Intermetallics 14 (2006) 1411e1416www.elsevier.com/locate/intermet

Shear punch tests for a bulk metallic glass

R.K. Guduru a, K.A. Darling a, R.O. Scattergood a,*, C.C. Koch a,K.L. Murty a, M. Bakkal b, A.J. Shih c

a Department of Materials Science & Engineering, NC State University, Raleigh, NC 27695, USAb Department of Mechanical Engineering, Istanbul Technical University, Istanbul, Turkey

c Department of Mechanical Engineering, University of Michigan, Ann Arbor, MI 48109, USA

Received 30 June 2005; accepted 10 January 2006

Available online 20 March 2006

Abstract

A shear punch test technique was used for characterization of the mechanical properties of Zre5Tie17.9Cue14.6Nie10Al bulk metallicglass. The ultimate shear stress values matched very closely with values derived from uniaxial compression tests reported in the literature.This is consistent with a lack of pressure sensitivity in compression reported for this particular metallic glass. Deformation response was strainrate insensitive up to a critical rate, beyond which softening occurred. The latter was attributed to thermal heating effects.� 2006 Elsevier Ltd. All rights reserved.

Keywords: B. Glasses metallic; B. Mechanical properties at ambient temperature; B. Fracture stress; E. Mechanical properties, theory; F. Mechanical testing

1. Introduction

The mechanical properties and failure mechanisms ofmetallic glasses are controlled by inhomogeneous shear banddeformation at low temperatures and homogeneous viscousflow deformation at temperatures above the glass transition[1e8]. Metallic glasses have high strength, relatively lowelastic modulus and are generally strain rate insensitive [9e18]. In tension and unconstrained compression, the initiationand propagation of a through-section shear fracture lead toearly failure with little plastic strain. The plastic deformationof metallic glasses has been attributed to the nucleation ofshear bands that accommodate the applied strain locally.This phenomenon can be observed in the form of load-curveserrations in compression and nanoindentation tests[7,10,13e18]. In contrast to crystalline metals, the plasticflow stress of metallic glasses can be pressure sensitive. Thisleads to changes in the shear plane angle and an associated

* Corresponding author. Department of Materials Science & Engineering,

NC State University, Campus box 7907, Raleigh, NC 27695-7907, USA.

Tel.: þ919 515 7843; fax: þ919 515 7724.

E-mail address: ron_scattergood@ncsu.edu (R.O. Scattergood).

0966-9795/$ - see front matter � 2006 Elsevier Ltd. All rights reserved.

doi:10.1016/j.intermet.2006.01.052

asymmetry in the uniaxial tensile vs. compressive flow stresses[2]. The pressure effect is attributed to free volume increasedue to shear; the critical stress for shear band propagation isreduced/increased when a tensile/compressive normal stresscomponent acts on the shear plane, as occurs in uniaxial tests[19]. Lowhapandu et al. [20] found inconsistency in thisbehavior for a ZreTieCueBe bulk metallic glass becauselarge shear plane angle changes occurred with no asymmetryin the flow stress. Other investigators have reported flow stressasymmetry in conjunction with changes in the shear planeangle. The intrinsic (zero pressure) plastic flow properties canbe determined by tests that generate shear-dominant loadingconditions, such as the shear punch test. This motivates thepresent study on the bulk metallic glass Zre5Tie17.9Cue14.6Nie10Al (denoted as BAA-11 [21]).

The shear punch test (SPT) technique has been used toobtain estimates of the tensile strength properties for polycrys-talline metals [22]. We have recently applied it to strain ratesensitivity effects [23]. The SPT technique produces the equiv-alent of a stressestrain curve in shear loading. It is similar tothe sheet metal blanking process in which shear deformationoccurs combined with some degree of compression, tensionand bending.

1412 R.K. Guduru et al. / Intermetallics 14 (2006) 1411e1416

2. Experimental procedure

A 6.4-mm diameter BAA-11 rod was prepared at the OakRidge National Laboratory by a rapid casting technique [21].An X-ray pattern for the as-received material is shown inFig. 1. The broad peak at 38 � confirms that an amorphousstructure has been retained. The tests were done using thesetup shown in Fig. 2a. A schematic of the test geometryand the idealized pure shear deformation zone is shown inFig. 2b. A detailed description of the SPT procedure usedhere is given in Ref. [22]. For the current tests, dpunch¼2.49 mm, ddie¼ 2.52 mm and the clearance c¼ (ddie�dpunch)/2¼ 15 mm. This choice for the test parameters willbe in an optimum range for maximizing the shear loadingfor a high strength material. Disk samples for testing werecut from the rod. The disks were ground flat and parallel using400, 600, 1200, 2400 and 4000 grit polishing paper. Final pol-ishing to obtain mirror surfaces was done using 0.1 mm alu-mina powder slurry. The polishing procedure adopted gavevery consistent test results with low scatter (<5%). However,the SPT results can be influenced by poor surface finish. Pre-liminary tests showed that, for example, a final surface finishwith 1200 grit paper produces results with larger amounts ofscatter (>15%). Sample thicknesses varied between 320 and395 mm. Load P vs. punch displacement d curves were ob-tained at punch displacement speeds ranging from 0.424 mm/sto 84.7 mm/s. Shear failure surfaces were observed using aHitachi S3200 scanning electron microscope.

The average shear stress t in an SPT is calculated using therelation

t¼ P

pdmeanhð1Þ

where dmean is the mean of the punch and die diameters and his the specimen thickness. This definition for the shear stresshas been shown to correlate well with uniaxial test resultsfor polycrystalline metals and alloys [22,24e26]. For pureshear, the strain is

150

100

50

0

Inte

nsity

(ar

bitr

ary

units

)

80604020

Angle (degrees)

Fig. 1. X-ray intensity vs. diffraction angle. Cu-Ka radiation.

g¼ d

cð2Þ

This relation is generally not used because the strain for anSPT is spatially distributed outward from the idealized zoneshown in Fig. 2b [24,26]. It is assumed that the average shearstrain and strain rate are proportional to the punch displace-ment and punch speed, respectively. The usual practice in re-porting SPT stress vs. strain curves is to use t vs. d. We followthat practice here, but use the normalized displacement d/hsince we have shown that this eliminates sample thicknesseffects [22].

3. Results and discussion

The t vs. d/h curves for BAA-11 are shown in Fig. 3. Thepunch speeds for test conditions #1e6 were 0.424, 2.12, 4.24,17, 33.9 and 84.7 mm/s, respectively. For clarity, the curves atdifferent speeds are offset from the origin. The slight upwardcurvature during initial loading is due to punch seating onto

Fig. 2. (a) Shear punch test setup. The die portion indicated contains

a clamping fixture for the sample with the die fixed underneath. (b) Schematic

for the shear punch test geometry.

1413R.K. Guduru et al. / Intermetallics 14 (2006) 1411e1416

the sample. The t vs. d/h curves show linear elastic loadingbehavior and an ultimate (maximum) stress point. Shear fail-ure occurs at this point and the remainder of the curve isdue to disk separation and punch-through. The deformationresponse for the lowest testing speed was not consistent withthe higher speed tests. Shear localization failures occurred dur-ing loading, producing reduced strength values and extremescatter in the results. A large shear serration can be seen onthe initial loading portion of the test #1 curve in Fig. 3. Thiseffect did not occur at the higher testing speeds. The ultimateshear stress tu is defined as the maximum stress reached ona t vs. d/h curve. Fig. 4 plots the tu values as a function ofthe punch speed. A data point for the lowest speed was omittedfrom the plot for the reasons mentioned. The tu values inFig. 4 represent averages of three or more tests done at eachspeed, and the �error bars are the maximum deviations forthe multiple tests. The key features of the test results are dis-cussed in the following sections.

1000

800

600

400

200

0

Str

ess

(MP

a)

1.51.00.50.0

Normalized displacement

1 2 3 4 5 6

Fig. 3. t vs. d/h curves for test conditions #1e#6.

1000

950

900

850

800

750

Ulti

mat

e S

hear

Str

ess

MP

a

12 3 4 5 6 2 3 4 5 6

10 1002

Speed µm/s

Fig. 4. tu vs. u for test conditions #2e#6.

3.1. Comparison to uniaxial compression tests

Compression tests on bulk metallic glasses can show sev-eral percent of plastic strain prior to failure. Tensile tests oftenreach failure in the elastic range. The ultimate stress can ex-ceed the elastic limit (deviation from linearity) in uniaxialtests, and this has been attributed by some authors to a risein the plastic flow resistance similar to strain hardening[2,27,28]. The SPT curves in Fig. 3 show an apparent elasticlimit. We have done FEA simulations for the test setup usedhere and have shown that the elastic limit for the t vs. d/hcurves does not correspond to a physically meaningful yieldpoint. Due to the gradual development of global plasticity inthe deformation zone in Fig. 2b, along with punchedie-samplecompliance effects that can be non-linear, a 1% plastic offsetin d/h is needed to reach the shear yield point for typical poly-crystalline metals [24]. Because of the much more extensiveelasticeplastic transition for metallic glasses (elastic strainsat yield can be 10� greater than in polycrystalline metals)and the expected lack of strain hardening capacity, the yieldpoint offset must be significantly larger and this has notbeen studied. In this paper we restrict attention to the deforma-tion response based on the ultimate shear stress at failure, tu,and compare this with test results reported in the literature.

The values for stress in Fig. 4 show a slight increase up toa speed of 17 mm/s, and then drop sharply at higher speeds. Wediscuss the sharp drop further in the next section. The averageof the stress values for the three lower speed tests (2.12e17 mm/s) is tu¼ 936� 8 MPa. Liu et al. [21] obtained an ulti-mate failure stress sc¼ 1865� 15 MPa for BAA-11 in uniaxialcompression tests. The latter authors also found that the shearplane angle in compression was q¼ 45�, indicating no effect ofa compressive normal stress on the shear plane (pressure effect)of the results; a pressure effect does occur for tensile tests. Re-cently completed ball indentation tests on samples cut from thesame BAA-11 rod used for the SPT samples confirm the lack ofa pressure effect in compressive loading [29]. The pressure inde-pendent ultimate shear stress derived from the compression testson BAA-11 [21] would then be tu¼ sc sin2 q¼ 932� 8 MPa,which is in excellent agreement with the SPT value,936� 8 MPa. This gives confidence that the SPT method canbe used to measure the intrinsic (pressure independent) shearstress for bulk metallic glasses, most of which show a pressureeffect for both tensile and compressive loading [2,29].

3.2. Loading rate effects

There are two effects of loading rate on the tu vs. d/hcurves in Fig. 3. Discrete shear band serrations can be seenon the curves at lower punch speeds; these are shown at higherresolution in Fig. 5. The serrations become less distinct asspeed increases and there is a transition to smooth curves in-dicative of homogeneous deformation. A similar diminishingof serrations with increasing strain rate has been observed incompression tests and nanoindentation tests [7,16]. Nanoin-dentation tests on BAA-11 showed a transition to smoothflow for indentation strain rates in the range 0.2e0.5/s [16].

1414 R.K. Guduru et al. / Intermetallics 14 (2006) 1411e1416

The SPT transition occurs between tests #3 and #4 in Fig. 3,which corresponds to shear strain rates in the range 0.3e1.1/s using Eq. (2). The transition to smooth flow at higher loadingrates has been attributed to a critical strain rate, above whichembryonic nucleation of shear bands leading to shear localiza-tion is kinetically restrained [30].

The most significant rate effect is seen in Fig. 4 as a sharpdrop in tu above a punch speed of 17 mm/s. We attribute thisdrop to thermal softening resulting from strain rate inducedheating. Bakkal et al. [31] observed a similar drop in the cuttingforce during machining of BAA-11, which was accompaniedby high temperatures and viscous flow features in the machin-ing chips. Phenomenological equations will be used here togain additional insight into the interplay of rate effects. Assumethat deformation response at lower speeds is due to shear bands(SB), while deformation at higher speeds is due to viscous flow(VF). Appropriate rate equations for each regime are

tSB ¼ C1um ð3Þ

tVF ¼ C2 expðD=TÞu ð4Þ

u¼ dd/dt is the punch speed, T is the temperature in K and Cdenotes a proportionality constant. Eq. (3) represents the tem-perature independent SB deformation with strain rate sensitiv-ity m, and Eq. (4) represents Newtonian VF with a constant Dparameter for viscosity. Following thermal analysis for ma-chining [32], the temperature T due to strain rate induced heat-ing is assumed be proportional to un where n< 1. Eq. (4) willthen have the form

tVF ¼ C2 expðC3=unÞu ð5ÞThe onset of thermal softening with increasing punch speed

becomes apparent at 17 mm/s in Fig. 4, and this is identified asthe common transition point separating low-speed tests from

1000

950

900

850

800

Str

ess

(MP

a)

0.350.300.250.200.150.10

Normalized Displacement

Fig. 5. tu vs. d/h for test #3 using expanded scales.

high-speed tests. Using regression analysis, three data pointsfor low-speed tests (2.12e17 mm/s) were fit to Eq. (3) and threedata points for high-speed tests (17e84.7 mm/s) were fit toEq. (5) to determine the parameters. The resulting equationsare (t and u in MPa and m/s, respectively)

tSB ¼ 925u0:0073 ð6aÞ

tVF ¼ 0:0843 exp�11:43=u0:2

�u ð6bÞ

These equations are plotted in Fig. 6 along with the data fromFig. 4. Eq. (6b) passes through a minimum at u z 65 mm/s be-cause of the opposing effects of viscosity decrease and linearstrain rate increase. It was not possible to do tests at higher punchspeeds to investigate a stress increase. The heating rate exponentn¼ 0.2. If the temperature at the low end of the testing speedrange is assumed to be ambient (300 K), then the temperaturerise at the high end would be 300(84.7/0.424)0.2¼ 865 K. Thisexceeds the glass transition temperature for BAA-11 (666 K)by a large margin, which would be consistent with the onset ofthermal softening effects as punch speed increases. The temper-ature estimated does not consider adiabatic heating in localizedshear bands. Temperatures in the bands could be higher, and theestimated value using the fitted exponent n¼ 0.2 is a lowerbound. Metallic glasses are not expected to have high rate sen-sitivity for temperatures below the glass transition. The best-fit value for m in Eq. (3) was 0.0073� 0.005, which indicatesnegligible strain rate sensitivity. Indentation tests on BAA-11showed no detectable strain rate sensitivity [29].

The phenomenological model proposed here to analyze thedecrease in shear stress at higher speeds is based on thermalsoftening. Alternate ‘‘softening’’ mechanisms could involvebrittle fracture due to high-strain rate stress concentration,and premature shear band failures. Fracture toughness valuesfor BAA-11 are reported to be in the range 50e60 MPa m1/2

Experiment Shear bands (eq. 6a) Viscous flow (eq. 6b)

1000

950

900

850

800

750

Ulti

mat

e S

hear

Str

ess

MP

a

12 3 4 5 6 7 72 3 4 5 6

10 1002

Speed µm/s

Fig. 6. Comparison of Eqs. (6a) and (6b) with experimental data from Fig. 4.

1415R.K. Guduru et al. / Intermetallics 14 (2006) 1411e1416

[21], which is not consistent with brittle behavior. PrematureSPT failures, accompanied by large serrations on the loadingcurve and large scatter in the test results, occurred only at thelowest test speed. As noted previously, these results were ex-cluded from the analysis. The SPT curves showed increasinglyuniform deformation response as the strain rate increased. Sim-ilar effects with an increase in ductility have been observed forother Zr-based metallic glasses [27]. Lacking temperature mea-surements, the thermal softening and strain rate effects pro-posed cannot be directly verified. A temperature rise abovethe glass transition for high-strain rate conditions and thelack of strain rate sensitivity for lower deformation rates,extracted from the parameters fit to the phenomenologicalequations, are consistent with other work on BAA-11 [29,31].

3.3. Shear-failure surfaces

Scanning electron microscopy (SEM) was used to character-ize features on failure surfaces of punched-out SPT disks.Fig. 7a shows a disk viewed along its edge. Fig. 7b shows shear

Fig. 7. SEM micrographs of a punched-out SPT disk for a punch speed of

2.12 mm/s. (a) Edge view with top T and bottom B surfaces indicated; the

top surface is against the punch face. (b) Top-down view of shear bands

formed in the region indicated by SB in Fig. 7a.

band offsets formed on the bottom surface of the disk, adjacent tothe edge (SB arrow). Shear-failure surfaces for metallic glassestested in tension or compression show unique features in theform of vein patterns [2]. Vein patterns were observed alongthe edge regions adjacent to the top (T) and the bottom (B)disk surfaces. These are the locations expected for shear-failureinitiation [24]. Typical features are shown in Fig. 8a, where thevein patterns display a branching-tree topology. No consistenttrends could be identified in these patterns as the testing speedincreased. The interior fracture surface of the disk, between thetop and bottom vein patterns, showed directional shear featuresrather similar to polycrystalline metals [22]. Fig. 8b showsa unique, micron-scale dimpled fracture surface that was ob-served in torn-off burrs formed along the top edge of a disk.

4. Summary and conclusions

A shear punch test technique (SPT) was used to character-ize the mechanical properties of Zre5Tie17.9Cue14.6Nie10Al

Fig. 8. SEM micrographs from the disk edge shown in Fig. 7. (a) Vein pattern

from region T in Fig. 7a. (b) Failure surface at a torn-off burr (arrow) in

Fig. 8a.

1416 R.K. Guduru et al. / Intermetallics 14 (2006) 1411e1416

bulk metallic glass. The ultimate shear stress values obtainedat lower testing speeds agreed very well with the shear failurestress derived from uniaxial compression tests. This is consis-tent with the lack of compressive flow stress pressure sensitiv-ity reported for this particular metallic glass. The agreementgives confidence that the SPT technique can be used to mea-sure an intrinsic (pressure independent) ultimate shear stressfor other bulk metallic glasses, most of which show pressuresensitivity for both compression and tension testing. At thehigher testing speeds, the ultimate shear stress droppedsharply. This was attributed to strain rate induced heatingand onset of viscous flow. Phenomenological-based rate equa-tions were fit to the data and showed that temperature in-creases above the glass transition can be expected. At lowertesting speeds, the deformation response showed negligiblestrain rate sensitivity.

Based on this study we conclude that the shear punch testtechnique is a useful complement to uniaxal tests for charac-terizing the shear deformation response of bulk metallicglasses. It has the advantage of being adaptable to small sam-ple sizes with relatively simple testing geometry.

Acknowledgements

This research was supported by the National Science Foun-dation grant number DMR-0201474. The authors are indebtedto Dr. C.T. Liu (ORNL) for providing the BAA-11 sampleused for this investigation.

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