Fabrication self-recovery bulge textures on TiNi shape memory alloy and its tribological properties in lubricated sliding

Tribology International 96 (2016) 11–22

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Fabrication self-recovery bulge textures on TiNi shape memory alloy and its tribological properties in lubricated sliding Guanghai Tang a,b, John KL Ho b,n, Guangneng Dong a,nn, Meng Hua b a b

Key Laboratory of Education Ministry for Modern Design and Rotor-Bearing System, Xi’an Jiaotong University, Xi’an 710049, PR China Department of Mechanical and Biomedical Engineering, City University of Hong Kong, Hong Kong, PR China

art ic l e i nf o

a b s t r a c t

Article history: Received 20 May 2015 Received in revised form 16 November 2015 Accepted 13 December 2015 Available online 18 December 2015

Effect of self-recovery bulge texture on tribological properties of TiNi under oil-lubricated condition was investigated. Fabrication of bulge textures on TiNi alloy was performed by a method with a sequence of indentation–polishing–heating. Due to the different levels of transformation of martensite to austenite, XRD pattern for the bulged TiNi was different after process of indentation–polishing–heating. Friction coefficient of the bulged TiNi was obviously lower than that of non-bulged counterpart. Study illustrated test conditions and height of bulges effectively varied the tribological behaviors of TiNi specimens. SEM analysis indicated that scratches produced by abrasive and oxidation wear were the major wear mechanisms of the bulged TiNi specimens. Furthermore, the implementation of bulge textures significantly improved the tribological properties of TiNi alloy. & 2015 Elsevier Ltd. All rights reserved.

Keywords: Titanium nickel Bulge texture Sliding friction Wear

1. Introduction The favorable properties like shape memory effect, pseudoelasticity and corrosion resistance of titanium nickel (TiNi) shape memory alloy have drawn extensive scientific research interests in the past decades. On the basis of these advantageous properties, TiNi alloy has been widely used as the material for designing and fabricating functional structures in most mechanical, bio-medical engineering and micro-electromechanical systems (MEMS) [1–3]. As the structural material in these systems may involve with sliding relatively over each other, it is regularly associated with the possible occurrence of interfacial friction and wear. To sustain the structural reliability and to save energy/power consumption, TiNi alloy is thus also expected to exhibit good friction performance and high wear resistance. However, TiNi alloy usually has poor processing performance and difficulty to lubricate. As a result, it is generally vulnerable to fail when it is used as tribological parts. Hence, effective and efficient use of TiNi alloy in the aforementioned systems requires the improvement of tribological performance to certain level. Under most operational conditions, material failure due to either the individual mode or the combined modes of severe friction, wear damage, erosion and fatigue always takes place at the contact interface [4]. Available literature [5–7] has indicated that effective implementation of different surface n

Corresponding author. Tel.: þ 852 3442 8425. Corresponding author. Tel.: þ 0086 29 82668552. E-mail addresses: [email protected] (J.K. Ho), [email protected] (G. Dong). nn

http://dx.doi.org/10.1016/j.triboint.2015.12.011 0301-679X/& 2015 Elsevier Ltd. All rights reserved.

textures on a surface can reduce friction, lower the rise of surface temperature and improve wear resistance. Hence, it is expected that suitable modification of TiNi shape memory alloy can improve its tribological properties specifically for engineering applications. There are many surface modification techniques available for texture processing in the existing literature. Typical approaches involve with milling with machine tool [8], irradiation with energy beam [9] and chemical etching [10], etc., which are mainly for manufacturing micro sunk structures on the surface of materials. Many available studies have demonstrated that groove and dimple textures exhibited good tribological performance under different experiment conditions [11,12]. Furthermore, under wet lubrication and sliding condition, lubricant in the sliding dimple depression may flow to create local hydrodynamic pressure, which likely generates upthrust to separate two mating surfaces [13]. Such thrust due to surface texture may contribute to increase the load carrying capacity of the sliding surfaces. However, study of Amanov, et al. [14] indicated that the friction and wear behaviors of dimpled specimens were improved further with the presence of certain level of bulge around the rim of dimples. Studies [15,16] also indicated that the suitable fabrication of protruding textures could also improve tribological behaviors of mating surfaces. Comparatively, it is anticipated that the tribological performance of protruding textures may be more superior to the dimpled surface as they provide channels for efficient draining of wear debris. However, most traditional techniques in the aforementioned literature are hardly used to manufacture protruding textures on a surface, and production of protrusion bulges, semi-spherical or dome textures thus requires the use of non-conventional

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techniques. Ando and Ino [17] used focusing ion beam to mill regular isotropic asperities on a silicon substrate. Stephens et al [18] applied LIGA technology to fabricate deterministic micro asperities on bearing and seals. Tang et al [19] combined hotfilament-assisted chemical vapor deposition and ion beam etching techniques to fabricate nano-tips on silicon substrate. The use of their texturing techniques [17–19] significantly improved the tribological properties of materials. As the protruding micro textures are normally used non-conventional techniques, which commonly require expensive equipment and are difficult to control for achieving any geometry of protruding textures. Thus, other economic and rapid approaches are necessary to be sought. Consequently, specifically use of TiNi shape memory alloy for fabricating the protruding textures is herewith proposed. In recent years, study on deformation recovery due to the shape memory effect of TiNi has been conducted. Generally, the plastically distorted TiNi shape memory alloy tended to experience partial deformation recovery to its original shape when it was under some form of stimulus [20]. The deformation recovery for different indented morphologies would be different. Typically, the recovery of the deformed TiNi alloy was relatively lower with pyramidal indenter than that with spherical indenter. This is mainly due to the different magnitude and spatial distribution of the induced stress [21]. The deformation recovery morphology of TiNi alloy usually varies with the recovering temperature [22,23]. Proper control of the level of recovering temperature allows the producing of different protruding morphologies on its surface. Results of our previous study [24] facilitated the identification of thermally induced partial deformation recovery on worn TiNi surface. Such phenomenon of recovery was possibly simulated/re-produced by suitable conjunction of the indentation with a process of mechanical polishing and heating treatment in sequence. It also affirmed that the implementation of bulge morphology on the surface of TiNi shape memory alloy by take advantage of deformation recovery was feasible. Such approach is feasibly used to fabricate various arrangements, geometries and shapes of various protruding textures on TiNi shape memory alloy. In view that tribological properties of a surface with bulge textures are in fact not yet investigated, it is thus anticipated that systematic study of the deformation recovery characteristics and their effect on tribological performance may furnish data to be suitably considered/ controlled in the engineering applications. This study was thus focused on fabricating bulge textures on TiNi shape memory alloy by a sequence of indentation–polishing– heating (IPH). It was then followed by a series of friction tests so as to investigate its influence on the friction and wear properties of TiNi alloy. Subsequently, the relationships between contact pressure, reciprocating speed and friction coefficient were established for the bulged textures on TiNi shape memory alloy. Detail analysis of friction and wear behaviors allowed the derivation of the associated mechanisms, which were then discussed in detail.

2. Material and methods

illustrated the martensitic finish phase transformation (Mf) on cooling to a temperature of about 18.3 °C and the austenite finish phase transformation (Af) on heating to a temperature of about 38.5 °C, as reported in [24]. According to DSC result, TiNi alloy was a mixture of martensite and austenite phases at room temperature of 20 72 °C. Prior to and post of processing of the anticipated textures (Section 2.2), the machined TiNi specimens were properly cleaned with acetone in an ultrasonic bath for 10 min, which was followed by careful wipe-cleaning with alcohol and then dried with a blower at room temperature. 2.2. Texture processing Fig. 1 shows the major procedural schematic in processing the self-recovery bulge textures on the surface of TiNi specimens. A 45# steel disc, on which a large amount of 1 mm diameter (2R) GCr15 balls were fixed (Fig. 1a), was mounted to Instron material testing machine, model 5569, with limit load of 100 kN. The so assembled Instron machine was used to compress/indent the surface of TiNi alloy specimens at room temperature condition. Three different compressions/indentations were performed individually with maximum load (F) to be set at 15, 20 and 25 kN, respectively. The loading rate was set at 0.2 mm/min under room temperature condition. The achievable indentation depth varied with compressive load, and tended to follow the trace of the individual curves as shown in Fig. 2. Due to pseudo-elastic recovery, the residual depth of the indented impression after unloading was subsequently recovered to h1 (Fig. 1b) from its insitu depth h0 before unloading. The indented impressions on the surface of TiNi alloy were arranged in a square array. Then, the indented surface was carefully and wetly polished with water at room temperature using 400#, 800# and 1000# metallographic sand papers until the verge of indent morphology disappeared (Fig. 1c). Attention was particularly paid to ensure the parallelism and uniformity throughout the polished surface. The polished TiNi specimens had surface roughness (Ra) about 0.135 μm, and were heated at 100 °C for 10 min in a KTL1600 vacuum tube furnace, China. Such heating process facilitated the appearance of bulge morphology (h2 – height of bulge) on the surface of TiNi shape memory alloy (Fig. 1d). An S-3000N scanning electron microscopy (SEM), Hitachi, Japan, was used to analyze the 2D morphology of bulged TiNi, while an OLS4000 confocal microscope, OLYMPUS, Japan, was used to observe its 3D counterparts and to measure the corresponding values of bulged surface height and diameter. The geometry of bulges and their spacing within a sampling length of 6.4 mm (Fig. 1e) which had been indented under a loading of F¼15 kN was measured. The measurements gave: (i) the value of bulge spacing (L) as 2.0 mm, (ii) the distribution of bulge diameter (2a1) and peak height (h2) for Sb1 to be scattering in range of 790– 820 μm and 23.0–24.5 μm, respectively and (iii) the corresponding skewness of the bulged surface peak height to be about 0.3, which is benefit to store lubrication oil. A D/MAX-2400 X-ray diffractometer (XRD), RIGAKU, Japan, was used to investigate the microstructural changes on the surface layer of non-bulged and bulged TiNi specimen after IPH process.

2.1. Material details 2.3. Friction testing procedure Commercial TiNi alloys which were prepared by vacuum induction melting (VIM) in graphite crucible furnace, were purchased from Saite Technology Co., Ltd., Xi’an, China. The asreceived 80 mm diameter  100 mm length TiNi alloy was then machined to dimensions of 5 mm thick  30 mm diameter circular plate specimens. The circular plate TiNi alloy was analyzed by energy dispersive X-ray spectroscopy and found to have chemical compositions in wt% as 55.44% Ni and 44.56% Ti. Scanning with a DSC2910 differential scanning calorimeter, TA Instruments, USA,

Tribological properties of TiNi alloy with bulge textures under oil-lubricated condition were investigated using a UMT-2 multifunctional tribo-tester, CETR, USA, with a pin-on-plate configuration at room temperature. As shown in Fig. 3, the Φ30  5 mm TiNi circular plate specimens with bulge textures having hardness of 269 Hv was used as a counter-mating material slid against the Φ6  15 mm static pin of 45# steel having hardness of 490 Hv. Tests were performed with the steel pin to be mounted into a pin

G. Tang et al. / Tribology International 96 (2016) 11–22

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Fig. 1. Schematic illustration of the IPH process.

Table 1 Friction tests parameters. Contact pressure P (MPa)

0.1, 0.4, 0.5, 0.7 and 1.0 (2.83, 11.32, 14.13, 19.81 and 28.3 N) Reciprocating speed (mm/s) 6.0, 12.0, 24.0,48.0 and 84.0 (0.5, 1.0, 2.0, 4.0 and 7.0 Hz) Lubricant Castor oil Oil supply (ml) 10 Room temperature (°C) 20 72 Viscosity 706.1 mm2 s  1 at 25 °C

Fig. 2. Typical indentation depth depending on compressive load curves at various loads, and the experimental setup is illustrated in the inset.

Fig. 3. Schematic representation of the pin on circle plate tribometer.

holder which was directly attached with a load sensor system to record friction forces during the friction tests. For achieving same contact condition with TiNi circular plate specimens, the upper steel pins were also polished sequentially with 400# and 800# grit abrasive papers under a contact pressure of 0.35 MPa and reciprocating sliding with a frequency of 2.0 Hz for 10 min. The so polished pin was then ultrasonically cleaned in acetone and its surface roughness was measured as Ra 0.22 μm. The pin was subsequently mounted at central region of the TiNi circular plate specimen, and slid with a stroke of 6.0 mm and reciprocating sliding frequencies of 0.5–7.0 Hz so as to give equivalent average velocity in range of 6.0–84.0 mm/s. Tests were performed in a laboratory having relative humidity of 30 72% at room temperature. Other settings of the friction tests were tabulated in Table 1. After the friction and wear tests, the corresponding worn tracks were observed and/or analyzed respectively by the SEM with an energy dispersive X-ray spectrometer (EDS), Oxford, UK. A PHI 5820 X-ray photoelectron spectroscopy (XPS) by Physical Electronics, Chanhassen MN, USA, was also used to analyze the chemical elements on the wear tracks of both non-bulged and bulged specimens individually slid for two hours. The analysis was undertaken over a small area of 800 μm  400 μm on the whole wear track. During operation, the power source of XPS was set at a 1486.6 eV AlKα mode with all data acquisitions to be powered at 350 W. The chemical composition analysis was performed with

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the XPS to set at high-resolution spectra mode and to be processed by the use of a XPS-Peak4.1 software.

3. Results and discussion 3.1. Morphology of self-recovery bulges and X-ray analysis The 2D and 3D morphologies of a typical TiNi specimen with bulge textures produced with indentation force of 25 kN, and the associated cross-sectional profile are illustrated in Fig. 4. It obviously reveals the successful production of bulge textures on the surface of TiNi alloy with the IPH method (Fig. 4a), which is also confirmed by the contours of the single bulge morphologies of 2D image (Fig. 4b) and 3D image (Fig. 4c), as analyzed with a confocal microscope. The plot of cross-sectional profile (Fig. 4d) allows the height and diameter of the bulge to be estimated as about 31 μm and 900 μm, respectively. Geometric parameters of all the bulged and non-bulged test specimens, either measured or estimated, are tabulated in Table 2. The variation of bulge height under different indent loads is mainly attributed to the distribution of induced stress due to different levels of indentation and polishing process [24–26]. Fig. 5 presents the XRD patterns of the non-bulged and bulged TiNi specimens prior to tribological tests. Fundamentally, there is not any difference in phase composition between the non-bulged and bulged TiNi to be observed. However, the patterns suggest that the phase in all TiNi specimens is mainly austenite (110)B2

phase embedding with minor level of martensitic variants over the range of 30–80 °. TiNi specimen post of IPH process significantly increases the diffraction intensity peak of (110)B2 (2θ ¼42.2°) because some martensitic phases have been thermally transformed to the austenitic phases [24]. Such phase transformation results in the possession of better pseudo-elasticity in the bulged TiNi than the non-bulged counterpart. According to the study of Jin et al [27], TiNi alloys in a state of pseudo-elasticity usually could absorb larger elastic deformation energy within contact and Table 2 Geometric parameters of all the bulged and non-bulged specimens used in these tests. Parameters

Nonbulged

Bulged specimens Sb1 (F¼ 15 kN) Sb2 (F¼ 25 kN) Sb3 (F¼ 20 kN)

Initial surface roughness Ra Spacing of bulge L (mm) Contact radius 2a1 (μm) Height of bulges h2 (μm) Radius of curvature cr (μm) h2/a1 h2/R (R-radius of GCr15 ball)

0.135

0.135

0.135

0.135

—

2.0

2.0

2.0

—

810

900

840

—

24.3

31.0

34.1

—

3177.4

3435.2

3594.8

— —

0.06 0.048

0.069 0.062

0.081 0.068

Fig. 4. 2D SEM image (a) of TiNi specimen with bulge textures after IPH process and the maximum indentation force F ¼ 25 kN, (b) 2D, (c) 3D morphologies, and crosssectional profile (d) of a single bulge.

G. Tang et al. / Tribology International 96 (2016) 11–22

Fig. 5. X-ray patterns of bulged and non-bulged TiNi specimens.

subsequently spring-back to recover when exerting load was reducing. Such spring-back recovery tends to exert a push to release two contact surfaces and thus results in less material volumetric loss in wear process. The appearance of additional diffraction peak (200)M (2θ ¼61.5°) of martensitic variant on the bulged specimen is envisaged as a result of a mechanism due to the nature of corresponding deformation. Generally, the representative strain for the domed indentation was expressed as 0.2 a/ R [28], where a was the contact radius of indentation and R was the radius of indenter (Fig. 1a). The study of Ni et al [29] further correlated the ratio of maximum indentation depth (hmax) and radius of indenter (R), hmax/R, to equal to the parameter of 0.2 a/R. Their studies also showed that slip plasticity occurred in martensitic phase and dislocations remained after heating to a temperature above Af when the representative strain was beyond 5%. Practically, the height of bulge produced by the process of IPH was commonly shorter than the depth of the corresponding indent [22–24]. In this study, the bulges on TiNi specimen produced with maximum indent load of 25 kN gives (i) the ratio of h2/R equals to 0.062 and (ii) a representative strain to be higher than 5%, it thus exerts more deformation energy to result in severer dislocation of martensitic phase which, to certain extent, tends to suppress the complete transformation of martensitic phase to austenitic phase. Consequently, the additional residual martensitic variant thus remains to exist in the bulged TiNi specimen even after heating. 3.2. Effect of bulge textures on the friction property Fig. 6 compares the variation of friction coefficient with sliding time of the non-bulged and bulged TiNi (Sb1, Sb2, and Sb3 – Table 2) specimens slid at reciprocating speed of 48 mm/s under contact pressure of 0.1 MPa. The friction coefficient curves of nonbulged and bulged specimens illustrate a trend to decrease quickly at the initial stage, which is followed by distinctly different behaviors depending on the height of bulges (h2 – as shown in Table 2). Although the friction coefficient curve for non-bulged TiNi alloy specimen exhibits certain level of mild fluctuation, it is then followed by stabilizing at a value of 0.2 from sliding time 1500 s onward. The friction coefficient curves of Sb1 and Sb2 seem to exhibit a similar trend of (i) decreasing to minimal value at sliding time about 600 s for Sb1 and about 300 s for Sb2 and (ii) slowly climbing up to a value about 0.17 for Sb1 and about 0.13 for Sb2. For the bulged specimen Sb3, the friction coefficient curve shows a tendency to decrease gradually and continuously to the end of the test after its rapid drop when sliding begins.

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Fig. 6. Variation of friction coefficient of the non-bulged and bulged TiNi as a function of sliding time.

Fig. 6 illustrates that the friction coefficients of all three bulged (Sb1, Sb2 and Sb3, respectively) TiNi specimens appear to be about 13.8%, 40.5% and 44.8%, respectively, lower than that of nonbulged specimen. Usually, the frictional force is derived from two major sources: (i) the adhesion force resulted from the atomic interaction between two contacting surfaces and (ii) the force to deform/shear the interacting asperities or to fracture/plough the surface by asperities and/or wear particles [30]. Value of friction coefficient (μ) may approximately be expressed as a sum of two components μ ¼ μa þ μp, where μa is the contribution of adhesion and μp is that due to the deformation/plough [31]. The friction coefficient due to either adhesion or deformation/plough is mainly dependent on the material properties and surface asperity peaks of the pair of contacting surfaces. Value of μa could be adequately reduced by suitably fabricating the bulge textures so as to minimize effectively the real contact area between two mating surfaces. As mentioned in Section 3.1, the TiNi alloy with bulge textures tends to have better pseudo-elasticity than their non-bulged counterpart (Fig. 5) under the compression contact mode. The bulge shape starts to restore deformation elastically when contact load is releasing. Furthermore, the creation of hydrodynamic flow of lubricant over the bulged profile may further thrust up mating surface in diminishing surface contact. The integrative action of pseudo-elasticity and hydrodynamic up-thrust could thus decrease partially the value of μp. Consequently, the implementation of bulged textures on TiNi alloy specimens tends to furnish with lower value of friction coefficient than that of their non-bulged counterpart (Fig. 6). The divergent spacing between the curves for Sb3 and Sb2 suggests a transition of height of bulge occurs between 34.1 μm and 31.0 μm, from which onward the dominant effect of pseudo-elasticity and hydrodynamic up-thrust inclines to be magnified positively. To understand further the effect of bulge textures on the tribological properties of TiNi alloy, a series of further friction tests slid with reciprocating speed of 6.0–84.0 mm/s were carried out under the contact pressure of 0.1–1.0 MPa. Fig. 7a compares the trend of the average friction coefficient varying with the reciprocating speed under a contact pressure of 0.1 MPa. The curve for non-bulged specimen shows a tendency of rapid increase when reciprocating speed increases to 12.0 mm/s. It is followed by a sharp drop in friction coefficient when the reciprocating speed further increases. Friction coefficients of both bulged Sb1 and Sb2 show similar trend, and have value lower than that of non-bulged specimen. Although the distribution of their friction coefficients over the range of tested reciprocating speeds exhibits diminishing trend, such decreasing character looks relatively consistent. The value of friction coefficient ratio (ημ) may approximately be

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expressed as ημ ¼(μnon-bulged–μbulged)/μnon-bulged, where μbulged and μnon-bulged are the friction coefficient of bulged and nonbulged specimens, respectively. The value of ημ is about 1.8– 20.42% for Sb1 and about 32.09–47.07% for Sb2 (Fig. 7b). The trend of friction coefficient for the bulged specimen Sb3 initially drops very sharply as the reciprocating speed increases from 6.0 mm/s to 12.0 mm/s. It then increases slightly to a reciprocating speed of 24.0 mm/s, and subsequently resumes decreasing trend with the value of ημ ranging between 27.7% and 55.3% with further increases in reciprocating speed to 84.0 mm/s. Generally, the friction coefficient is varying with and controllable by both surface and bulk properties according to the following rule μ ¼ τ/Hb, where τ is shear strength at contact surface and Hb is the hardness of the bulged TiNi [31,32]. Hence, the drop of friction coefficient values with the increase of reciprocating speed is mainly attributed to the reduction of shear strength at the contact interface. Such reduction may be a result of forming oil film between the pin and bulges [14], which, under sliding condition, subsequently creates higher hydrodynamic up-thrust to reduce interfacial contact. Moreover, sliding activating flow of oil lubricant tends to dissipate friction heat away more rapidly from contact interface with increasing of reciprocating speed. Heat dissipation in such way constrains the surface temperature rise to soften the material and keeps the friction coefficient to certain level.

The change behaviors of average friction coefficient with contact pressure for the non-bulged and the bulged TiNi alloy specimens are plotted in Fig. 8. Certainly, friction coefficient of nonbulged TiNi alloy clearly demonstrates to be slowly increasing over the range of contact pressure (Fig. 8a). In aspects of surface-tosurface contact, a large number of asperities are pressed against each other. However, the effect of pseudo-elasticity on asperities deformation is weak, and the asperities wear out quickly. This phenomenon of non-bulged TiNi specimen as shown in Fig. 8a can be elucidated partially as the likely closer actual interfacial contact area with the increase in contact pressure. Different behaviors of friction coefficient are observed for the bulged specimens. Typically, the friction coefficient for Sb1 has a minimum value of 0.15, and its value is increasing when the contact pressure is higher than 0.4 MPa. The trend of its friction curve resumes a sagged characteristic. Moreover, it gives the value of ημ ranging between 13.76 and 25.35% (Fig. 8b). Comparatively, the trend of friction coefficient for Sb2 resumes similar characteristics as those of the bulged Sb1 except its level of sagging is relatively small and its value of ημ to be in the range of 36.68–45.43%. However, the friction coefficient curve for the Sb3 is evidently fluctuating within the range of 0.11–0.16 with a lowest value occurring at the contact pressure of 0.1 MPa (Fig. 8a). In addition, its value of ημ falls within the range of 25.34–43.56%. Results illustrate that the friction coefficient of bulged specimens increases as the contact pressure is

Fig. 7. (a) Average friction coefficient with respect to the contact pressure at reciprocating speed of 48 mm/s, (b) the value of friction coefficient ratio for all bulged specimens under different reciprocating speed.

Fig. 8. (a) Average friction coefficient with respect to the reciprocating speed at contact pressure of 0.1 MPa, (b) the value of friction coefficient ratio for all bulged specimens under different contact pressures.

G. Tang et al. / Tribology International 96 (2016) 11–22

increasing from 0.4 to 1.0 MPa, and such increase mainly because of the reduction in the effect of pseudo-elasticity and the increase real contact between the pin and the bulges. Elastic deformation of a micro bulge (δ) at low contact pressure P when it is contacting with a rigid plane could be evaluated with the following expression δ ¼(9P2/16E2cr)1/3, where E is elastic modulus and cr is curvature radius of the bulge [33]. As cr increase the value of δ decreases, and the real contact area between the pin and the bulged specimens thus also decreases. The friction coefficient of bulged specimen with height of 34.1 μm (i.e., bulged specimen Sb3 – Table 2) gives lowest value under reciprocating speed of 48.0 mm/s when it is compared with its counterparts with the heights of 24.3 μm and 31.0 μm (Fig. 7a). With the contact pressure higher than 0.4 MPa, the deformation of the single bulge may be in elastic-plastic state, which results in non-uniform distribution of strain and higher level of permanent deformation. Such permanent deformation may change the side profile of the bulges and subsequently diminish hydrodynamic effect. Among the three types of bulged TiNi alloy specimens, the Sb2 specimen gives the lowest friction coefficient as the contact pressure increases from about 0.4 MPa onward (Fig. 8a). This may be attributed to the resulted mechanical properties of bulges due to (i) IPH process in achieving its bulge height and (ii) the net effect of hydrodynamic performance. The linear and horizontal trends of its friction coefficient curve may imply the balance off of the detrimental effect due to the increase in level of deformation with the favorable effect by the IPH mechanical properties and hydrodynamic performance. 3.3. Analysis of the worn surface Fig. 9 exhibits the SEM images of wear scars on the surface of both non-bulged and Sb3 bulged TiNi specimens rubbed against 45# steel pin under contact pressure of 0.1 MPa and reciprocating speed of 48.0 mm/s. The two specimens illustrate the occurrence of distinctly different wear phenomena. Very rough with appearance of ploughed furrows along the sliding direction is observed on the rubbed surface of non-bulged specimen (Fig. 9a). The rubbed scar also covered with many irregular abrasion marks (Fig. 9c), which clearly suggests the wear mode to be abrasive in nature and mainly attributed to plastic deformation of friction surface [34]. The bulge textures on the surface of Sb3 still remain certain level of intact and are not completely worn out (Fig. 9b), implying pin surface not in direct and close contact with some portions on the bulges at low contact pressure. Furthermore, the intact regions at either end of wear scar of bulges are almost having similar coverage, suggesting similar hydrodynamic effect to be generated in either sliding direction of the reciprocating movement. Scratches on the worn scar of bulges appear to be smooth and flat (Fig. 9d) without any sight of gathering wear debris along the slid wear track. The absence of wear debris on worn surface prevents the occurrence of delamination wear; hence, the wear behavior of materials is significantly improved. Fig. 9e illustrates the profiles across the wear scars (A–A as shown in Fig. 9a and b) of both the non-bulged and that of a single bulge on Sb3 specimens. The profiles give the maximum depth of wear tracks for non-bulged and a single bulge to be about  27 μm and  13.1 μm, respectively, which surely confirms the significant enhancement of wear resistance for the bulged TiNi specimens in comparison with that of the non-bulged counterpart. Comparing the 3D morphologies of a single bulge on the bulged specimens slid under contact pressure of 1.0 MPa and reciprocating speed of 48.0 mm/s as shown in Fig. 10. The results suggest: (i) some portion of the bulge on the Sb1 specimen is completely worn out, and the depth of worn track is about –5.0 μm (Fig.10a), (ii) the 3D worn morphology shows that residual bulge with

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height about 12.8 μm for Sb2 and about 15.9 μm for Sb3, respectively, is still to be observed (Fig. 10b and c) and (iii) wear tracks existed around the bulge is mainly attributed to the rough surface of pin during sliding process and the area decreases with the height of bulge increases. The residual bulges in (ii) above indicate that they are not completely worn out, and it can be seen that the cross section profiles of bulges appear smooth. This feature on the wear scars is attributed to accelerate running in the contact interface between the pin and bulged TiNi specimens at high contact pressure. In addition, correlating the results in Figs. 9e and 10c also indicate that increase in contact pressure enlarges the contact area between the pin and bulged TiNi specimens and subsequently leads to increase the friction coefficient. Furthermore, two hours sliding friction tests on both nonbulged and bulged TiNi specimens, specifically for investigating the ability and duration of bulged textures to maintain at low friction coefficient values, are also conducted. Results for all the specimens slid at a reciprocating speed of 48 mm/s and contact pressure of 0.5 MPa are shown in Fig. 11a. Friction coefficient curves for both bulged (Sb2 and Sb3) and non-bulged specimens seem to resume a constant value trend without any sign of fluctuation throughout the test. Typically, it gives value of about 0.13 for Sb2, about 0.14 for Sb3, and about 0.21 for non-bulged. Furthermore, the trend of their individual curves looks rather smooth although the coefficient curve for bulged specimen Sb1 seems to increase slowly after being slid for 3000 s. Such increasing trend is identified as a result of enlarging the contact area. Comparing the results in Figs. 6 and 11a, it clearly indicates that the friction coefficient for the bulged TiNi specimens tends to acquire a stable value relatively quicker and subsequently to maintain at low value magnitude of friction coefficient for a long period of time. Such analysis of slid surfaces was unable to reveal any obvious material loss or transfer from the mating pin. Aiming at understanding the contribution of pin wear to the friction coefficient of TiNi discs, the micrographs of pin before and after 2 h sliding friction tests were subsequently analyzed. Pin surface morphology before and after sliding against non-bulged and bulged specimens is shown in Fig. 11b and c. As the upper pins are suitably polished with abrasive papers for 10 min with a contact pressure of 0.35 MPa and a reciprocating speed of 24 mm/s, their pre-slid surfaces are generally rough and also scattering with some scattering pits. Their counterpart surface slid against non-bulged TiNi specimen appears to have some irregular abrasion marks along sliding direction (Fig. 11b). Worn surface (Fig. 11c) of their counterpart slid against bulged TiNi specimen looks smooth and flat with sign of some randomly distributed pits. Although the analysis do not show any obvious difference among the slid pin surfaces, difference in real contact area between the pin and individual TiNi discs could be considered as the major attribution to the increase and decrease in friction coefficient for the corresponding nonbulged and bulged specimens (Fig. 11a). Chemical compositions in the wear scars were analyzed with an EDS. Spectra of various elements of the non-bulged and bulged TiNi specimens (i.e., zone of B as shown in Figs. 9c and d) are shown in Fig. 12a and b. A small amount of about 22.23 at% Fe and 3.23 at% Cr is detected on the wear track of non-bulged TiNi suggesting the taking place of adhesive wear transferring metal from one to each other [30]. However, analysis also reveals neither chemical change nor element transfer from 45# steel pin is detected on the bulged specimen. Moreover, about 34.86 at% of oxygen (O) under a 0.1 MPa pressure is detected on the wear track of bulged TiNi specimen, indicating the occurrence of oxidation in the sliding process. Usually, the formation of oxide layer prevents direct metal-to-metal contact. If such oxide layer was harder than base material of two mating surfaces, it served to reduce friction and wear [35,36]. Hence, the oxide layer on bulge textures

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Fig. 9. SEM images of the worn surface for the non-bulged ((a) (40  )) and ((c) (500  )) and bulged ((b) (70  )) and ((d) (500  )) specimen Sb3, (e) Wear track profiles (A– A) of the non-bulged and that of a single bulge on slid TiNi as shown in (a) and (b).

evidently depletes to reduce friction and wear of the specimens. To gain further insights on the chemical compositions of the slid surfaces, XPS analysis of worn specimens (in Fig. 11a) was performed specifically on the surface of TiNi specimens. XPS spectra of both non-bulged and bulged (Sb3) specimens, respectively, slid under contact pressure of 0.5 MPa and reciprocating speed of 48 mm/s are illustrated in Fig. 12c. It shows the existence of titanium (Ti 2 p), nickel (Ni 2 p), oxygen (O 1 s), and iron (Fe 2 p) peaks on all samples, and the existence calcium (Ca 2 p) on the bulged specimens only. Two obvious peaks individually at

458.9 eV and 464.7 eV are detected from the Ti 2 p spectrum in both non-bulged and bulged TiNi specimens, which clearly demonstrates the existence of Ti 2 p3/2 and Ti 2p1/2, respectively. The Ti 2p peak mainly indicates the availability of Ti as Ti4 þ in form of stable TiO2. The detection of Ti3 þ (Ti2O3) suggests the likely formation of sub-oxides within the oxide film [37]. Detection of O 1 s peaks at 529.8 eV and 531.0 eV generally reveals the availability of metal oxide (O2  ) species and hydroxyls (  OH) species, respectively, in the film. The peaks of Fe 2p3/2 (respectively around 705.0 eV, 711.2 eV and 720.9 eV) are fundamentally

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Fig. 10. The 3D morphology of a single bulge on Sb1 (a), Sb2 (b) and Sb3 (c) slid under contact pressure of 1.0 MPa and reciprocating speed of 48 mm/s.

Fig. 11. (a) Variation of friction coefficient of the non-bulged and bulged TiNi as a function of sliding time under 0.5 MPa and 48 mm/s (2 h), Worn surface of the pin before and after sliding against non-bulged (b) and bulged (c) specimens.

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suggesting the existence of metallic iron, Fe(II), and Fe(III) [38], which implies the transfer of Fe onto the TiNi specimens and/or the formation of Fe oxides in the film. The analysis of XPS spectra suggests that the oxide layer on the worn surface is mainly TiO2 and Fe2O3, which may be the main sources attributing to the resumption of a constant trend of friction coefficient curves without any obvious sign of fluctuation when sliding time is longer than 1000 s. Fig. 13 shows the schematic diagram anticipating the possible wear process. The wear process on the non-bulged specimen (Fig. 13a) involves with the gathering of wear debris along the wear track. Those not swept out debris tend to accumulate to the boundary of the sliding track or to be sandwiched to slide to and

fro continuously between interface. Such wear debris may exert high contact stress to deteriorate the friction and wear properties of the mating surfaces [39]. The schematic contact mode of 45# steel pin slid against the bulged TiNi circular plate is illustrated in Fig. 13b. Because of the effect of pseudo-elasticity, the applied contact pressure in this study is insufficient to cause any visible plastic deformation on the bulges and substrate at the initial friction test. Hence, the process of continuous sliding tends to flatten the top surface of bulge (Fig. 9b) so as to run in property with the pin. During sliding, oil film formed along the bulge side wall is beneficial to enhance hydrodynamic effect in reducing shear stress. The bulge profile and hydrodynamic effect act to flush the gathered wear particles off the sliding interface, which

Fig. 12. EDS analysis of the element peaks on worn scars: (a) non-bulged and (b) bulged Sb3 specimens under contact pressure of 0.1 MPa and reciprocating speed of 48 mm/ s, (c) XPS of worn surface of non-bulged and bulged (Sb3) specimens under contact pressure of 0.5 MPa and reciprocating speed of 48 mm/s.

Fig. 13. Schematic diagram of sliding wear process.

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subsequently reduce the abrasive effect as the role of dimple [5]. Increase in height of bulge tends to rise up the relative curvature radius on the bulge side wall and enlarge the projection area to admit lubricant flow; it thus decreases the contact area (i) between the pin and substrate and (ii) between the pin and bulge textures (Fig. 10). The reduction in real contact area thus decreases the friction coefficient of bulged TiNi specimens. In summary, integration of the deformation nature of bulges, the absence of wear debris between the contact interface, and the effective dissipation of friction heat significantly improves the friction and wear behaviors of TiNi alloy, especially that with suitably bulged texture.

4. Conclusions This study investigated the effect of bulge textures on the friction and wear properties of TiNi alloy under different contact pressures and reciprocating speeds. The study allows following conclusions to be drawn: 1) Processing sequence of indentation-polishing-heating allowed fabricating bulge textures on the surface of TiNi alloy specimens. The distinct difference of XRD analyses of non-bulged and bulged TiNi specimens illustrated their individually distinct patterns, which were identified mainly to be attributed to the different levels of (i) the thermally-induced transformation of martensite to austenite and (ii) the residual martensitic variant. 2) The bulged TiNi specimens with bulge heights of 24.3 μm, 31.0 μm and 34.1 μm, respectively, exhibited lower friction coefficient than their non-bulged counterpart by 13.8%, 40.5% and 44.8%, respectively. Lower friction coefficient for bulges with heights of 34.1 μm and 31.0 μm than that of counterpart with bulge height of 24.3 μm was basically attributed to the deformation nature of the bulge and the effect of their individual hydrodynamic flows. 3) Generally, the observation of decrease in the friction coefficient of bulged TiNi with reciprocating speed was mainly due to (i) the decrease in interfacial shear strength as a result of oil film formation to promote hydrodynamic flow and (ii) suppression of the temperature rise of interfacial surfaces. 4) Friction coefficient of the bulged TiNi rose as contact pressure increased from 0.4 to 1.0 MPa. Such rise was identified to be attributed to the reduction in the effect of pseudo-elasticity and to the increase in real contact area between interfacial surfaces. 5) Mild scratches produced by abrasive and/or some oxidation wear were the major wear mechanisms in the bulged TiNi specimens. However, the mechanism involved in the nonbulged specimen was basically dominated by (i) the forming of ploughing furrows, (ii) entrapment of fine wear particles between interfacial surfaces to initiate abrasive and delamination wear so as to generate irregular abrasion marks, and (iii) the occurrence of possible adhesive wear. The implementation of bulge textures significantly improves tribological properties of TiNi alloy.

Acknowledgment This study was financially supported by a matching Grant no.: 9220053, City University of Hong Kong, PR China. References [1] Shaw JA, Grummon DS, Foltz J. Superelastic NiTi honeycombs: fabrication and experiments. Smart Mater Struct 2007;16:S170–8.

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