Wear characteristics of titanium alloy Ti54 for

Tribology International 43 (2010) 2345–2354

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Wear characteristics of titanium alloy Ti54 for cryogenic sliding applications N.S.M. El-Tayeb n, T.C. Yap, P.V. Brevern Faculty of Engineering and Technology, Multimedia University, Jalan Ayer Keroh Lama, Bukit Beruang, 75450 Melaka, Malaysia

a r t i c l e in fo

abstract

Article history: Received 18 August 2009 Received in revised form 30 August 2010 Accepted 31 August 2010 Available online 17 September 2010

Cryogenic wear behaviour of Ti–5Al–4V–0.6Mo–0.4Fe (Ti54) alloy sliding against tungsten carbide is investigated at different speeds, loads and distances. Empirical models based RSM are developed to predict wear characteristics of Ti54 alloy as a function of sliding conditions. It is found that experimental and predicted results are in good agreement. Besides, cryogenic wear is substantially lower than dry wear. SEM and EDS analyses of worn surfaces and wear debris reveal that cryogenic sliding is significantly influenced by changing material properties along with boundary lubrication performance. The study has shown that modes in dry sliding are adhesion and delamination whereas in cryogenic sliding they are abrasion and delamination. & 2010 Elsevier Ltd. All rights reserved.

Keywords: Titanium alloy Cryo-Tribology Wear modeling Electron scanning microscopy

1. Introduction The use of titanium alloys in sliding applications is limited because of their poor wear resistance [1] and galling properties [2]. This is due to low resistance of titanium alloys to plastic shearing as well as low protection by surface oxide formed as a consequence of high flash temperatures (induced by frictional heating) during dry sliding [3]. One of the main reasons behind poor tribological properties of titanium alloys is the low thermal conductivity of these alloys. During sliding, the generated heat dissipates slowly and rises up the interface temperature, which in turn deteriorates the tribological performance of sliding titanium. Thus temperature of the sliding surface needs to be controlled. In high speed machining of titanium and inconel alloys, conventional cutting fluids reportedly failed to reduce the cutting temperature [4] and cryogenic fluid [5] was recommended to replace the conventional cutting fluid. Earlier researches on cryogenic machining showed the benefit of using liquid nitrogen in machining. Turning to cryogenic sliding of titanium alloys, apparently very little work is available in the literature apart from the work conducted by Basu et al. [6] concerning friction and wear of high purity titanium and our previous published work [7] on cryogenic frictional behaviour of titanium alloys Ti–6Al–4V (Ti64) and Ti–5Al–4V–0.6Mo–0.4Fe (Ti54) sliding against tungsten carbide wheel. Basu et al. [6] found that friction and wear of high purity titanium decreased when sliding against steel in liquid nitrogen. Other researchers from German Federal Institute

n

Corresponding author. Tel.: +202 26890000x1468; fax: + 202 26875889/97. E-mail addresses: [email protected], [email protected] (N.S.M. El-Tayeb). 0301-679X/$ - see front matter & 2010 Elsevier Ltd. All rights reserved. doi:10.1016/j.triboint.2010.08.012

for Materials Research and Testing (BAM) studied the tribological behaviour of various material pairs [8]. They used specially designed cryo-tribometers to investigate wear and friction of several pairs of material at room and low temperatures. They reported that tribological behaviours of steel/steel and TiN coating/steel are unchanged at 77 K compared to the tribological behaviour of same the materials at room temperature. However, wear and friction behaviour of polymers against steel changed at low temperature; i.e. wear decreases from room temperature to 77 K. Similar results were reported by Theiler et al. [9], who tested PTFE against steel at room temperature, 77 K (in liquid ¨ nitrogen) and 4.2 K (in liquid helium). Besides this, Hubner et al. [10] and Pinto et al. [11] reported that at extreme low temperature (4–20 K), austenite in stainless steel and steel were transformed into martensite and this caused the stainless steel and steel to become harder, which resulted in higher wear resistance. The above review indicates that there is a strong potential in controlling the tribological behaviour of sliding titanium alloys by introducing cryogenic fluid at the sliding interface. This is one of the main objectives of the current work. To the best of the authors’ knowledge, none of the past researches gave attention to investigate the wear behaviour of Ti alloy (Ti54) sliding against tungsten carbide under cryogenic conditions. Besides no information is available about the combined effect of test parameters or their interactions (i.e. the combined effects of load/speed, load/sliding time and speed/sliding time) on wear behaviour of Ti alloys. All published work focused only on the study of the effect of one parameter at a time. Thus, the current work focuses on two main objectives. The first is to evaluate the wear behaviour of newly developed titanium alloy Ti–5Al–4V–0.6Mo–0.4Fe (Ti54) (supplied by Timet Inc., USA) sliding against tungsten carbide

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N.S.M. El-Tayeb et al. / Tribology International 43 (2010) 2345–2354

wheel. The second is to construct wear maps of Ti54 titanium alloy under dry and cryogenic conditions. The sliding parameters considered are load, speed and test duration in dry and cryogenic environments. In order to reduce the experimental activities and maximize results quality, a design of experiment (DoE) is implemented using a factorial design of 23 along with response

surface methodology (RSM) to develop wear maps for titanium alloys at both room and cryogenic temperatures.

2. Experimental work 2.1. Material and testing procedure

Table 1 Chemical composition and properties of titanium alloy [7]. Element

Weight percentage

Al V Mo Fe C O N Ti Phase Density (g/cm3) Ultimate tensile strength (MPa) Thermal conductivity (W/m K) Specific heat (J/g K) Young’s modulus (GPa) Shear modulus (GPa)

4.5–5.5 3.0–5.0 0.4–1.0 0.2–0.8 0.1 max 0.2 max 0.03 max Balance a+b 4.44 NA 6.96 at 25 1C 0.54 at 25 1C NA NA

* Ti54: Ti–5Al–4V–0.6Mo–0.4Fe.

A newly developed titanium alloy Ti–5Al–4V–0.6Mo–0.4Fe (Ti54) was studied in this research. More details about this alloy can be found elsewhere [7]. The compositions and properties of the alloy are listed in Table 1. Detailed descriptions of a pin-on-ring tester (shown in Fig. 1) and tribo tests and under ambient and cryogenic conditions are given earlier in [7]. Briefly, specimens (pins) of titanium alloy Ti54 of size 5 mm 5 mm 20 mm were tested against a rotating tungsten carbide disc of 25 mm diameter and 9 mm thickness. The square face of the titanium alloy pins were rubbed against the tungsten carbide counterface (detailed specifications of the carbide disc are given elsewhere [12]). Initial surface roughnesses of tested surfaces of titanium Ra were 0.083–0.148 mm. After tests, the roughnesses Ra of the tested surfaces were measured and found in the range 0.466–0.669. Also, hardness of Ti54 was measured and found to be 325 HV10 (equivalent to 68 HRA). The hardness of the carbide disc is 92 HRA, with an initial surface roughness of 1.16 mm Ra. After the test, surface roughness of the disc was reduced to around 0.459–0.536 mm. In this study three independent variables with 5 levels were selected, i.e. normal loads (6.4–22.96 N), sliding speed (0.1295–0.971 m/s) and sliding time (2.64–9.36 min). Before and after each test, pins and discs were weighted by digital scale (Setra EL-410S) with 0.001 g resolution. The difference in wear volume was calculated from the weight loss. Two sets of experiments were performed, the first was conducted at room temperature under dry condition while the second was in liquid nitrogen (LN2). The LN2 jet was directed to the interface between the titanium alloy pin and counterface disc using a nozzle as

Fig. 1. Schematic diagram of pin-on-disc apparatus. Table 3 Analysis of variance and estimated regression coefficients for wear model Eq. (7) (Ti54D). Table 2 Design matrix and experimental results. Input Speed

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17

1 1 1 1 1 1 1 1 1.682 1.682 0 0 0 0 0 0 0

Source Response

Load

1 1 1 1 1 1 1 1 0 0 1.682 1.682 0 0 0 0 0

Time

1 1 1 1 1 1 1 1 0 0 0 0 1.682 1.682 0 0 0

Measured wear volume (mm3) Ti54D

Ti54C

3.3784 2.2523 1.6517 0.8258 2.0270 1.3514 0.9760 0.7508 2.4775 0.7508 2.4775 1.1261 2.4775 0.7508 1.5015 1.8018 1.7267

2.7027 2.0270 1.2763 0.6006 1.1261 1.5015 0.4505 0.4505 1.9520 0.8258 2.1772 0.6757 2.1772 0.6757 1.2012 1.5015 1.3514

Regression Linear Square Interaction Residual Error Lack-of-fit Pure error Total

DF 9 3 3 3 7 5 2 16

Seq SS

Adj SS

9.36795 8.82997 0.04763 0.49035 0.14873 0.09988 0.04885 9.51668

9.36795 0.07751 0.04763 0.49035 0.14873 0.09988 0.04885

Adj MS

F

P

1.04088 0.02584 0.01588 0.16345 0.02125 0.01998 0.02442

48.99 1.22 0.75 7.69

0.000 0.372 0.557 0.013

0.82

0.630

Estimated regression coefficients Term Coef. SE Coef.

T

P

Constant Speed Load Time Speed Speed Load Load Time Time Speed Load Speed Time Load Time

0.860 1.375 1.312 0.111 0.637 0.891 0.637 2.549 3.642 1.821

0.418 0.211 0.231 0.915 0.544 0.402 0.544 0.038 0.008 0.111

0.82299 1.61161 0.08713 0.01803 0.44280 0.00161 0.00692 0.10714 0.37537 0.00957

0.95695 1.17201 0.06643 0.16291 0.69461 0.00180 0.01085 0.04203 0.10307 0.00525

S¼ 0.145764, PRESS¼ 0.873185, R2 ¼ 98.44%, R2(pred) ¼90.82%, R2(adj)¼ 96.43%.


shown in Fig. 1. The wear results are presented as a function of load, speed and test duration. In order to elucidate the wear processes and thus the fundamental mechanisms, worn surfaces of the alloy pins and wear debris were examined by scanning electron microscopy (ZEISS SUPRA 35VP) equipped with energy-dispersive spectroscopy (EDS). Chemical analyses of the elements in the worn surfaces and wear debris were also done using EDS. 2.2. Design of experiment Five levels of each factor are involved in this design. These levels are for speed (0.1295, 0.3, 0.55, 0.8 and 0.9705 m/s), for load (6.464, 9.81, 14.715, 19.62 and 22.956 N) and for sliding time (2.636, 4, 6, 8 and 9.364 min). The transforming equations used to transform the natural variables into coded form ( 1.682, 1, 0, 1 and 1.682.) are given for each of the independent variables as x1 ¼

speed0:55 ; 0:25

x2 ¼

load14:715 ; 4:095

x3 ¼

time6 2

ð1Þ

A central composite design (CCD) with nc ¼3 center runs and

a ¼1.682 was used to describe the response surface of wear volume. The design consists of a factorial portion, an axial portion and a center point. The factorial portion is a complete (2k) factorial design with factor levels coded by 1 and +1. The axial portion or star points are points located on the coordinate axes of the factorial portion at a distance a from the design center. The center point is given for a number (n0) of repetition runs conducted for the condition in (0 0 0). Table 2 shows the design matrix and experimental results for alloys in both dry (D) and cryogenics (C) conditions.

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Table 4 Analysis of variance and estimated regression coefficients for wear model Eq. (8) (Ti54C). Source

DF

Seq SS

Regression Linear Square Interaction Residual Error Lack-of-fit Pure error Total

9 3 3 3 7 5 2 16

6.95734 6.40389 0.00462 0.54883 0.54485 0.49976 0.04509 7.50219

Adj SS 6.957342 0.182321 0.004624 0.548827 0.544849 0.499759 0.045090

Adj MS

F

P

0.773038 0.060774 0.001541 0.182942 0.077836 0.099952 0.022545

9.93 0.78 0.02 2.35

0.003 0.541 0.996 0.159

4.43

0.194

Estimated regression coefficients Term Coef. SE Coef.

T

P

Constant Speed Load Time Speed Speed Load Load Time Time Speed Load Speed Time Load Time

0.033 1.100 0.058 0.630 0.243 0.083 0.083 2.188 1.427 0.476

0.974 0.308 0.955 0.549 0.815 0.936 0.936 0.065 0.197 0.649

0.06110 2.46730 0.00736 0.19647 0.32336 0.00029 0.00173 0.17602 0.28153 0.00478

1.83159 2.24320 0.12714 0.31181 1.32948 0.00345 0.02077 0.08044 0.19728 0.01005

S¼ 0.278990, PRESS ¼ 3.92409, R2 ¼92.74%, R2(pred) ¼ 47.69%, R2(adj)¼ 83.40%.

3. Development of wear models for titanium alloy Ti54 The traditional way to evaluate material tribology behaviour is to perform experiments by changing one variable while the other factors are constant. This requires a large number of experiments and is time consuming. Response surface methodology (RSM) is a combination of statistical and mathematical techniques based on a few experiments, which is useful for developing, improving and optimizing processes [13]. It is commonly applied in situations where several input variables potentially influence the response of a process. This technique was used in our previous work [7] for modeling cryogenic friction performance and will be used in the current work as well. The input variables are sliding speed (x1), load (x2) and sliding distance (x3) and the response is wear volume (y). The wear volume as a function of sliding speed, load and sliding distance can be expressed as y ¼ f ðx1 ,x2 ,x3 Þ þ e

ð2Þ

where e is a random error. The second order model is y ¼ b0 þ

k X i1

bi xi þ

k X i¼1

bii xi 2 þ

XX

bij xi xj þ e for io j

i

ð3Þ

j

The b parameters of the polynomials are estimated by method of least squares. The matrix approach has been adopted to solve Eq. (3) in which y is defined by a vector (n 1) and x is a matrix (n p) of independent variables; b is a vector ðp 1Þ of parameters (or regression coefficients) to be estimated and e is a vector (n 1) of errors. Eq. (3) may be rewritten in the matrix form as y ¼ bx þ e

ð4Þ

Fig. 2. Combined effect of load and speed on wear behaviour of Ti54: (a) ambient temperature and (b) cryogenic temperature.

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If b is replaced by matrix b, then Eq. (4) can be rearranged to the normal equation and expressed as

ð8Þ

Analysis of variance (ANOVA) was performed to determine the significant parameters and the validity of the equations for all models (Eqs. (7) and (8)) for a confidence level of 95%. The analysis of wear models (Eqs. (7) and (8)) is given below. The value ‘P’ in Table 3 is less than 0.05, indicating that the model is significant, which is desirable as it indicates that the terms in the model have a significant effect on the response [14]. From Table 3 also, it is obvious to recognize that the lack of fit of the full quadratic model was insignificant since the P-value is greater than 0.05. The significance of each coefficient was determined by the t-values and P-values, which are listed in Table 3. The large magnitude of t-value and the small magnitude of P-value indicate the high significance of the corresponding coefficient. The main effects of speed, load and the two level interactions of speed load (x1x2) and speed time (x1x3) and load time (x2x3) are significant model terms. The ANOVA was repeated for Ti54C (Eq. 8) and the results are given in Table 4. Eq. (8) is significantly based on the AVONA and two level interactions of speed load (x1x2) and speed time (x1x3) are important factors in controlling the wear volume. The R2 for Eqs. (7) and (8) are 98.44% and 92.74%, respectively. These equations are valid for sliding the titanium alloys against

Fig. 3. Combined effect of speed and time on wear behaviour of Ti54: (a) ambient temperature and (b) cryogenic temperature.

Fig. 4. The combined effect of load and time on wear behaviour of Ti54: (a) Ambient temperature (b) Cryogenic temperature.

xT ðxÞb ¼ xT y

ð5Þ

The solution of the normal equations Eq. (5) can be written as b ¼ ðxT xÞ1 xT y,

ð6Þ

where xT is the transpose of matrix x and (xTx) 1 is the inverse of (xTx). Calculation of the coefficient values was done using the MATLAB software. The equations of wear volume of Ti 5Al 4V 0.6Mo 0.4Fe (Ti54) for dry and cryogenic sliding in coded forms are given, respectively, in Eqs. (7) and (8) as yTi54D ¼ 1:6783þ 0:4325x1 þ 0:3753x2 þ 0:5644x3 0:0277x21 þ 0:0387x22

0:0277x23 þ 0:1314x1 x2 þ0:1877x1 x3 þ0:0938x2 x3

ð7Þ

yTi54C ¼ 1:361 þ0:364 x1 þ 0:2564 x2 þ 0:5202 x3 0:0202 x21 0:0069x22

0:0069 x23 þ 0:2158x1 x2 þ 0:1408 x1 x3 0:0469 x2 x3


tungsten carbide within the operating ranges of 0.1295o speed o0.9705 m/s, 6.464 oload o22.956 N and 2.636 osliding time o9.364 min. 4. Results and discussion Figs. 2–4 show the effects of various combinations of input (speed, load and time) on surface response (wear volume) of titanium alloys sliding against tungsten carbide under both dry and cryogenic conditions. The graphs were produced by using Eqs. (7) and (8). 4.1. Dry and cryogenic wear behaviour of titanium alloy Ti54 Eqs. (7) and (8) were used to plot the relationships between wear volume and speed, load and time for dry and cryogenic slidings of Ti54, as shown in Figs. 2–4. The combined effect of load and speed on wear volume of Ti54 under dry sliding condition is presented in Fig. 2a. The wear volume of titanium increases with increase in speed and load especially at higher levels, the combined effect of both factors is dramatically larger for both dry and cryogenic slidings. As can be seen, under dry sliding, wear volume of Ti54 changes marginally when load increases at the lowest speed (0.1295–0.3 m/s). In cryogenic sliding, the trend of wear volume of Ti54 is similar to dry sliding but with much lower magnitude. This obviously shows the impact of liquid nitrogen in reducing wear volume. Fig. 3 shows increase in trend of wear volume of Ti54 with increase in sliding time and speed under both dry and cryogenic sliding. The increase in speed is relatively less compared with the increase in time. At highest levels of time and speed, liquid nitrogen was able to reduce the wear by about 40% but at low

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levels of both factors, liquid nitrogen was not that effective, i.e. no much difference between dry sliding and cryogenic skidding at lower level of time and speed. This might be due the to rise in temperature under dry sliding, which increases with sliding time. Again the combined effect of both load and time on wear volume of Ti54D under dry sliding is relatively higher compared to cryogenic sliding, Fig. 4. Under dry sliding, wear volume increases linearly with increase in either load or sliding time but under cryogenic, nonlinear increase is the characteristic feature of wear volume with increase in either load or sliding time. It can also be seen that wear volume increases with increase in load, especially at longer sliding time and it is clear that the effect of load is less significant at shorter sliding time. Fig. 4b shows cryogenic results similar to the dry ones in which wear volume increases with load and time. It should be noted that the effect of time on wear volume is more significant in cryogenic sliding. It is commonly known that heat accumulation at the sliding interface causes softening of the material and if liquid nitrogen cannot neutralize the generated heat then the surfaces become weaker and vulnerable to wear. In general, wear volume of titanium alloy for both dry and cryogenic conditions increases with increase in sliding speed, applied load or sliding distance and this increase is more pronounced at higher levels of the variables. It is well known that interface temperature increases with increase in speed, load and sliding distance (time) [6]. In addition to increase in interface temperature, application of load increases the real contact area as well. On the other hand, mechanical properties of sliding surfaces are usually changing at higher interface temperature due to thermal softening [15]. This usually occurs when surface temperature exceeds 0.42Tm (melting temperature) of the alloys [16]. When surface temperature exceeds this value, it reduces the flow strength/hardness and therefore increases the wear rate [15].

Mag= 100 X

Detector: SE2 WD= 10 mm

Mag= 100 X


Mag= 100 X


Mag= 100 X


Fig. 5. SEM micrographs ( 100) of worn surfaces of Ti54 subjected to different sliding speeds at 14.715 N and 6 min: (a) dry sliding at 0.13 m/s, (b) dry sliding at 0.97 m/s, (c) cryogenic sliding at 0.13 m/s and (d) cryogenic sliding at 0.97 m/s.

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Besides this, increase in surface temperature also causes increase in contamination of the sliding surface due to oxidation [15]. Thus, the combined effect of any two factors of load, speed or time, at their highest level produces relatively higher wear. Under cryogenic sliding, liquid nitrogen can minimize heat generation and prevent thermal softening, resulting in prevention of adhesion wear and galling. However, when sliding speed/load increases, the rate of heat generated is greater than the heat dissipated to the surroundings by the liquid nitrogen jet. Besides this, when sliding distance (time) increases, the accumulated heat

Mag= 1.00 K X


10 µm

at the sliding interface becomes more and causes material softening. Therefore, liquid nitrogen jet cooling is less effective in the region of high load, high speed and long sliding distance. This phenomenon is observed from all the wear maps generated (Figs. 2–4), the combination of two factors at high level causes the highest wear volume. It is well known that sliding of metals produces very large plastic shear strains at the sliding interface and large strain gradients in the near surface material [17]. At high strain rates, metals and alloys frequently show narrow zones of highly

Mag= 1.00 K X


10 µm

Fig. 6. SEM micrographs ( 1000) of worn surfaces of Ti54 subjected to different sliding speeds at 14.715 N and 6 min: (a) dry sliding at 0.97 m/s and (b) cryogenic sliding at 0.97 m/s.

en

p

Mag= 500 X


Mag= 500 X


Mag= 500 X


Mag= 500 X


Fig. 7. SEM micrographs ( 500) of worn surfaces of Ti54 subjected to different loads at 0.55 m/s and 6 mins: (a) dry sliding at 4.46 N, (b) dry sliding at 22.96 N, (c) cryogenic sliding at 4.46 N and (d) cryogenic sliding at 22.96 N.


localized deformation, referred to as adiabatic shear bands (ASB) [18]. The localization of plastic flow in shear bands, which always leads to fracture of material, is usually attributed to the plastic instability resulting from thermal softening, which can overcome the effect of strain hardening in a deformed region. This instability can occur when the local rate of heat generation due to plastic flow exceeds the rate of heat dissipation to the surrounding materials [18]. The shear bands are favored initiation sites for failure, which occurs by void nucleation, growth and coalescence inside the thermally soften regions [19]. As mentioned earlier, titanium alloys have low thermal conductivity and are very likely to fail by adiabatic shear band when subjected to high strain rate deformation [20]. The injection of LN2 to the sliding interface works on controlling the generated heat by reducing the extent of local heat generation rate (due to plastic flow) compared to heat dissipation rate. Thus liquid nitrogen is expected to prevent the formation of adiabatic shear band and also hardens the titanium alloys [21]. This will be further examined microscopically in the following section.

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produce worn surface at low speed (Fig. 5c), wear particles were totally disappeared and only very few particles were remained on the worn surface at high speed (Fig. 5d). This is very likely because liquid nitrogen washed all wear debris away from the interface. Apparently, theses images of the worn surfaces of Ti54 (Fig. 5) have only abrasive marks and no cracks or other features of damage can be detected but when magnified images ( 1000) were produced, fine or tiny cracks and other features of damage on the worn surface were observed as shown in Fig. 6a and b. Surprisingly, tiny cracks were also detected on the worn surface even under dry sliding condition at high sliding speed as marked with dotted ellipse in Fig. 6a. This may indicate the inability of Ti54 alloy to plastic deformation. Fig. 7a–d compares the worn surfaces of Ti54 when subjected to different loads under dry and cryogenic sliding conditions. It seems

Ti54D

5. SEM observations of worn surfaces and wear debris 5.1. Worn surfaces of Ti54 alloy The effects of sliding speed on the worn surfaces of Ti54 sliding under dry and cryogenic conditions are shown in Fig. 5. On the worn surfaces produced by dry sliding at low and high speeds (Fig. 5a and b), other than abrasion marks occurring by microcutting with relatively narrow scars in the direction of sliding and some wear debris adhering to the worn surfaces, nothing else could be detected by SEM at a magnification of 100. When similar tests were conducted under cryogenic condition to

Mag= 150 X


Fig. 9. Wear debris collected after dry sliding of Ti54 at 0.55 m/s, 14.72 N and 9.36 min.

T

Mag= 500 X


Mag= 500 X


Mag= 500 X


Mag= 500 X


Fig. 8. SEM micrographs ( 500) of worn surfaces of Ti54 subjected to different sliding distances at 0.55 m/s and 14.72 N: (a) dry sliding at 86.99 m (2.64 min), (b) dry sliding at 309 m (9.364 min), (c) cryogenic sliding at 86.99 m (2.64 min) and (d) cryogenic sliding at 309 m (9.364 min).

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Fig. 10. SEM and EDS analysis of the wear debris collected from dry sliding test of Ti54: (a) SEM micrograph showing different types of wear debris and (b) EDS spectrum.

Fig. 11. SEM and EDS analysis of the wear debris collected from cryogenic sliding test of Ti54: (a) SEM micrograph showing different types of wear debris and (b) EDS spectrum.


that under dry sliding, the worn surfaces (Fig. 7a and b) are significantly deteriorated when the load increased from 4.46 to 22.96 N. Furthermore, wider and deeper ploughing marks can also be seen on the worn surface marked ‘p’ in Fig. 7b. This is due to the increase in actual contact areas during the sliding process when the load increased. Under cryogenic condition, when the load increased, no significant difference was observed as shown in Fig. 7c and d. In general, the influence of increase in speed, as discussed in previous section, was more pronounced compared to the influence of load. Fig. 7b shows wider ploughing grooves on the worn surface of Ti54, which produced by dry sliding at higher load (22.96 N). These grooves are well known as the characteristic features of abrasive wear mechanism produced sometimes by a third body, i.e. entrapped wear particles at the interface as shown in Fig. 7b marked with ‘en’. Under cryogenic sliding, at low load, ploughing marks and material displacement to both sides of the tracks are well clear but when the load was increased these features became very less, compare Fig. 7c and d. Different wear mechanisms were observed when Ti54 alloy was subjected to different sliding distances under dry and cryogenic conditions (Fig. 8). Under dry sliding and at shorter distance (Fig. 8a), fine wear debris and clear ploughing marks can be seen while at longer distance (Fig. 8b), the worn surface showed signs of three-body abrasion marked ‘T’ in Fig. 8b. Under cryogenic sliding, on the other hand, relatively smooth wear tracks were produced at shorter distance (Fig. 8c). Further increase in the sliding time showed in addition to tiny cracks more deteriorated surface (Fig. 8d), which was characterized by brittle fracture of the alloy. The hardening effect due to interposing liquid nitrogen became more obvious when Ti54 was subjected to a longer sliding distance (Fig. 8d). Again this may confirm that Ti54 became harder and more brittle, which in turn leads to cracks and delamination as seen in Fig. 8d. Considering the worn surfaces of titanium alloy Ti54 tested at both dry and cryogenic conditions (Figs. 5–8), it is clear that different wear mechanisms occurred in different environments. Under dry sliding condition, sliding motion between the titanium alloy and tungsten carbide generated heat in both surfaces. Titanium alloy, because of its low thermal conductivity, dissipates the accumulated heat with a lower rate than the rate of heat generation, especially at high levels of speed and load. As a result, adiabatic shear band is likely formed in titanium alloys under dry sliding condition. Further strain (by further sliding) causes void nucleation within the shear band. The voids grow further and eventually coalesce, creating complete separation [19]. As a result, delamination occurs and flake like metallic debris dominate in dry sliding. The flakes like debris were reported earlier to occur in sliding condition conducive to adiabatic type shear localization phenomena [22]. Besides this, high interface temperature also may cause change in microstructure and thermal softening [16]. In sliding under cryogenic condition, interface temperature was decreased by interposing liquid nitrogen and thus liquid nitrogen prevents/delays the formation of adiabatic shear band and delamination wear. Besides this, liquid nitrogen, because of low temperature, hardens the surface of titanium alloy [21]. This resulted in abrasion wear by microcutting of titanium alloys on the worn surfaces.

5.2. Analysis of wear debris Fig. 9 shows micrographs of wear debris collected after dry test for titanium alloy Ti54. The wear debris is in the form of metallic flake. The lengths of the plate like debris are 250–350 mm and the width is 50–70 mm. In addition to plate like debris, a small amount of fine debris can also be seen. Furthermore, the flake like debris dominate.

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EDS analysis of wear debris collected from dry sliding of Ti54 (Fig. 10) shows that the main elements of all the debris are Ti and Al. Moreover, the plates like wear debris are confirmed to be from the titanium alloy Ti54, suggesting that delamination wear mode dominated during dry sliding process. It is also observed that small portion of tungsten (W) was detected from the wear debris and definitely they are from tungsten carbide of the grinding wheel. Wear debris collected from cryogenic tests of Ti54 and their EDS analysis are shown in Fig. 11. Cryogenic sliding with liquid nitrogen induced wear debris in the form of long chip with length of 500 mm and width of 40 mm (Fig. 11a). EDS analysis (Fig. 11b) on the wear debris of Ti54 confirmed that the cutting chip like wear debris were from the titanium alloys. This shows that microcutting (abrasion wear) is likely the main wear mode. The debris collected from the cryogenic test was mainly in the form of cutting chips of cylindrical shapes. The worn surface also shows brittle fracture feature. This was due to sliding and the impact occurring during the tests under cryogenic condition; titanium alloys behave in a brittle fashion and become more sensitive to repeated impact compared to tougher materials [23].

6. Conclusions (i) The relationship between wear volume and applied load, sliding velocity and sliding distance (sliding time) has been successfully developed using RSM. The predicted results are in good agreement with the measured ones. These relationships are applicable within the ranges of tested parameters. (ii) In general, wear volume increases with load, sliding speed and sliding time. Under cryogenic sliding, wear volume is consistently lower than wear volume in dry sliding condition. Furthermore, load and speed influence the wear with greater extent at longer distance. (iii) Different wear mechanisms were observed, i.e. in dry sliding, delamination wear due to formation of ASB is the dominated wear mechanism under dry and cryogenic conditions. Liquid nitrogen hardens the titanium alloys and prevents/delays thermal softening. Under cryogenic sliding, the main wear mechanism was abrasion; besides, some evidences of threebody abrasion due to entrapped wear particles were observed. Cracks and fracture of brittle nature were detected also under cryogenic sliding due to alteration of titanium mechanical properties. (iv) Plates like wear debris, collected from dry sliding tests, were confirmed to be from the titanium alloy Ti54, which suggested that delamination wear mode dominated during the dry sliding process. Under cryogenic condition, wear debris were in the form of cylindrical shapes.

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