Mechanics of Materials 116 (2018) 3–10

Contents lists available at ScienceDirect

Mechanics of Materials

journal homepage: www.elsevier.com/locate/mechmat

High-strain-rate plastic deformation and fracture behaviour of

Ti-5Al-5Mo-5V-1Cr-1Fe titanium alloy at room temperature

Chun Ran

a , Pengwan Chen

a , ∗, Ling Li a , Wangfeng Zhang

b , Yanlong Liu

a , Xiao Zhang

a

a State Key Laboratory of Explosion Science and Technology, Beijing Institute of Technology, Beijing 10 0 081, PR China b Beijing Institute of Aeronautical Materials, Beijing 10 0 095, PR China

a r t i c l e i n f o

Article history:

Received 4 November 2016

Revised 14 June 2017

Available online 24 August 2017

Keywords:

Ti-55511

High strain rate

Dynamic compression

Shear bands

Room temperature

a b s t r a c t

To study the plastic deformation and fracture behaviour of Ti-5Al-5Mo-5V-1Cr-1Fe (Ti-55511) alloy un-

der high strain rate loading conditions, a series of dynamic compression tests on Ti-55511 alloy have

been performed at constant strain rates ranging from 350 s −1 to 2900 s −1 by means of split Hopkin-

son pressure bar technique at room temperature. The different strain and strain rate loading conditions

are realized by changing the length and velocity of the striker bar, and high intensive localized shear

region is induced in Ti-55511 alloy. The dynamic compression stress-strain response, strain rate harden-

ing effect and strain rate sensitivity, and the fracture behaviour are discussed. The experimental results

demonstrate that: The strain rate hardening effect and strain rate sensitivity of Ti-55511 alloy are appar-

ent; Brittle shear bands form at high strain rate loading conditions, and the formation of a shear band

does not mean the occurrence of phase transformation; Collapse of the specimens occurs along a plane

inclined at an angle of about 45 ° to the compression axis at room temperature for both quasi-static and

dynamic compression loading; The shear-compression zone and tension-shear-compression zone coex-

ist in the fracture surface, and collapse of the specimens is attributed to shear failure mechanism for

Ti-55511 alloy under compression loading at room temperature.

© 2017 Elsevier Ltd. All rights reserved.

1

a

o

c

s

a

C

h

b

g

X

o

t

t

a

s

2

f

f

t

a

w

r

b

M

s

a

1

H

t

e

e

p

a

t

b

h

0

. Introduction

Titanium alloys have been extensively utilized in the fields of

erospace and automotive as key structural components because

f their high strength-to-weight ratio, good hardenability and ex-

ellent combinations of corrosion, toughness and crack growth re-

istance ( Boyer, 1996 ). Dynamic mechanical behaviours of titanium

nd titanium alloys have been studied by many prior researchers.

hichili et al. (1998) pointed out that even when the material was

eavily twinned; the amount of plastic deformation contributed

y twinning was much less than that of dislocations by investi-

ating the high strain rate response of α-Ti at room temperature.

u et al. (2006) investigated the dynamic compression behaviour

f Ti6Al4V (subsequently referred to as TC4) alloy and found that

he equiaxed and distortion-free grains within the bands were

he results of dynamic recrystallization (subsequently referred to

s DRX). Based on dynamic compression tests (shear compres-

ion specimen, subsequently referred to as SCS), Rittel et al. (2006,

0 08a,b, 20 09 ) argued that DRX not only precedes adiabatic shear

ailure but it was also likely to be a dominant micromechanical

∗ Corresponding author.

E-mail addresses: rangle123@163.com (C. Ran), pwchen@bit.edu.cn (P. Chen).

f

a

h

ttp://dx.doi.org/10.1016/j.mechmat.2017.08.007

167-6636/© 2017 Elsevier Ltd. All rights reserved.

actor in the generation of the band. Bai et al. (1994) investigated

he microstructure evolution during shear localization of TC4 alloy

nd pointed out that the parabolic dimples on the fracture surface

ere the result of elongated α phase close to the localized shear

egion and no phase transformation was observed within the shear

and zone. A series of corresponding work has been carried out by

eyers et al. (1994), Xue et al. (20 01, 20 02 ), Peirs et al. (2010) and

o forth.

Ti–5Al–5Mo–5V–1Cr–1Fe (subsequently referred to as Ti-55511)

lloy is superior as an aircraft structural material considering its

5–20% weight loss as compared to TC4 alloy ( Liu et al., 2014a ).

ence, the increasing attention has been received with respect

o Ti–55511 alloy by many material scientists since 2013 ( Ahmed

t al., 2013, 2014; Liu et al., 2016a,b, 2014b; Peng et al., 2013; Shi

t al., 2015 ).

However, the focus of prior researches on Ti-55511 alloy was

rimarily on low strain rate loading conditions ( < 100 s −1 ). In fact,

ircraft structures made of Ti-55511 alloy are inevitably subjected

o high strain rate ( > 100 s −1 ) loading conditions, and limited num-

er of publications in this area have been reported. Recent studies

rom Wang et al. (2015) have shown that the grains in the bound-

ry of the shear band were elongated. However, the mechanical be-

aviour and failure mechanism of Ti-55511 alloy under high strain

4 C. Ran et al. / Mechanics of Materials 116 (2018) 3–10

Table 1

Chemical composition of Ti–5Al–5Mo–5V–1Cr–1Fe alloy

(wt%).

Al Mo V Cr Fe C N

5.50 4.82 4.82 1.02 1.02 0.02 0.03

H O Zr Si Ti

0.001 0.1 0.15 0.10 Balance

Fig. 1. Typical original microstructure of Ti-55511 alloy; a) lower magnification and

b) higher magnification of A.

i

ε

ε

σ

w

Y

r

o

b

i

t

s

b

r

t

c

B

p

v

i

o

T

t

s

s

rates are still not well understood. Therefore, a systematic study

of the effect of strain rate on the dynamic plastic deformation and

fracture behaviour of Ti-55511 alloy is required.

The purposes of this work are to obtain insights into: (a) the in-

fluence of strain rate hardening effect and strain rate sensitivity on

dynamic plastic deformation behaviour of Ti-55511 alloy at strain

rates ranging from 350 s −1 to 2900 s −1 at room temperature using

SHPB technique and (b) fracture features at high strain rate.

2. Material and experimental techniques

The Ti-55511 alloy used in the present investigation was in the

form of forged bar with a diameter of 155 mm from AVIC Beijing

Institute of Aeronautical Materials, PR China. The β transus tem-

perature of the as-received bars is approximately 1163 K via met-

allographic observation method, and the chemical composition is

listed in Table 1 ( Ran et al., 2017 ). Al is α-stabilizer element, which

makes the alloy weldable. Mo and V are β isomorphous stabilizer

elements, and Fe and Cr are the β eutectoid stabilizer element. The

β phase stabilizers enhance its hardenability. Others are additions.

As the β-stabilizer content increases, the hardenability increases,

but welding becomes more difficult ( Boyer, 1996 ). The initial mi-

crostructure of this alloy consists of structure of β transforma-

tion (matrix) and α phase (platelet α and equiaxed α), as shown

in Fig. 1 . The cylindrical specimens were all machined from the

forged bar and turned to 6 mm in diameter and 6 mm in height.

Dynamic compression tests were carried out at constant strain

rates ranging from 350 s −1 to 2900 s −1 at room temperature by

means of split Hopkinson pressure bar (subsequently referred to

as SHPB) technique. The setup used in this work was nickel (18Ni)

bars with a diameter of 14 mm and the lengths of the striker, in-

cident and transmission bar were 0.2 m, 1.2 m, and 1.2 m, respec-

tively, and three cylindrical specimens were used to test dynamic

compression properties under each strain rate loading condition.

Based on GB/T 7314-2005 (Metallic Materials-Compression testing

at ambient temperature), an INSTRON 5985 testing machine was

used to conduct quasi-static compression experiments at strain

rate ranging from 10 −4 s −1 to 10 −3 s −1 . It should be pointed out

that the bar-specimen interfaces were sufficiently lubricated (for

both quasi-static and dynamic compression tests) in order to avoid

any barrelling of the specimens.

The samples for microstructure observation were cut along ax-

ial direction by electrical discharge machining and metallographic

specimens were prepared by standard mechanical polishing and

etched in the Kroll’s reagent. The microstructures were examined

by LEICA DMI 30 0 0M optical microscope (OM) and HITACHI S-

4800 scanning electron microscope (SEM).

3. Results and discussion

3.1. Mechanical response of Ti-55511 alloy at high strain rates

The dynamic compression tests were conducted with cylindrical

specimens on SHPB at room temperature. When one-dimensional

stress-waves in the bars are achieved and the specimen is in a

state of uniform stress, the strain rate, strain and stress histories

n the specimen can be determined by ( Shukla et al., 2010 ):

˙ (t) = −2

C 0 L s

ε r (t) (1)

( t ) = 2

C 0 L s

∫ ε r (t) dt (2)

s (t) =

A 0 E 0 ε t (t)

A s =

(d 0 d s

)2

E 0 ε t (t) (3)

here A 0 is the cross-sectional area of the bars; E 0 and C 0 are the

oung’s modulus and elastic bar wave speed in the bar material,

espectively; A s and L s are initial cross-sectional area and length

f the specimen, respectively; d 0 and d s are the diameters of the

ar and specimen, respectively. Here εi ( t ), εr ( t ) and εt ( t ) represent

ncident, reflected and transmitted strain histories in the bars at

he specimen ends, respectively.

Fig. 2 shows the typical true stress and strain rate versus true

train at room temperature and a strain rate of 1100 s −1 , it should

e noted that the macro shear failure does not occur at this strain

ate loading condition. As shown in Fig. 2 , (a) In the stage of AB,

he flow stress increases from 1400 MPa to 1460 MPa with in-

reasing true strain due to strain hardening; (b) In the stage of

D, the strain hardening effect and thermal softening effect com-

ete with each other, and the flow stress reaches the maximum

alue (1476 MPa) at D point. In this stage, strain hardening effect

s greater than that of thermal softening effect; (c) In the stage

f DE, the flow stress reduces slowly with increasing true strain.

his phenomenon can be explained as that the strength loss due

o thermal softening becomes greater than that of increasing in

trength due to strain hardening; (d) In the stage of EF, the flow

tress reduces rapidly from 1472 MPa to 1210 MPa with increasing

C. Ran et al. / Mechanics of Materials 116 (2018) 3–10 5

Fig. 2. Typical stress and strain rate as functions of true strain for Ti-55511 alloy at

room temperature and a strain rate of 1100 s −1 .

Fig. 3. Typical true stress vs. true strain curves of Ti-55511 alloy deformed at dif-

ferent strain rates conditions and room temperature.

s

t

p

1

D

e

i

o

d

T

l

s

s

i

s

i

T

o

w

p

i

t

s

w

r

t

Fig. 4. Variations of strain rate sensitivity as a function of strain rate for specimen

deformed at room temperature under ε = 6%. Note that y is the strain rate sensitiv-

ity and x is the strain rate.

n

5

R

1

0

3

t

(

β

w

p

s

2

c

s

o

fi

i

t

w

t

s

t

s

t

o

h

t

3

p

(

c

T

train. The stress drop is about 262 MPa, and the plastic deforma-

ion becomes unstable in this stage. Various mechanisms are pro-

osed for the stress drop, such as thermal softening ( Wang et al.,

992; Dodd and Bai 2015 ), dislocation ( Chichili et al., 1998 ) and

RX ( Xu et al., 2006 ). The thermal softening may be more influ-

ntial according to Rittel and Wang’s report (2008b) , and further

nvestigation is still need to clarify the mechanism; (e) In the stage

f FG, strain and stress decrease simultaneously because of elastic

eformation recovery.

Fig. 3 presents typical true stress versus true strain curves of

i-55511 alloy obtained for quasi-static (0.001 s −1 ) and dynamic

oading (ranging from 480 s −1 to 2300 s −1 ). Apparently, the true

tress-true strain curves of Ti-55511 alloy for dynamic loading

how clearly oscillations due to the reflection of waves at the spec-

men surfaces and the incident bar. As shown in Fig. 3 , the yield

tress of Ti-55511 alloy is about 1200 MPa at quasi-static load-

ng condition, and this is higher than that of Miao (1110 MPa).

his discrepancy may come from the different microstructures

f Ti-55511 alloy. The microstructure of Ti-55511 alloy in Miao’s

ork (2008) is basketweave, while it is bimodal structure in the

resent study. Shi et al. (2015) also reported that the mechan-

cal behaviour of Ti-55511 alloy was affected by the microstruc-

ure. The quasi-static yield stress and flow stress at a fixed plastic

train (6%) are σ YS ≈ 1200 MPa and σ FS ≈ 1210 MPa, respectively,

hile for dynamic loading, σ YD > 1300 MPa and σ FD > 1400 MPa,

espectively. Hence, compared with quasi-static loading condi-

ion, the yield stress and flow stress increase sensibly under dy-

amic loading, implying that the strain rate hardening effect of Ti-

5511 alloy is apparent. Similar phenomena have been reported by

ittel et al. (2008c) and Miao (2008) . When the strain rate is above

370 s −1 , macro shear failure occurs and the failure strain is about

.16.

.2. Effects of strain rate

The strain rate sensitivity can be approximately estimated as

he slope of the flow stress versus the logarithm of the strain rate

Lee and Lin, 1998 ):

=

σi − σ0

log 10 ˙ ε i − log 10 ˙ ε 0 =

σi − σ0

log 10 ( ̇ ε i / ̇ ε 0 ) (4)

here the quasi-static compressive stresses σ 0 and dynamic com-

ressive stresses σ i are obtained in tests conducted at constant

train rates ˙ ε 0 (0.001 s −1 ) and ˙ ε i (ranging from 350 s −1 to

900 s −1 ), respectively. It should be pointed out that the dynamic

ompression stresses are conducted at the same value of compres-

ive plastic strain (6%).

Fig. 4 shows the variations of strain rate sensitivity as a function

f strain rate for specimen deformed at room temperature. At a

xed plastic strain value (6%), strain rate sensitivity increases with

ncreasing strain rate range, and the increases of strain rate sensi-

ivity is more pronounced at high strain rates.

The yield stress and flow stress at a fixed plastic strain (6%)

ere experimentally determined and are given in Fig. 5 , along with

he data of Miao (2008) , which shows the variations of the yield

tress and flow stress at room temperature with the logarithm of

he mean strain rate. As shown in Fig. 5 , the yield stress and flow

tress increase with increasing strain rate. It is important to note

hat the results obtained in our study are consistent well with that

f Miao (2008) . According to the above analysis, the strain rate

ardening effect of Ti-55511 alloy is apparent from the experimen-

al results.

.3. Calculation of temperature in shear localization region

Based on the adiabatic assumption, a large proportion of the

lastic work is converted into heat to rise the local temperature

Dodd and Bai, 2015 ), and the maximum temperature during the

ompression period can be estimated as:

= �T + T 0 =

β∫

τdγ

ρC + T 0 (5)

6 C. Ran et al. / Mechanics of Materials 116 (2018) 3–10

Fig. 5. Yield stress as a function of strain rate for Ti-55511 alloy at room temper-

ature, b) Influence of strain rate on flow stress at a constant plastic strain of 6%.

Note that x is the strain rate, and y is Yield stress in Fig. 5a) while Flow stress in

Fig. 5b).

Fig. 6. Calculated maximum temperature in the shear band of Ti-55511 alloy as a

function of strain rate at room temperature. Note that y is the maximum tempera-

ture in the shear band and x is the strain rate.

Fig. 7. a) Top view of the specimen after test at 1070 s −1 , showing arc type of shear

band. b) Side view of the specimen after test at 1580 s −1 , showing a planar type of

shear band. c) Optical micrograph of shear band obtained from the specimen after

test at 1370 s −1 and the width of the shear band is about 5 μm.

o

m

r

a

r

i

a

i

m

f

t

f

s

e

n

e

B

3

t

i

a

(

F

a

c

t

r

e

where ρ is the mass density, C is the specific heat, T 0 is the ambi-

ent temperature and β is the fraction of plastic work converted to

heat, which is taken as 0.9 in this work. For Ti-55511 alloy, ρ and

C are 4625 kg/m

3 and 523 J/(kg K), respectively. Here, T 0 = 293 K.

The calculated maximum temperature of Ti-55511 alloy is

shown in Fig. 6 , in which the continuous line represents the fitting

f quadratic polynomial to the various strain rates of interest. The

aximum temperature increases almost linearly at the strain rate

anging from 350 s −1 to 1400 s −1 , While the maximum temper-

ture increases slowly with increasing strain rate when the strain

ate is above 1400 s −1 . The corresponding maximum temperature

s about 425 K ( Fig. 6 ) for Ti-55511 alloy loaded at room temper-

ture and a strain rate of 2300 s −1 based on Eq. (7), so the max-

mum temperature is much lower than that of the phase transfor-

ation temperature 1163 K, which means no α → β phase trans-

ormation occurs in the localized shear region.

According to Rittel et al.’s work (20 08b, 20 09 ), the maximum

emperature (below 573 K) was lower than that of the phase trans-

ormation temperature (1263 K) even when the shear stress and

train were the maximum values in their tests. Therefore, to some

xtent, for Ti-55511 alloy, the formation of a shear band does

ot mean the occurrence of phase transformation. The present

xperimental results are in agreement with those of Xu and

ai (2007) and Yang and Cheng (2002) .

.4. Fracture behaviour

After mechanical testing, optical microscopy and scanning elec-

ron microscope studies were performed on the deformed spec-

mens in an effort to identify fracture characteristics and mech-

nisms. In the top view of the sample deformed at 1070 s −1

Fig. 7 a)), an arc-shaped shear band can be seen. As shown in

ig. 7 b), fracture occurred at an angle of 45 ° to the compression

xis. Fig. 7 c) depicts a flat shear band.

The Ti-55511 alloy failed in dynamic compression at nominal

ompressive strains of approximately 16% or less. The true stress-

rue strain curves show a maximum stress level followed by a

apid loss of load-carrying capacity ( Fig. 3 ). There is no appar-

nt barrelling or frictional constraints for specimens deformed to

C. Ran et al. / Mechanics of Materials 116 (2018) 3–10 7

Fig. 8. Typical scanned electron micrograph of a shear band for dynamic compres-

sion at 1370 s −1 and room temperature. a) a shear band and an associated crack, b)

microstructure of a shear band and c) microsturcture adjacent to a shear band.

s

s

1

a

i

F

s

1

l

t

p

c

s

o

m

Fig. 9. 3D schematic diagram showing the fracture features and orientation of shear

band of Ti-55511alloy subjected to dynamic compression to strain rates greater than

1370 s −1 .

t

t

s

c

t

t

i

a

v

r

c

s

m

(

w

i

(

c

l

o

s

r

t

o

c

h

f

c

t

L

s

F

i

f

c

s

t

P

i

e

trains just less than that required for failure strain (16%). Yet when

pecimens are deformed to strain levels nominally greater than

6%, shearing failure occurs along a plane inclined at an angle of

bout 45 ° to the bar axis ( Fig. 7 b)). Such failures are observed dur-

ng both the quasi-static and dynamic compressive deformations.

ig. 7 c) shows the microstructures of the well-developed localized

hear regions deformed at room temperature and a strain rate of

370 s −1 . As shown in Fig. 7 c), the plastic deformation is highly

ocalized in a narrow region with a width of about 5 μm.

The typical scanned electron micrograph of a shear band and

he microstructures adjacent to the shear band for dynamic com-

ression at 1370 s −1 and room temperature are shown in Fig. 8 a)–

). Fig. 8 a) shows a shear band and an associated crack, and the

harp crack extends along the shear band. Typical microstructure

f a shear band is shown in Fig. 8 b), and Fig. 8 c) shows the typical

icrostructure adjacent to the shear band. Compared with the ini-

ial microstructure of Ti-55511 alloy ( Fig. 1 b)), the α phases close

o the shear band stretch along the shear direction, so the α in the

tructure of β transformation close to the shear band is ( Fig. 8 a)-

)). However, the α phase and α in the structure of β transforma-

ion far from the shear band are similar to the initial microstruc-

ure of Ti-55511 alloy (upper right coner of Fig. 8 c)). Therefore, it

s evident that α phases close to the shear band are subjected to

large deformation and stretched along the shear direction and a

ery distinctive boundary separates the shear band from the sur-

ounding structures.

According to the study of Dodd and Bai (2015) , the shear band

an be classified as brittle shear band (no microvoids) and ductile

hear band (with equiaxed or elliptical voids). It is noted that no

icrovoids and microcracks occur in the shear band in our study

Fig. 8 ), implying that brittle shear bands formed in Ti-55511 alloy

hen it is deformed to large plastic strain at high strain rate. This

s different with the results from Meyers et al. (1994), Bai et al.

1994) , and Lee and Lin. (1998) , in which microvoids and/or micro-

racks occurred in the shear band in CP titanium and TC4 alloy.

To describe the fracture mechanisms operating in Ti-55511 al-

oy, a schematic diagram of the observed fracture features and the

rientation of the adiabatic shear band are shown in Fig. 9 . In the

eparated fracture surface, two regions (tension-shear-compression

egion and shear-compression region) are indicated. It is known

hat in uniaxial compression, localized deformation bands occur

n planes of maximum shear stress oriented at 45 ° to the axis of

ompression. However, when there is a bulge on the specimen, a

oop stress appears at the equatorial plane of the cylindrical sur-

ace which induces a tensile loading state. Therefore, the shear-

ompression zone and tension-shear-compression zone coexist in

he fracture surface. And this result is consistent well with that of

ee and Lin (1998) .

The typical fracture morphology of Ti-55511 alloy under quasi-

tatic compression loading (0.001 s −1 ) is shown in Fig. 10 a).

ig. 10 b) is the higher magnification of A in Fig. 10 a) and c)

s the higher magnification of B in Fig. 10 b). It can be seen

rom Fig. 10 a) that the fracture surface can be divided into two

haracteristic zones, a smooth region and a shear region. The

mooth areas are caused by rubbing between the fragment and

he fracture surface ( Liao and Duffy, 1998 ; Rittel and Wang, 2008).

arabolic dimples are elongated along the shear direction, which

ndicates that large plastic deformation takes place. The smooth ar-

as and the dimple areas coexist in the fracture surface along the

8 C. Ran et al. / Mechanics of Materials 116 (2018) 3–10

Fig. 10. Typical fracture morphologies under quasi-static compression loading

(0.001 s −1 ): a) lower magnification, b) higher magnification of A in Fig. 10a), and

c) higher magnification of B in Fig. 10b).

Fig. 11. Typical fracture morphologies under dynamic compression loading

(1580 s −1 ): a) lower magnification, b) higher magnification of A, and c) higher mag-

nification of B.

s

c

l

t

t

o

T

G

f

i

n

i

T

c

f

shear direction ( Fig. 10 b)). This result is consistent with those of

Goods et al. (1979) and Timothy et al. (1985) .

Fig. 11 a) shows the typical fracture surface of Ti-55511 alloy un-

der dynamic compression loading (1580 s −1 ). Fig. 11 b) is the higher

magnification of A in Fig. 11 a) and c) is the higher magnification

of B in Fig. 11 b). Similar to the quasi-static case, the fracture sur-

face can be divided into smooth region and shear region. As shown

in Fig. 11 b), the size of the parabolic dimples is much smaller

and shallower than that under quasi-static loading . Similar results

were also reported by Liu et al. (2005) and Zhang Et Al. (2011) .

Figs. 12 a)–c) show the typical fracture surface of Ti-55511 alloy

under strain rates 1370 s −1 , 2100 s −1 , and 2900 s −1 at room tem-

perature, respectively. It is interesting to note that some parabolic

dimples exist on the fracture surface even when the strain rate

reaches 2900 s −1 , and the size and depth of the dimples decrease

with increasing strain rate.

According to ASM Handbook (Volume 12, pp. 36–75), dimple

hape is governed by the state of stress. When the state of stress

hanged, the Ti-55511 alloy can exhibit not only smaller and shal-

ower dimples, but also a change in the fracture mode in response

o the restriction on plastic deformation. These changes in frac-

ure mode are most evident in the general region of the fracture

rigin and may not be present over the entire fracture surface.

he result obtained in our study is consistent well with that of

rebe et al. (1985) , and the parabolic dimples on the fracture sur-

ace are the relic of elongated phase close to the shear band, which

s well consistent with that of Bai et al. (1994) .

According to the above analysis, for both quasi-static and dy-

amic compression loading, smooth areas and dimple areas coex-

st in the fracture surface along the shear direction ( Figs. 10–12 ).

herefore, for both quasi-static and dynamic loading conditions,

ollapse of the specimens is attributed to shear failure mechanism

or Ti-55511 alloy at room temperature.

C. Ran et al. / Mechanics of Materials 116 (2018) 3–10 9

Fig. 12. Typical fracture morphologies under dynamic compression loading: a)

1370 s −1 ; b) 2100 s −1 ; and c) 2900 s −1 .

4

l

r

m

t

a

s

a

t

s

b

t

p

r

t

i

c

u

A

S

a

e

N

a

t

A

a

R

A

A

B

B

C

D

G

G

L

L

L

L

L

L

L

M

M

P

P

R

R

R

R

R

R

S

. Conclusions

The plastic deformation and fracture behaviour of Ti-55511 al-

oy subjected to high strain rate at room temperature have been

eported. Results obtained from the mechanical testing show that

acro shear failure occurs when strain rate is above 1370 s −1 , and

he failure strain is about 0.16. Collapse of the specimens occurs

long a plane inclined at an angle of about 45 ° to the compres-

ion axis, and the strain rate hardening effect of Ti-55511 alloy is

pparent. Microstructure observations reveal that α phases close

o the shear band stretch along the shear direction, and brittle

hear bands form with further plastic deformation. Whilst, shear

ands are the precursor to the crack formation and fracture, and

he formation of a shear band does not mean the occurrence of

hase transformation. Analyses of fracture morphology show that

elative smooth areas and ductile dimple areas coexist in the frac-

ure surface for both quasi-static and dynamic compression load-

ng. The elongated shallow dimples on the fracture surface indi-

ate shear failure is the main failure mechanism for Ti-55511 alloy

nder compression loading at room temperature.

cknowledgements

This research was financially supported by the National Natural

cience Foundation of China (Grant Nos. 11472054 and; 11221202 )

nd the Opening Project of State Key Laboratory of Explosion Sci-

nce and Technology ( Beijing Institute of Technology ) with Grant

o. KFJJ16-02M . One of the authors (Chun Ran) acknowledges the

ssistance of Ph.D. Haozhe Liang in conducting the high strain rate

ests. Useful discussion with Senior Engineer Junfang. Li of Chinese

cademy of Inspection and Quarantine for Advanced Materials is

lso acknowledge and appreciated.

eferences

hmed, M. , Savvakin, D.G. , Ivasishin, O.M. , et al. , 2013. The effect of cooling rates

on the microstructure and mechanical properties of thermo-mechanically pro-cessed Ti–Al–Mo–V–Cr–Fe alloys. Mater. Sci. Eng. 576, 167–177 .

hmed, M. , Savvakin, D.G. , Ivasishin, O.M. , et al. , 2014. The effect of ageing on mi-crostructure and mechanical properties of powder Ti-5Al-5Mo-5V-1Cr-1Fe alloy.

Mater. Sci. Eng. 605, 89–97 .

ai, Y.L. , Xue, Q. , Xu, Y.B. , et al. , 1994. Characteristics and microstructure in the evo-lution of shear localization in Ti-6A1-4V alloy. Mech. Mater. 17, 155–164 .

oyer, R.R. , 1996. An overview on the use of titanium in the aerospace industry.Mater. Sci. Eng. 213, 103–114 .

hichili, D.R. , Ramesh, K.T. , Hemker, K.J. , 1998. The high-strain-rate response of al-pha-titanium: experiments, deformation mechanisms and modeling. Acta Mater.

46, 1025–1043 . odd, B. , Bai, Y.L. , 2015. Introduction to Adiabatic Shear Localization, revised ed.

Imperial College Press .

oods, S.H. , Brown, L.M. , 1979. Overview No. 1: The nucleation of cavities by plasticdeformation. Acta Metall. 27, 1–15 .

rebe, H.A. , Pak, H.R. , Meyers, M.A. , 1985. Adiabatic shear localization in titaniumand Ti-6 pct Al-4 pct V alloy. Metall. Trans. A 16, 761–775 .

ee, W.S. , Lin, C.F. , 1998. Plastic deformation and fracture behaviour of Ti–6Al–4Valloy loaded with high-strain-rate under various temperatures. Mater. Sci. Eng.

241, 48–59 .

iao, S.C. , Duffy, J. , 1998. Adiabatic shear bands in a TI-6Al-4V titanium alloy. J.Mech. Phys. Solids 46, 2201–2231 .

iu, C.M. , Lu, Y. , Tian, X.J. , et al. , 2016a. Influence of continuous grain boundaryα on ductility of laser melting deposited titanium alloys. Mater. Sci. Eng. 661,

145–151 . iu, L.F. , Dai, L.H. , Bai, Y.L. , et al. , 2005. Initiation and propagation of shear bands in

Zr-based bulk metallic glass under quasi-static and dynamic shear loadings. J.

Non-crystall. Solids 351, 3259–3270 . iu, L. , Shangguan, Y.J. , Tang, H.B. , et al. , 2014a. Fretting wear behavior of laser-ni-

trided Ti–5Al–5Mo–5V–1Cr–1Fe alloy fabricated by laser melting deposition.Appl. Phys. A 116, 1993–20 0 0 .

iu, L. , Shangguan, Y.J. , Li, H.Z. , et al. , 2016b. Sliding wear behavior of laser-ni-trided and thermal oxidation-treated Ti–5Al–5Mo–5V–1Cr–1Fe alloy fabricated

by laser melting deposition. Appl. Phys. A 122, 1–7 .

iu, S.F. , Li, M.Q. , Luo, J. , et al. , 2014b. Deformation behavior in the isothermal com-pression of Ti–5Al–5Mo–5V–1Cr–1Fe alloy. Mater. Sci. Eng. 589, 15–22 .

eyers, M.A. , Subhash, G. , Kad, B.K. , et al. , 1994. Evolution of microstructure andshear-band formation on alpha-hcp titanium. Mech. Mater. 17, 175–193 .

iao, P. , 2008. Study On Adiabatic Shear Behavior of TC18 Titanium Alloy andLow Alloy Ultra High Strength Steel [D]. Beijing Institute of Technology, Beijing,

pp. 29–30 .

eirs, J. , Verleysen, P. , Degrieck, J. , et al. , 2010. The use of hat-shaped specimens tostudy the high-strain-rate shear behaviour of Ti–6Al–4V. Int. J. Impact Eng. 37,

703–714 . eng, W.W. , Zeng, W.D. , Zhang, Y.W. , et al. , 2013. Oxidation behavior and effect of

oxidation on microstructure and tensile properties of Ti–5Al–5Mo–5V–1Cr–1Fealloy. J. Alloys Compd. 577, 633–642 .

an, C. , Chen, P.W. , Li, L. , et al. , 2017. Dynamic shear deformation and failure of

Ti-5Al-5Mo-5V-1Cr-1Fe titanium alloy. Mater. Sci. Eng. 694, 41–47 . ittel, D. , Wang, Z. , Merzer, M. , 2006. Adiabatic shear failure and dynamic stored

energy of cold work. Phys. Rev. Lett. 96, 075502 . ittel, D. , Landau, P. , Venkert, A. , 2008a. Dynamic recrystallization as a potential

cause for adiabatic shear failure. Phys. Rev. Lett. 101, 165501 . ittel, D. , Wang, Z.G. , 2008b. Thermo-mechanical aspects of adiabatic shear failure

of AM50 and Ti6Al4V alloys. Mech. Mater. 40, 629–635 . ittel, D. , Wang, Z.G. , Dorogoy, A. , 2008c. Geometrical imperfection and adiabatic

shear banding. Int. J. Impact Eng. 35, 1280–1292 .

ittel, D. , 2009. A different viewpoint on adiabatic shear localization. J. Phys. D 42,214009 .

hi, X.H. , Zeng, W.D. , Shi, C.L. , et al. , 2015. The fracture toughness and its predic-tion model for Ti–5Al–5Mo–5V–1Cr–1Fe titanium alloy with basket-weave mi-

crostructure. J. Alloys Compd. 632, 748–755 .

10 C. Ran et al. / Mechanics of Materials 116 (2018) 3–10

X

Shukla, A. , Ravichandran, G. , Rajapakse, Y. , 2010. Dynamic Failure of Materials andStructures. Springer .

Timothy, S.P. , Hutchings, I.M. , 1985. The structure of adiabatic shear bands in a tita-nium alloy. Acta Metall. 33, 667–676 .

Wang, B.F. , Sun, J.Y. , Wang, X.Y. , et al. , 2015. Adiabatic Shear Localization in a NearBeta Ti-5Al-5Mo-5V-1Cr-1Fe Alloy. Mater. Sci. Eng. 639, 526–533 .

Wang, L.L. , Yu, X.T. , Li, Y.C. , 1992. Advances in Impact Dynamics. Press of Universityof Science and Technology of China .

Xu, Y.B. , Bai, Y.L. , Meyers, M.A. , 2006. Deformation, phase transformation and recrys-

tallization in the shear bands induced by high-strain rate loading in titaniumand its alloys. J. Mater. Sci. Technol. 22, 737–746 .

u, Y.B. , Bai, Y.L. , 2007. Shear localization, microstructure evolution and fracture un-der high-strain rate. Adv. Mech. 496–516 .

Xue, Q. , 2001. Spatial Evolution of Adiabatic Shear Localization in Stainless Steel,Titanium, and Ti-6Al-4V Alloy, Materials Science and Engineering. University of

California, San Diego . Xue, Q. , Meyers, M.A. , Nesterenko, V. , 2002. Self-organization of shear bands in ti-

tanium and Ti–6Al–4V alloy. Acta Mater. 50, 575–596 . Yang, Y. , Cheng, X.L. , 2002. Current status and trends in researches on adiabatic

shearing. Chin. J. Nonferrous Met. 12, 401–408 .

Zhang, J. , Tan, C.W. , Ren, Y. , et al. , 2011. Adiabatic shear fracture in Ti-6Al-4V alloy.Trans. Nonferrous Metals Soc. China 21, 2396–2401 .