Fatigue properties of a nanocrystalline titanium based bulk metallic glassy alloy

In contrast, the different crack growth rate was obtained in L-Region II: the higher crack growth resistance was found for Ti-Al6V4, as compared to Ti-BMG.. This was attributed to the re[r]

Trang 1

Original Article

Fatigue properties of a nanocrystalline titanium based bulk metallic

glassy alloy

Graduate School of Natural Science and Technology, Okayama University, 3-1-1 Tsushimanaka, Kita-ku, Okayama, 700-8530, Japan

a r t i c l e i n f o

Article history:

Received 23 August 2018

Received in revised form

5 October 2018

Accepted 14 October 2018

Available online 25 October 2018

Keywords:

Crack closure

Metallic glass

Titanium

Crack growth

Fatigue failure mechanism

a b s t r a c t

To obtain a better understanding of the fatigue properties and crack growth characteristics of a nano-crystalline titanium based bulk metal glasses (Ti-BMG) made by vacuumed casting process, the fatigue failure mechanisms of Ti-BMG have been investigated via Se N and da/dN eDK tests For comparison, the crystalline Ti alloy Ti-Al6V4 was also employed The fatigue strength in the early fatigue stage was high for Ti-BMG due to the high tensile strength However, the fatigue strength decreased signiﬁcantly in the late fatigue stage The higher slope of Se N relation was detected for Ti-BMG, which crossed that for the Ti-Al6V4 sample around 5 103cycles In the higher Region II, the fatigue crack growth rate was of similar level for both Ti-BMG and Ti-Al6V4 due to their similar strain energy In the lower Region II, however, the lower crack growth resistance was obtained for Ti-BMG, as compared to Ti-Al6V4 This was attributed to the high crack driving force for Ti-BMG, caused by the weak roughness-induced crack closure Such crack closing characteristics of Ti-BMG were systematically investigated by various experimental techniques

© 2018 The Authors Publishing services by Elsevier B.V on behalf of Vietnam National University, Hanoi This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/)

1 Introduction

Since metallic glasses have been developed by the process of

rapid cooling in 1960, a large number of scientists has developed

the metal glasses for engineering applications, thanks to their high

corrosion resistance, good electromagnetic properties and high

tensile strength Because the fatigue failure in the engineering

components and structures is more than 90% of the total, the

fa-tigue properties are signiﬁcantly important This is especially true

for metallic glasses, as their fatigue strength is not so high despite

the high tensile strength Fatigue crack growth behavior in the

amorphous structure is considered to be similar to that in

poly-crystalline steel and aluminum alloys [1] Moreover, the crack

growth mechanism is associated with alternating blunting and

re-sharpening of the crack tip The plastic deformation zone ahead of a

crack is a source for heat generation, which leads to a change of the

fracture mechanism and toughness Crack growth occurs quite

readily due to the lack of microstructural barriers, i.e., no grain

boundary, resulting in a low fatigue strength

The investigation of fracture and fatigue in thin ribbons of a nickel-base metallic glass was carried out by Alpas et al.[2] They have found that fatigue crack growth behavior of the high tensile strength and high toughness amorphous alloy is caused by abnormal microstructure and unusual form of plastic deformation Severe deformation in metallic glasses is considered to arise from ﬂow in localized shear bands, in which the veined fracture surface

is obtained[3] Compact tension specimens were made from the bulk plates of Zr-Ti-Cu-Ni-Be base alloy to examine the fatigue crack growth behavior, which revealed fracture toughness of

KIC¼ 55 MPa m1/2[4] Yokoyama et al.[5]have examined fatigue properties of various Zr-based systems, and they have concluded that the W€ohler curve is different from those of ordinary crystalline structural alloys, as the fatigue strength of the Zr-based BMG is very low due to their low slip resistance The low fatigue endurance limit

of partially crystallized BMGs with respect to that of fully amor-phous alloys was also reported[6] Menzel and Dauskardt[7]have examined the fatigue damage for a Zr-based bulk metallic glass, in which shear bands or mixed-mode cracks, propagating at ~ 49to the applied stress axis after a few cycles, make the low fatigue strength To understand accurately the crack growth behavior, an examination of the crack growth characteristics in detail is signif-icantly important

* Corresponding author Fax: þ81 86 251 8025.

E-mail address: mitsuhiro.okayasu@utoronto.ca (M Okayasu).

Peer review under responsibility of Vietnam National University, Hanoi.

Contents lists available atScienceDirect Journal of Science: Advanced Materials and Devices

j o u r n a l h o m e p a g e : w w w e l s e v i e r c o m / l o c a t e / j s a m d

https://doi.org/10.1016/j.jsamd.2018.10.001

Journal of Science: Advanced Materials and Devices 3 (2018) 478e484

Trang 2

Since Christensen reported the fatigue crack closure as a major

problem in the crack propagation study in 1963[8], this

phenom-enon has been the signiﬁcant parameters to understand the fatigue

crack growth characteristics The concept of crack closure has been

widely applied In this case, Chen et al have proposed valuable

crack closure models [9,10] It was considered that, without the

crack closure parameter, the fatigue crack growth rates cannot be

predicted In particular, the fatigue crack growth behavior in the

near-threshold regime is strongly affected by the crack closure

ef-fect Although the crack growth behavior of bulk glassy alloy has

been examined, there is apparently lack of the study for crack

closing characteristics Thus, the aim of this work is to investigate

the effect of the extent of crack closure on the crack growth

char-acteristics for a nanocrystalline bulk metallic alloy

2 Experimental

2.1 Material preparation

In the present study, the titanium-based nanocrystalline bulk

metallic glass was selected (Ti-BMG: Ti41.5Zr2.5Hf5Cu42.5Ni7.5Si1) A

tubular Ti-BMG sample (f2.0 (OD)f1.6 (ID) 200 mm) was

made by the high speed vacuumed casting technique, in which

copper molds were employed to make the high-speed cooling In

the vacuumed casting process, the material was melted in an arc

melter under vacuum at approximately 2.5 103Pa, and injected

rapidly into a copper mold using a vacuumed casting system

designed by Makabe Giken Co., Ltd, where the solidiﬁcation rate of

this vacuumed casting is about 103K/s The Ti-BMG sample selected

has been used for the colliori-type massﬂow meter under the

vi-bration In this examination, a commercial crystalline material

(ti-tanium alloy Ti-Al6V4) was also employed to compare the

mechanical properties the Ti-BMG samples under investigation To

study the mechanical properties of Ti-BMG, the test specimens for

both samples were originally designed in a dumbbell shape, which

machined through electro-discharge machining.Fig 1(a) shows the

photograph of the specimen and the testingﬁxture, andFig 1(b)

indicates the specimen conﬁgurations for the mechanical testing

For the crack growth test, a sharp notch was made to monitor easily

the fatigue crack growth rate.Fig 2(a) displays the transmission

electron microscope (TEM) image of Ti-BMG and X-ray diffraction

(XRD) patterns of Ti-BMG and Ti-Al6V4 For Ti-BMG, a

nano-crystalline structure and a halo pattern for amorphous structure

were obtained in TEM and XRD, respectively On the contrary,

crystalline structure of a sharp peak was observed for Ti-Al6V4 To

further understand the amorphous structure of Ti-BMG, the

elec-tron backscatter diffraction (EBSD) analysis was carried out

Fig 2(b) displays the inverse poleﬁgure maps of BMG and

Al6V4 As seen, crystal formation is completely collapsed for

BMG, while the crystal structure was apparently formed for

Ti-Al6V4 Furthermore, after the annealing at 900C for 15 min, the

recrystallization occurred in Ti-BMG with grain size of about 20mm,

which is 4 times higher than that for Ti-Al6V4

2.2 Mechanical properties

In the present work, tensile and fatigue tests were carried out

using a screw driven type universal testing machine with 50 kN

capacity The tensile test was conducted at 1 mm/min until the

specimen was fractured completely The stress and strain values

were measured by a standard load cell and strain gauge,

respec-tively, which were monitored during the tensile test using a data

acquisition system in conjunction with a computer The fatigue

strength was examined by two different methods: S N and da/dN

-DK tests The relationship between the applied stress amplitude

and number of cycles toﬁnal failure was investigated in the S - N approach The tensilee tensile cyclic loading was applied at a low cycle frequency 1 Hz up to 105cycles

The da/dN - DK relations were examined after a pre-crack 1.0 mm from the notch tip was created The pre-crack was made under small-scale yielding conditions, where the crack tip was not blunted before the da/dN -DK test The crack length was monitored directly during the cyclic loading using a traveling light type mi-croscope with a resolution of 0.01 mm On the basis of the ASTM standard,DK value was calculated from Eq(1)with the parameters

of crack length (a) and applied cyclic stress range (Ds)[11]:

DK¼Dspffiffiffiffiffiffiffipa

1:12 0:231 a

W

þ 10:55 a

W

2

21:72 a

W

3

þ 30:39 a

W

4

(1)

Fig 1 Specimen and testing ﬁxture for the tensile and fatigue tests: (a) photograph and (b) schematic diagram.

M Okayasu, T Shigeoka / Journal of Science: Advanced Materials and Devices 3 (2018) 478e484 479

Trang 3

To estimate the crack driving force, the relationship between the

applied load and crack opening value (strain) was examined at

several stages in Region II of the da/dN vs.DK The loade strain

curves were measured by a standard load cell and a strain gauge

attached on the specimen To understand the failure characteristics,

the fracture surface characteristics, including the surface

rough-ness, were investigated using a scanning electron microscope

(SEM) and a laser scanning microscope (OLYMPUS, LEXT-OLS4100)

3 Results and discussion 3.1 Tensile properties

Fig 3 presents the representative engineering stresse engi-neering strain curves for Ti-BMG and Ti-Al6V4 As seen, the stress

vs strain curve for Ti-BMG is located at a high level compared to that for Ti-Al6V4: the ultimate tensile stress (sUTS) for Ti-BMG is about 1800 MPa, which is about 60% higher compared to that for Ti-Al6V4 It is also clear that the linear stressestrain relation with the high elastic constant was observed for the Ti-BMG sample, result-ing in no clear plastic deformation In contrast, the plastic defor-mation (or work hardening behavior) was detected for Ti-Al6V4, e.g., the fracture strain (εf) is about 16%

3.2 Fatigue properties

Fig 4shows the relationship between stress amplitude and cyclic number to failure for Ti-BMG and Ti-Al6V4, i.e., Se N curve As seen, different trend of fatigue properties was detected in both samples, where the fatigue strength for BMG is higher than that for

Ti-Fig 2 (a) TEM image of Ti-BMG and XRD patterns of Ti-BMG and Ti-Al6V4 and (b) IPF

maps of the Ti-BMG and Ti-Al6V4 specimens.

Fig 3 Tensile stressestrain curves for Ti-BMG and Ti-Al6V4.

Fig 4 Stress amplitude vs number of cycles to failure for Ti-BMG and Ti-Al6V4.

Trang 4

Al6V4 in the early fatigue stage, but the lower fatigue strength for

Ti-BMG in the late fatigue stage Namely, the higher slope of S - N

re-lations is obtained for Ti-BMG, which crosses that for Ti-Al6V4

around 5 103 cycles as enclosed by the dashed circle inFig 4

Due to the different slope of S vs N, the endurance limit at 105cycles

for Ti-BMG is more than 10 times lower than that for Ti-Al6V4,

namely, 7.6 MPa for Ti-BMG and 117 MPa for Ti-Al6V4 It is

re-ported in the previous study in Ref.[12]that there is linear

rela-tionship between the ultimate tensile strength (UTS) and fatigue

limit for crystalline materials, while UTS for BMGs does not directly

affect their fatigue limit Similar result is reported in Ref.[13], where

no clear relationship between the fatigue limit and the yield strength

is detected The reason behind this is not clear at the moment, but it

could be affected by the vacancy of atom in BMG[12]

To understand clearly the fatigue strength of both Ti-BMG and Ti-Al6V4 samples, their Se N curves were quantitatively evaluated

by a power law dependence of stress amplitude (sa) and cyclic number to failure (Nf)[14]:

where sf is the fatigue strength coefﬁcient and b is the fatigue exponent The values of sf and b for Ti-BMG and Ti-Al6V4 are

sf¼ 33.5 GPa and b ¼ 0.51 (for Ti-BMG) andsf¼ 1.49 GPa and

b¼ 0.13 (for Ti-Al6V4) In this case, a high fatigue strength is ex-pected for highsfand low b value From this estimation, highsfand high b values, obtained for Ti-BMG, are related to the high and low fatigue strength in the early and the late fatigue stage, respectively

To verify the crack growth characteristics in detail, the fatigue crack growth behaviors were investigated.Fig 5 shows the rela-tionship between crack growth rate and the stress intensity factor range (da/dN -DK) for both samples In this case, their crack growth rate could be related to that in Region II This is because the lowDK value for both sample are about 200 MPa mm1/2, which is much higher than that for the related BMG: 56.9 MPa mm1/2for Zr-based bulk metallic glass[15] In this case, two distinct regions of fatigue crack growth are identiﬁed: the lower and higher Region II L-Re-gion II is the range of crack growth rate above the threshold stress intensity, where the crack propagation speed is slow rate H-Region

II is a linear relationship between log da/dN and logDK, i.e., Paris region:

da

where C and m are fatigue crack growth parameters; and C and m values for Ti-BMG are 4 1011and 1.4, which are relatively closed to those for Zr41.25Ti13.75, Ni10Cu12.5Be22.5 bulk metallic glass (C¼ 2.4 1010and m¼ 1.7)[7] As seen inFig 5, different crack growth characteristics are observed The resistance to crack growth

in L-Region II appears to be substantially lower for Ti-BMG,

Fig 5 Relationship between crack growth rate and stress intensity factor range in

Region II (da/dN eDK) for Ti-BMG and Ti-Al6V4.

4

Trang 5

compared to that for Ti-Al6V4, although the crack growth rate for

both samples is similarly observed in H-Region II Such crack growth

rate in L-Region II might be associated with their endurance limits

due to the low crack propagation rate, as shown inFig 4, namely the

higher crack growth rate in L-Region II is related to the lower

endurance limits for Ti-BMG On the other hand, similar crack growth rate in H-Region II is attributed to the similar mechanical properties However, it could be questionable, since the mechanical properties of both samples were quite different, e.g., tensile and fatigue strength It

is general consideration that the strain energy is attributed to their

Fig 7 The models of the ideal crack and crack closure: (a) ideal crack opening and closing and (b) roughness-induced crack closure.

Fig 8 Stress intensity factor (K) vs strain showing the stress intensity factor at crack closing (K cl ), the maximum (K max ) and the minimum stress intensity factors (K min ) for (a)

Trang 6

crack growth rate, as the strain ahead of crack tip can absorb the

crack driving force From the stressestrain curves inFig 3, the strain

energy for BMG and Al6V4 can be estimated as 135 MPa (for

Ti-BMG) and 124 MPa (for Ti-Al6V4) Because of similar strain energy of

both Ti-BMG and Ti-Al6V4, the high crack growth rate for Ti-BMG in

L-Region II is inconsistent, whereas the similar crack growth rate of

both samples in H-Region II is applicable

To understand the different fatigue properties, fracture surface

observation was carried out using the laser scanning microscope

after the fatigue crack growth tests.Fig 6 displays the fracture

surfaces of Ti-BMG and Ti-Al6V4 It should be noted that both

samples were fracture after cyclic loading of more than 104cycles It

is clear that the crack growth characteristics are obviously different

A relatively smooth crack growth surface could be seen for Ti-BMG,

while a coarse fracture surface was obtained for Ti-Al6V4 The

mean surface roughness was Ra¼ 968 nm for Ti-Al6V4, which is

about twice rougher than that for Ti-BMG Such a difference in the

roughness of the fracture surface could make a change of the crack

growth resistance, because of different severity of crack closure, i.e.,

roughness-induced crack closure Such crack closing characteristic

can be interpreted as follows It is considered that

roughness-induced crack closure occurs severely in the low crack growth

rate (L-Region II), and the mechanism of the roughness-induced

crack closure can be interpreted usingFig 7 [16] To understand

the crack closing mechanism easily, ideal crack opening and closing

are indicated inFig 7(a) Due toﬂat crack surfaces without plastic

deformation, roughness- and plasticity-induced crack closures do

not occur signiﬁcantly In this case, the crack surfaces are opened

and those are closed completely after removed the applied load In

contrast, because of the rough fracture surfaces in Fig 7(b), the

crack surfaces make a contact each other before removing the

loading This occurrence makes reduction of the crack driving force

leading to the low crack growth rate, i.e., roughness-induced crack

closure for Ti-6Al4V in L-Region II

3.3 Crack closure characteristics

Due to the difference in the crack growth rate in L-Region II for

Ti-BMG and Ti-Al6V4, the extent of crack closure has been investigated

Fig 8displays the relationship between K and strain for both Ti-BMG

and Ti-Al6V4 samples obtained in Region II Note that the strain value

was measured in the specimen behind the ligament to estimate the crack opening displacement value FromFig 8, the K vs strain ex-hibits a concave shape with different slope, in which the high slope of

K vs strain is detected at the lowDK level, which signifying an ac-celeration in the reduction of the measured strain value at the minimum stress intensity factor (Kmin) It is also seen that a concave unloading portion is apparently reﬂected at the low K level (Kcl), which is the stress intensity factor at crack closing action (fracture surface contact)[17] As inFig 8, the Kclvalue depends on theDK level, where the lower theDK, the higher the Kcl With increasing the

DK value, the slope of K vs strain decreases which would be affected

by severe deformation around the crack tip

There are several crack closure models to quantify the actual crack driving force The incorporation of crack closing effects in terms of the effective stress intensity factor range involves the maximum and the minimum stress intensity factor Based upon this, the effective stress intensity factor range (DKeff) can be esti-mated byDKeff¼ Kmaxe Kcl Based on theDKeffvalues, the variation

Fig 9 Variation of the ratio ofDK eff andDK (U) as a function ofDK for Ti-BMG and

Ti-Al6V4.

Fig 10 Relationship between log da/dN and logDK (logDK eff ) in L-Region II for (a)

Trang 7

of U (DKeff/DK) as a function ofDK for both samples is shown in

Fig 9 From this, it is appeared that the value of U is altered

depending on DK, where the U value for the Ti-BMG is overall

higher than that for Ti-Al6V4 This occurrence is reﬂected by the

weak crack closure for Ti-BMG, resulting in the high crack growth

rate at the lowDK region In this case, severe crack closure occurred

for Ti-Al6V4 due to the rough fracture roughness Because of ductile

properties for Ti-Al6V4, crack blunting and constraining of the

shear bands may have also enhanced fatigue crack growth

resis-tance, i.e., plasticity-induced crack closure

Based on theDKeffvalues obtained, the relationship between log

da/dN and logDK (logDKeff) in the L-Region II for both Ti-BMG and

Ti-Al6V4 was indicated inFig 10 As seen, the da/dN vs.DKefffor

Ti-Al6V4 is shifted to the left-hand side due to the severe crack

closure In contrast, the da/dN vs DKefffor Ti-BMG did not shift

signiﬁcantly compared to the Ti-Al6V4 one It is convinced from

this result that the crack growth rate is not delayed for Ti-BMG,

because of weak crack closure arising from the smooth fracture

surface Note that, in this case, no clear microstructural barrier of

Ti-BMG is also signiﬁcant factor It is therefore the BMG samples do

not have high fatigue properties

4 Conclusion

An examination has been made of the fatigue and crack growth

properties for Ti-BMG and Ti-Al6V4 crystalline structures, the

fa-tigue failure characteristics of Ti-BMG have been clariﬁed The

fa-tigue strength for Ti-BMG was high in the early fafa-tigue stage due to

the high tensile strength However, the fatigue strength decreased in

the late fatigue stage The higher slope of Se N relations was

ob-tained for Ti-BMG, which crossed those for the Ti-Al6V4 sample

around 5 103cycles Fracture surface for Ti-BMG after the fatigue

test was dominated by the smooth surface Rough fracture surface

was obtained for the Ti-Al6V4, which was about twice higher than

that for Ti-BMG The fatigue crack growth rate in H-Region II of the

fatigue stage for Ti-BMG was similarly observed for the Ti-Al6V4

sample, which was attributed to the similar strain energy level In

contrast, the different crack growth rate was obtained in L-Region II:

the higher crack growth resistance was found for Ti-Al6V4, as

compared to Ti-BMG This was attributed to the reduction in the

crack driving force arising from the different severity of crack

closure, e.g., roughness-induced crack closure Because of the

smooth fracture surface for Ti-BMG, the crack growth rate enhanced

Acknowledgments The authors would like to express their appreciation to Professor Mitsuru Watanabe for his helpful comments and suggestions on the manuscript

References

[1] C.J Gilbert, V Schroeder, R.O Ritchie, Mechanisms for fracture and fatigue-crack propagation in a bulk metallic glass, Metall Mater Trans A 30A (1999) 1739e1753

[2] A.T Alpas, L Edwards, C.N Reid, Fracture and fatigue crack propagation in a nickel-base metallic glass, Metall Trans A 20 (1989) 1395e1409

[3] K.M Flore, R.H Dauskardt, Local heating associated with crack tip plasticity in Zr-Ti-Ni-Cu-Be bulk amorphous metals, J Math Res 14 (1999) 638e643 [4] C.J Gilbert, R.O Ritchie, W.L Johnson, Fracture toughness and fatigue-crack propagation in a Zr-Ti-Ni-Cu-Be bulk metallic glass, Appl Phys Lett 71 (1997) 476e478

[5] Y Yokoyama, K Fukaura, H Sunada, Fatigue properties and microstructures of Zr55C30Al10Ni5 bulk glassy alloys, Mater Trans JIM 41 (2000) 675e680 [6] G.Y Wang, P.K Liaw, Y Yokoyama, M Freels, R.A Buchanan, A Inoue, C.R Brooks, Effects of partial crystallization on compression and fatigue behavior of Zr-based bulk metallic glasses, J Mater Res 22 (2007) 493e500

[7] B.C Menzel, R.H Dauskardt, Stress-life fatigue behavior of Zr-based bulk metallic glass, Acta Mater 54 (2006) 935e943

[8] R.H Christensen, Fatigue crack growth affected by metal fragments wedged between opening-closing crack surfaces, Appl Mater Res (Oct 1963) 207e210

[9] D.L Chen, B Weiss, R Stickler, A model for crack closure, Eng Fract Mech 53 (1996) 493e509

[10] D.L Chen, B Weiss, R Stickler, Contribution of the cyclic loading portion below the opening load to fatigue crack growth, Mater Sci Eng A 208 (1996) 181e187

[11] T Kunio, H Nakazawa, I Hatashi, H Okamura, Hakai Rikigaku Jikken Hoho (Fracture Mechanics Experimental Method), vol 4, Asakura Publishing Co., Ltd., 1993, pp 240e241

[12] M Okayasu, S Takasu, T Bitoh, D.H Shin, M Makabe, Fatigue and tensile properties of titanium based bulk glassy alloys, Int J Mater Eng Technol 5 (2) (2011) 77e90

[13] G.Y Wang, P.K Liaw, Y Yokoyama, A Inoue, C.T Liu, Fatigue behavior of Zr-based bulk-metallic glasses, Mater Sci Eng A 494 (1e2) (2008) 314e323 [14] E.S Puchi-Cabrera, M.H Staia, D.T Quinto, C Villalobos-Gutierres, E Ochoa-Perez, Fatigue properties of a SAE4340 steel coated with TiCV by PAPVD, Int J Fatig 29 (2007) 471e480

[15] Y Nakai, S Hosomi, Fatigue crack initiation and small-crack propagation in Zr-based bulk metallic glass, Mater Trans 48 (2007) 1770e1773

[16] S Ishihara, Y Sugai, A.J McEvily, On the distinction between plasticity- and roughness-induced fatigue crack closure, Metall Mater Trans 43 (9) (2012) 3086e3096

[17] S Suresh, R.O Ritchie, Some considerations on the modeling of oxideinduced fatigue crack closure using solutions for a rigid wedge inside a linear elastic crack, Scripta Metall 17 (1983) 575e580

Định dạng
Số trang	7
Dung lượng	2,39 MB