evaluation of constitutive models (Ref 26) and dynamic deformation and fracture behavior (Ref 27, 28).

Determination of the dynamic flow curve of a material from a rod impact test is made by computationally

simulating the experiment with a two-dimensional wave propagation computer code. The flow parameters are

varied until the computed profiles agree with those determined experimentally at various times during the

deformation history.

Symmetric Rod Impact Test. In the early 1980s, Erlich et al. (Ref 29) implemented two major modifications to

the classic Taylor technique as a method for obtaining the entire stress-strain flow curve for materials

undergoing high strain rate compressive and shear loading to large plastic strains. The first was to use ultrahigh-

speed photography to monitor the deformation history of the specimen rod. This allows intermediate as well as

final deformation states to be compared with two-dimensional computer simulations, thus improving the

reliability of the flow curve determination.

The second modification was to replace the rigid plate with another rod of the same geometry and material as

the impacting rod (Fig. 6b). This arrangement, referred to as the “symmetric rod impact” technique, allows the

impacting ends of the two specimen rods to deform together symmetrically, thus eliminating boundary

condition uncertainties in the analysis that arise from the unknown friction conditions at the rod/plate interface

and from the deformation of the rigid plate adjacent to that interface. By using the symmetric rod impact

technique, dynamic flow curves at ambient temperature were obtained for 6061-T6 aluminum (Ref 29), a

titanium alloy, and 4340 steel over a wide range of initial hardnesses (Ref 30).

A typical experimental arrangement for the symmetric rod impact test at ambient temperature is shown in Fig.

7. The specimen rods are identical right circular cylinders, the ends of which are machined flat and parallel to

within about 0.01 mm (0.0004 in.). The impacting rod is mounted on the front end of a projectile, which is

accelerated by expanding helium in a gas gun. Projectile velocity is recorded by a series of contact pins located

near the end of the gun barrel. The stationary rod is held in place by six ceramic fingers attached to a target-

mounting fixture, which in turn is affixed to an alignment plate at the muzzle of the gun. The position of the

latter rod can be adjusted by rotating the threaded bars into which the ceramic fingers are inserted. Alignment

of the two rods is critical to ensure that the impacting ends are parallel and coaxial.

Fig. 7 Symmetric rod impact tests at ambient temperature

The specimen rods are backlit by a quick-pulse, high-intensity light source (xenon flash tube or exploding

bridge wires) triggered just before impact. The silhouettes of the deforming rods are recorded by a high-speed

framing camera at framing rates between one half and one million frames per second. After deformation is

complete (about 30–40 μs after impact), the specimen rods fly into a recovery pipe filled with rags or other

energy-absorbing materials that minimize additional deformation. The pipe is sufficiently narrow to prevent the

projectile from entering the specimen and impacting it again. The recovered rods are then sectioned along the

axis and examined metallographically to ascertain the extent of internal damage.

The impact velocity must be low enough to suppress the formation of tensile voids (which can occur at early

times by the focusing of the radial release waves on the rod axis) or shear bands (which can occur at later times

as a result of large plastic deformation near the impact end). Although a small amount of incipient damage can

be tolerated, any significant amount of damage can affect the shape of the deforming rod profiles.

Asymmetric Rod Impact Test. As mentioned previously, one advantage of the asymmetric rod impact test is

that elevated temperature tests can be performed without heating a moving projectile, as required with the

symmetric rod impact test. In this variation, a rigid plate is launched into a stationary specimen rod that is

preheated to the desired temperature. An elevated-temperature test with the asymmetric rod impact technique

only requires that the target specimen be preheated to the desired test temperature. The specimen is preheated

with three infrared line heaters. Radiation from each linear filament is focused onto the specimen by an

elliptical reflector. In 1982, Gust (Ref 31) used a reverse ballistics variation of the classic Taylor test to measure

fractional changes in the length of various metallic rods at initial temperatures up to about 1000 °C (1830 °F). If

frictional effects in the impact area are not a significant factor for a given specimen material and target, the

asymmetric rod impact technique may be used to determine dynamic flow curves at ambient or elevated

temperatures.

References cited in this section

4. E. Orowan, The Calculation of Roll Pressure in Hot and Cold Flat Rolling, Proc. Inst. Mech. Eng., Vol

150, 1943, p 143

5. N. Loizou and R.B. Sims, The Yield Stress of Pure Lead in Compression, J. Mech. Phys. Solids, Vol 1,

1953, p 234

6. 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

7. P.M. Cook, True Stress-Strain Curves for Steel in Compression at High Temperatures and Strain Rates

for Application to the Calculation of Load and Torque in Hot Rolling, Conf. on The Properties of

Materials at High Rates of Strain, Institution of Mechanical Engineers, 1957, p 86

8. R.R. Arnold and R.J. Parker, Resistance to Deformation of Aluminum and Some Aluminum Alloys, J.

Inst. Met., Vol 88, 1959–1960, p 255

9. J.A. Bailey and A.R.E. Singer, A Plane-Strain Cam Plastometer for Use in Metal-Working Studies, J.

Inst. Met., Vol 92, 1963–1964, p 288

10. J.E. Hockett, Compression Testing at Constant True Strain Rates, Proc. ASTM, Vol 59, 1959, p 1309

11. J.M. Jacquerie, The Plasticity of Zinc, C.R.M. Metall. Rep., No. 9, 1966, p 51 (in French)

12. A. Hannick and J.M. Jacquerie, The Compression Test with the Cam Plastometer and its Application to

the Determination of Rolling Pressures, C.R.M. Metall. Rep., No. 3, 1965, p 49 (in French)

13. H. Suzuki, S. Hashizume, Y. Yabuki, Y. Ichihara, S. Nakajima, and K. Kenmochi, Studies on the Flow

Stress of Metals and Alloys, Rep. Inst. Ind. Sci., Univ. Tokyo, Vol 18 (No. 3, Serial No. 117), 1968

14. M.J. Stewart, Hot Deformation of C-Mn Steels From 1000 to 2200 F (600 to 1200 C) With Constant

True Strain Rates From 0.5 to 140 s

-1

, The Hot Deformation of Austenite, J.B. Ballance, Ed., The

Metallurgical Society of AIME, 1977, p 47

15. G.T. van Rooyen and W.A. Backofen, A Study of Interface Friction in Plastic Compression, Int. J.

Mech. Sci., Vol 1, 1960, p 1

16. W.F. Adler, T.W. James, and P.E. Kukuchek, “Development of a Dynamic Compression Fixture for

Brittle Materials,” Technical Report CR-81-918, Effects Technology, Inc., Santa Barbara, April 1981

17. D.R. Ireland, “Procedures and Problems Associated with Reliable Control of the Instrumented Impact

Test,” in “Instrumented Impact Testing,” STP 563, ASTM, 1974, p 3–29

18. H.J. Saxton, D.R. Ireland, and W.L. Server, “Analysis and Control of Inertial Effects During

Instrumented Impact Testing,” in “Instrumented Impact Testing,” STP 563, ASTM, 1974, p 50–73

19. D.R. Ireland, “Critical Review of Instrumented Impact Testing,” Paper 5 presented at the Dynamic

Fracture Toughness Conference, London, July 1976

20. B. Hopkinson, A Method of Measuring the Pressure Produced in the Detonation of Explosives or by the

Impact of Bullets, Philos. Trans. A, Vol 213, 1914, p 437

21. R.M. Davies, A Critical Study of the Hopkinson Pressure Bar, Philos. Trans. R. Soc. (London) A, Vol

240, 1948, p 375

22. H. Kolsky, An Investigation of the Mechanical Properties of Materials at Very High Rates of Loading,

Proc. R. Soc. (London) B, Vol 62, 1949, p 676

23. G.I. Taylor, The Use of Flat-Ended Projectiles for Determining Dynamic Yield Strength, Part I:

Theoretical Considerations, Proc. R. Soc. (London) A, Vol 194, 1948, p 289–299

24. A.C. Whiffin, The Use of Flat-Ended Projectiles for Determining Dynamic Yield Strength, Part II: Tests

on Various Metallic Materials, Proc. R. Soc. (London) A, Vol 194, 1948, p 200–232

25. M.L. Wilkins and M.W. Guinan, Impact of Cylinders on a Rigid Boundary, J. Appl. Phys., Vol 44,

1973, p 1200–1206

26. J. Buchar, S. Rolc, V. Hruby, On the Explosive Welding of a Ring to the Axisymmetric Body, J. Mater.

Process. Technol., Vol 85 (No. 1–3), Jan 1999, 171–174

27. R.W. Armstrong, F.J. Zerilli, Deformation Twinning: From Atomic Modeling to Shock Wave Loading,

Advances in Twinning, Minerals, Metals and Materials Society/AIME, 1999, p 67–81

28. H. Couque, On the Use of the Symmetric Taylor Test to Evaluate Dynamic Ductile Compression

Fracture Properties of Metals, Structures under Shock and Impact V, Computational Mechanics

Publications, 1998, p 579–589

29. D.C. Erlich, D.A. Shockey, and L. Seaman, Symmetric Rod Impact Technique for Dynamic Yield

Determination, Second Topical Conference on Shock Waves in Condensed Matter, Menlo Park, CA,

AIP Conf. Proc., No. 78, 1981, p 402–406

30. D.C. Erlich, D.A. Shockey, and L. Seaman, Symmetric Rod Impact Test for Dynamic Yield

Determination, Shock Waves in Condensed Matter, American Institute of Physics, 1983, p 402–406

31. W.H. Gust, High Impact Deformation of Metal Cylinders at Elevated Temperatures, J. Appl. Phys., Vol

53 (No. 5), 1982, p 3566–3575

High Strain Rate Tension and Compression Tests

High Strain Rate Tension Testing

As previously noted, valid stress-strain data can be obtained in conventional load frames for strain rates in a

range up to about 10 to 100 s

-1

if suitable precautions are observed with regard to wave propagation effects and

instrument response times. For tension testing at higher strain rates, or for cases when wave propagation effects

are important, then more specialized testing techniques may have to be used, as described subsequently.

Expanding Ring Test

The expanding ring test is a highly sophisticated technique for subjecting metals to tensile strain rates over 10

-1

(Ref 32, 33, 34). Although the testing principle is simple, its performance requires specialized equipment

available in only a few laboratories. The ring test can determine the high-rate stress-strain relationships, but a

simplified, more widely used version can be employed to determine ultimate strain only (Ref 35, 36).

This test involves the sudden radial acceleration of a ring due to detonation of an explosive charge or

electromagnetic loading. The ring rapidly becomes a free-flying body, expanding radially and decelerating due

to its own internal circumferential stresses. A thin ring must be used for the analysis to be valid; the wall

thickness should be less than one-tenth the ring diameter, which is typically 25 mm (1 in.). If R is the radius of

the ring and σ is the hoop stress:

(Eq 17)

To obtain stress-strain data, radial displacement as a function of time must be calculated. Strain is proportional

to change in radius (just as engineering strain in tension is ΔL/L

); thus:

(Eq 18)

where R

is the initial radius. Stress may be computed from Eq 17 by double differentiation of radial

displacement data as a function of time. Ring displacement can be obtained through the use of high-speed

photography, streak cameras, displacement interferometers, or other methods for measuring radius as a function

of time.

It is difficult to determine stress accurately by double differentiation of displacement data. Several laboratories

have used a laser velocity interferometer to measure ring velocity directly (Ref 37, 38). Thus, only a single

differentiation is necessary to calculate stress, and precision is improved considerably.

Advantages of the Ring Test. The ring test has two principal advantages. The expanding ring test subjects the

material to a state of dynamic uniaxial stress without the wave propagation complications that accompany other

high strain rate tests. Also, the maximum strain rate available in the ring test is higher than in any other

common tension tests involving large plastic strains.

Limitations of the Ring Test. Strain rate in the expanding ring test is not usually constant. The strain rate is

computed from (dR/dt)/R, and both of these terms vary continually. Strain rate is usually greatest at the start of

ring deceleration, when strain is smallest. Values in excess of 10

-1

are readily obtained. If the ring does not

rupture, the strain rate falls to zero at the end of the test.

Ring specimens also experience a compressive preload in the radial direction that often exceeds the yield stress

during the acceleration phase. Because load history is known to affect the subsequent stress-strain behavior of

many materials, data obtained from expanding ring tests do not always agree with results from other tests at

slightly lower strain rates.

The difficulties, expenses, and limitations of the expanding ring test preclude its use as a standard test technique

for generating high strain rate stress-strain data in tension. Only a few laboratories are capable of performing

this test. However, if subjecting a material to high strain rates in tension without determining stress-strain data

is of primary interest, the expanding ring test is much easier to conduct. A number of investigators have used

this test to determine strain to failure under dynamic loading (Ref 35, 36). Here, the accurate determination of

radial displacement versus time is not as critical because stresses are not calculated. Less precise displacement

data provide reasonably accurate determinations of strain rate. The ambiguity arising from possible strain rate

history effects still exists when the expanding ring test is used in this simpler manner.

The expanding cylinder test, a variation of the ring test, provides a dynamic stress state equivalent to that

produced in a quasi-static tensile test on a wide sheet versus a thin strip of material. A difficulty encountered in

this type of test is the need for an impulse to be generated simultaneously in time along the axis of the cylinder.

Because explosive detonation along a wire, for example, propagates at a finite wave speed, uniform

deformation along the length of the axis cannot be ensured. Dimensions, detonation wave speeds, and

synchronization of multiple detonation all must be considered carefully to ensure that the cylinder is deformed

as uniformly as possible and that axial stress waves are not generated (Ref 39).

Flyer Plate and Short Duration Pulse Loading