Tribology of titanium boride-coated titanium balls against alumina ceramic: Wear, friction, and micromechanisms

Tribology of titanium boride-coated titanium balls against alumina ceramic: Wear, friction, and micromechanisms

Available online at Wear 265 (2008) 375–386 Tribology of titanium boride-coated titanium balls against alumina ceramic: Wear, ...

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Available online at

Wear 265 (2008) 375–386

Tribology of titanium boride-coated titanium balls against alumina ceramic: Wear, friction, and micromechanisms Curtis Lee a , Anthony Sanders b , Nishant Tikekar a , K.S. Ravi Chandran a,∗ a

Department of Metallurgical Engineering, The University of Utah, 135 South, 1460 East Rm. 412, Salt Lake City, UT 84112, United States b Ortho Development Corporation, Draper, UT 84020, United States Received 3 April 2007; received in revised form 22 October 2007; accepted 20 November 2007 Available online 16 January 2008

Abstract The tribological performance of titanium alloy (Ti–6Al–4V) balls coated with a dual boride layer comprised of titanium diboride (TiB2 ) and titanium boride (TiB) whiskers mated against alumina ceramic disks has been determined using lubricated ball-on-disk wear testing. Measurements of coefficient of friction values and volumetric wear were made and electron microscopic investigation of wear spots and tracks was performed. The wear rate of the boride-coated titanium alloy balls was 40 times less than that of 97% dense alumina balls. Measurements of wear track width and depth corroborated this result. The superior wear resistance is attributed to the hardness and the unique structure of the dual (TiB2 + TiB) whisker layer and the consequent smoothness of the wear surface created during the wear process. The material removal mechanism is abrasive in nature in the boride-coated balls compared to grain fracture and pullout in alumina. © 2007 Elsevier B.V. All rights reserved. Keywords: Borided Ti–6Al–4V; Tribology; Ball-on-disk wear testing; Alumina; Titanium boride coating; Titanium alloy

1. Introduction The higher specific strength and the superior corrosion resistance of titanium and its alloys, relative to other competing metals, make them desirable in many high performance applications. However, its tendency to gall and seize due to low work hardening behavior is a barrier to its use in applications where contact stresses and relative sliding between surfaces are high [1]. Several surface modification techniques, including chemical vapor deposition (CVD) [2], physical vapor deposition (PVD) [3], laser surface treatment [4], thermal oxidation [5], ion nitriding [6], and solid-state diffusion [7,8] have been developed to create hard and wear-resistant layers comprised of one or more hard compounds of titanium. While the common objective behind these techniques has been either the creation of hard surface layers of nitrides, carbides or borides or subsurface layers rich in N, C and O, there are some inherent disadvantages. For example, the thicknesses of layers that can be made by CVD and PVD techniques are limited thicknesses of 0.5–10 ␮m

Corresponding author. Tel.: +1 801 581 7197; fax: +1 801 581 4937. E-mail address: [email protected] (K.S.R. Chandran).

0043-1648/$ – see front matter © 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2007.11.011

[9], and result in coating only line-of-sight surfaces. The maximum hardness levels achievable by the impregnation of O and N atoms by thermal oxidation and ion nitriding are often limited to around 12 GPa (HV ) [10]. Methods based on laser surface treatment suffer from process complexity and cost, surface oxidation during melting, and the general restriction to planar surfaces [4]. Surface hardening of titanium by the solid-state diffusion of B is advantageous because the process itself is simple and inexpensive, does not need complicated equipment, and because complex geometries can be coated. In this process, B from a powder mixture surrounding the specimen is diffused into the titanium surface at high temperatures. The diffusion leads to a dual layer coating consisting of a continuous monolithic TiB2 outer layer and an inner layer consisting of discrete TiB whiskers, generally penetrating normal to the surface (Fig. 1(a)). The dual layer architecture provides a gradual transition from a hard ceramic outer layer (hardness of TiB2 layer is ∼30 GPa (HV ) [11]), through the next layer (hardness of TiB whisker layer is ∼20 GPa (HV ) [12]), to the ductile metal substrate (hardness of titanium is ∼4 GPa [13]). This dual layer structure holds tremendous potential for sustained wear resistance and low friction under contact conditions, for the layers are


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Fig. 1. Microstructure of (a) dual layer boride coating consisting of a TiB2 layer and TiB whisker layer, as observed in Ti–6Al–4V disk sample and SEM images of (b) alumina disk, and (c) alumina ball.

not discretely structured as in the case of CVD and PVD coatings. Additionally, the hardness levels are much higher than those achievable by thermal oxidation and ion nitriding. The objective of this study was to investigate the tribological characteristics of the dual layer boride coating against a hard ceramic. The wear resistance of this dual layer coating is expected to be favorable due to several factors: (i) TiB2 has a very high hardness and is a wear-resistant ceramic, (ii) TiB whisker tensile strength has been estimated to be as high as 5500 MPa [14] and (iii) the abrasive wear rate of Ti + 40% TiB2 sintered composite was found to be about 500× lower than the base Ti–6Al–4V alloy [15]. The present study employed balls and disks of alumina (Al2 O3 ) ceramic for comparative evaluation of the Ti–6Al–4V alloy balls coated with the dual boride layers. The structure of the TiB2 + TiB whisker dual layer coating, surface roughness outcomes of coated/polished balls, friction coefficient, wear performance, and the mechanisms of wear have been determined and are discussed in this work.

2. Materials and experiments 2.1. Materials and ball preparation Grade 200 Ti–6Al–4V alloy balls, Ø6.35 mm (Bal-Tec, Los Angeles, CA) were chosen for this study. Surface roughness measurements of these balls by contact profilometer yielded average values: Ra and Rz of 0.05 ␮m and 0.70 ␮m, respectively. Grade 25 alumina (Al2 O3 ) balls, Ø6.35 mm (Hoover, Sault Ste. Marie, MI) were used for comparison. Surface roughness levels of these balls were: Ra and Rz of 0.04 ␮m and 0.6 ␮m, respectively. For all the ball-on-disk wear tests, the counterface material was a Ø76.2 mm alumina disk (CoorsTek, Golden, CO). The disks were lapped to a flatness of 2.5 ␮m and Ra of 0.5 ␮m. The material property data for the alumina balls and disks supplied by the manufacturer are summarized in Table 1. SEM images of the unworn disk and ball surfaces appear in Fig. 1(b) and (c). The dual layer boride-coated balls were prepared by coating procedures discussed in previous work [7,8]. Briefly, this

Table 1 Property data for alumina balls and disks Article

Density (g/cm3 )



Flexural strength (MPa)

Compressive strength (MPa)

Al2 O3 ball Al2 O3 disk

3.87 3.92

99.5% 99.8%

1700 HV 83 RC @45 N

324 372

2,068 2,068

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Table 2 Diameter, sphericity, and roughness of Ti–6Al–4V balls before and after boriding and after polishing Ball condition

Diameter (mm)

Sphericity (␮m) average (S.D.)

Ra (␮m) average (S.D.)

As-received Borided Borided + polished (average) Borided + polished (best areas) Borided + polished (dull areas)

6.359144 6.374562 6.355537 – –

0.152 (0.051) 19.80 (11.9) 9.25 (5.21) – –

0.050 (0.0053) 0.555 (0.1960) 0.114 (0.0290) 0.024 (0.0089) 0.125 (0.0521)

involved surrounding the Ti–6Al–4V balls in a powder mixture composed of a B source, an activator, and a de-oxidant in a graphite crucible. The powder mixture was prepared by ball milling for 16 h. The balls were situated in the crucibles such that each ball was surrounded by at least 2.54 cm of powder. The boriding treatment was done at 1050 ◦ C for 24 h followed by slow cooling to room temperature. One treated ball was sectioned for metallographic examination. After boriding, the balls were polished by a commercial ball bearing manufacturer (ITI, Dexter, MI). Six balls were used in determining the average diameter, surface roughness, and sphericity prior to and after polishing; the results are given in Table 2. The polished surfaces had some roughness variation due to differing degrees of engagement with the ball polishing counterfaces, caused by the asphericity of the as-borided balls. Consequently, specific locations that provided the best finish and the thickest coating were identified for wear testing. Further, Vickers micro-indentation hardness measurements (LECO M-400, St. Joseph, MI, USA), made on balls secured in a collet, were supplemental in identifying the spots of coating suitable for wear testing. The hardness measurements revealed that the hardest coating locations corresponded to spots on the ball that had a relatively smooth surface texture. 2.2. Wear testing Wear testing was performed on a custom-made ball-on-disk wear-testing machine designed to meet the ASTM G99 specification. Basically, it consisted of a rotary stage in which the counterface alumina disk was secured, and above it, a collet into which the test ball was mounted. The rotary stage was designed to function inside a reservoir containing the lubricating liquid. The collet had a rear screw that was used to tightly grab and secure the ball along its equator, preventing ball movement during testing. The collet was attached to a force–torque sensor (ATI Automation, Apex, NC, USA; max. normal force: 200 ± 0.6 N, max. shear: 65 ± 0.2 N) for measurement of the normal and tangential forces on the ball during wear. The collet–sensor assembly was attached to the bottom of a loading stage that carried weights for normal loading. Wear tests were run under a load of 5.0 kg (49 N) and at a speed of 0.12 m s−1 . The disk and the ball were submerged in a mineral oil (PTI Process Chemicals, McHenry, IL) with an s.g. of 0.8395 at 25 ◦ C, and a viscosity of 106 cP at 40 ◦ C. The test track length was approximately 1.0 km. (Summary of test conditions provided in Table 3.) A National Instruments DAQPad-6020E data acquisition system was used to continuously monitor the

normal and tangential loads on the ball. The coefficient of friction (COF) was calculated and smoothed using a central moving average technique:  j+(n/2)−1  1/n i=j−(n/2) Fi (1) μj = j+(n/2)−1 1/n i=j−(n/2) Ni where μj is the average COF at a discreet distance (dj ), n is the number of data points in the average, and Fi and Ni are the measured horizontal and normal forces at a given distance (di ), respectively. A moving average is used to report the COF because the sensitivity of the sensor to both the normal and horizontal forces resulted in large scatter in the data. In part, the scatter was also exacerbated due to stick-slip friction between the ball and the disk, especially in the uncoated Ti–6Al–4V ball. Using a moving average reduced scatter and produced a relatively smooth line that could be analyzed and used to compare the average COF of different materials. 2.3. Determination of volumetric wear Volumetric wear on the balls was determined by two techniques. The first technique followed the ASTM G99-04 recommendations, assuming that the wear spot is an ideally flat feature. The wear spot was imaged using a Nikon Nexiv VMR 3020 optical coordinate measuring machine (CMM) at an appropriate magnification. The wear spot edge was identified at 10◦ increments, and a best-fit circle was fit to 36 edge points. The wear volume was then calculated using the following equations:   2 d 2 h=r− r − (2) 4   2  πh 3d 2 V = (3) +h 6 4 where h is the spot depth, r is the ball radius, d is the spot diameter, and V is the volume of material lost by wear. The second technique used a Zygo New View 5032 scanning white light interferometer (SWLI) (Zygo Corporation, Middlefield, CT), Table 3 Wear testing parameters Load Temperature Sliding distance Lubrication

5 kg Room temperature 1000 m Light mineral oil


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which used a direct method to measure volumetric wear. In this technique, the wear spot, including some adjacent regions, was scanned and used to create a 3D image (Fig. 2(a)). From the scanned data, the instrument calculated a best-fit sphere for the unworn regions and then subtracted that sphere from the entire data set. In the resulting data set, the unworn surface appeared flat and the worn surface appeared as a spherical crater (Fig. 2(b)). The instrument’s software then analytically determined the volume of the crater, which was the volumetric wear. There were small discrepancies between the results of the two wear measurement techniques due to problems in practice. In some instances, there was difficulty in measuring the alumina wear spots with the SWLI technique because the slope of the unworn surface around the spot deflected light away from the instrument’s objective. As a result, only partial data was obtained from the unworn surface. In the borided balls, two measurements resulted in significant discrepancies (47% and 29%), but three of the five measurements had discrepancies less than 3%. The average discrepancy for the alumina balls was 17%. In cases where there were discrepancies, CMM data was relied upon more than SWLI because the wear spots appeared flat and circular, and because there is precedence for using the wear spot diameter to calculate volumetric wear in alumina balls [16]. Also, the difficulty in obtaining data from the unworn regions of the alumina balls with SWLI makes the volume calculations using that method suspect for error.

Fig. 2. (a) Raw scan of a wear spot on a ball and (b) after subtraction of a best-fit sphere, as determined from the unworn regions in (a).

Fig. 3. (a) Microstructure of the coating on the ball after boriding treatment, and (b and c) after ball polishing. (b) Represents an area of maximal coating thickness on the ball and (c) corresponds to an area where minimal coating thickness remained.

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3. Results 3.1. Microstructure of dual layer boride coating After boriding, the Ti–6Al–4V balls were examined macroscopically and microscopically and no uncoated areas were found. Thus, the process was effective in uniformly treating the entire outer surface. The coating thickness of a sectioned ball was measured using an optical microscope. The monolithic outer TiB2 layer was measured as 5 ± 1 ␮m and the total coating thickness (consisting of both layers) was found to be 17 ± 9 ␮m. The coating appeared to be fully dense. Microstructures of the coating and the substrate are shown in Fig. 3(a)–(c). The boriding process resulted in a dual layer TiB2 + TiB whisker coating on the ball surface and a lamellar ␣ + ␤ microstructure in the interior. The interior microstructure is typical in heat treatment conditions involving slow cooling from above the ␤-transus temperature of the Ti–6Al–4V alloy. The outer layer of TiB2 and the substrate are separated by the intermediate TiB whisker layer; the TiB whiskers are visible as gray, while the TiB2 layer appears essentially white. The surface finish produced by ball polishing was not entirely uniform. Macroscopically, the surface appeared patchy with well-polished regions interspersed with relatively dull regions. Post-polish metallographic examination of cross-sectioned balls revealed regions which varied in coating thickness (Fig. 3(b) and (c)). The duller surface regions coincided with areas where a significant part of the coating was apparently removed during polishing (Fig. 3(c)). Such regions often appeared to lack the outer TiB2 layer; however, they typically retained a layer of TiB whiskers penetrating into the substrate. 3.2. Coefficient of friction The COF of metals and alloys are typically higher than hard materials like ceramics. Uncoated titanium and its alloys exhibit high friction even under lubricated conditions, due to the ductile nature and low work hardening behavior of titanium [13]. The COF in a test run on an uncoated Ti–6Al–4V ball ranged from 0.3 to 0.4 in the present test conditions as shown in Fig. 4(a). The COF values (taken as the RMS value for the entire test record) for the five tests of coated balls (along with the other tests) are

Fig. 4. (a) COF of uncoated Ti–6Al–4V ball against alumina and (b) COF of boride-coated Ti–6Al–4V and alumina balls, both against alumina. Tests were run using light mineral oil as a lubricant and under a 5 kg load. (Note the difference in scales between the two graphs.)

presented in Table 4. In Fig. 4(b), representative COF data from a test of a borided Ti–6Al–4V ball mated to an alumina disk is compared with COF data from a test of an alumina ball mated to an alumina disk. The boride-coated Ti–6Al–4V balls performed as well or slightly better than the alumina balls under the present test conditions. The RMS COF ranged between 0.11 and 0.12 for the tests run on borided Ti–6Al–4V balls and was 0.12 for the alumina balls. It is clear that the dual layer boride coating on titanium led to a decrease in the COF along with an increase in the hardness, relative to the uncoated titanium alloy. This is consistent with other surface-hardened titanium specimens where a significant

Table 4 Wear and COF for the materials tested Specimen name


70306 60929 60926 61004 61018 61020 61106 61103 60922

Ti–6Al–4V TiB2 + TiB TiB2 + TiB TiB2 + TiB TiB2 + TiB TiB2 + TiB Alumina Alumina Alumina


Distance (km)

Wear spot Ø (mm)

0.399a 1.021 1.027 1.021 1.021 1.012 1.012 1.012 1.212

4.56 0.478 0.716 0.738 0.556 0.722 1.552 1.644 1.729

Wear volume (mm3 )

Wear rate (mm3 /(N m))





8.35 0.0008 0.0041 0.0046 0.0015 0.0042 0.0914 0.1157 0.1417

– 0.0015 0.0041 0.0045 0.0021 0.0042 0.1103 0.0898 0.1332

4.27 × 10−4 1.60 × 10−8 8.14 × 10−8 9.19 × 10−8 3.00 × 10−8 8.46 × 10−8 1.84 × 10−6 2.33 × 10−6 2.38 × 10−6

– 3.00 × 10−8 8.14 × 10−8 8.99 × 10−8 4.19 × 10−8 8.46 × 10−8 2.22 × 10−6 1.81 × 10−6 2.24 × 10−6

Test was terminated after 400 m because nearly half of the ball was worn away.


0.35 0.11 0.11 0.11 0.10 0.11 0.12 0.12 0.12


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0.1095 (0.005) 0.120 (0.002) 6.55 × 10−8 (1.75 × 10−8 ) 2.09 × 10−6 (2.44 × 10−6 ) 6.08 × 10−8 (3.51 × 10−8 ) 2.19 × 10−6 (2.99 × 10−6 )

3.3. Volumetric wear of balls

Standard deviations appear in parentheses.

0.642 (0.118) 1.642 (0.089) TiB2 + TiB Alumina

1.020 (0.006) 1.078 (0.115)

16.7 (5.7) 108.2 (11.9)

0.0030 (0.0018) 0.1163 (0.0251)

0.0033 (0.0014) 0.1111 (0.0217)


COF (RMS) SWLI Wear rate (mm3 /(N m))


SWLI Wear volume (mm3 )

Height (␮m) Wear spot, Ø (mm) Distance (km) Material

Table 5 Multi-specimen averages of wear and COF

reduction in COF was seen [17,18]. For example, Itoh et al. [17] reported a COF of 0.05–0.2 for N+ implanted Ti–6Al–4V mated against itself, using a load of 460 mN and immersed in lubricating oil. Met et al. [18] reported a COF of 0.03–0.08 using a smooth fine-grained diamond coating on Ti–6Al–4V mated against itself under a 13 N load and lubricated with ringers solution.

Table 4 provides a compilation of volumetric wear and specific wear rates observed in individual balls. There is some variability in the wear of coated Ti–6Al–4V balls. This may be due to the variability in the coating thickness, and is probably not linked to the variations in surface roughness. The volumetric wear data of all of the TiB2 + TiB coated balls is at least an order of magnitude less than the alumina balls. The worst COF value for the coated balls is less than the best COF value for the alumina balls. Table 5 provides average wear and COF results. Optical micrographs of the smallest and largest borided ball wear spots are presented in Figs. 5 and 6, respectively. These demonstrate the nature of wear progression through the dual layer boride coating. Fig. 5(a) and (b) illustrates the general morphology of the wear spot and the surface topography of the ball. In Fig. 5(b), a dual contrast can be seen, indicating the remnant TiB2 layer (white areas in optical micrographs and the darker phase in the SEM images) and the TiB whisker layer (gray areas). This outcome can be imagined as the result of a planar cross-section through the ball, where the plane transects the coating while grazing the substrate. In Fig. 6(a) and (b), showing the specimen with the largest wear spot, the TiB2 layer is clearly visible, but the TiB layer is much less evident. Both are visible only along the edge (Fig. 6(b)). This is indicative of wear progression beyond the TiB whisker layer, such that the imagined intersecting plane now transects the substrate in large part. It is likely that close to the end of the test some of the load was actually born by the TiB2 layer (the elastic modulus of TiB2 is 560 GPa, compared to that of Ti–6Al–4V, which is 110 GPa), thus explaining how the ball was able to resist wear despite the inner regions of the wear spot seeming to consist almost entirely of the Tialloy matrix. It is also possible that these regions also contained some TiB whiskers that contributed to the wear resistance, in view of the presence of some long, sporadic TiB whiskers in the TiB layer of the coating (Fig. 3(a)). SEM images of a wear spot from a representative alumina ball are presented in Fig. 7(a)–(d). Unlike the worn surfaces of borided Ti–6Al–4V balls, the alumina balls exhibited grain fracture during the wear process. This is indicated by the relatively rough topography of the wear spot and the associated pits corresponding to the grain-pull-outs. 4. Discussion 4.1. Ball polishing, sphericity and uniformity of boride coating The overall surface roughness Ra of the borided balls after polishing, ∼0.10 ␮m, meets the specification for an AFBMA

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Fig. 5. Micrographs of the smallest wear spot of the borided Ti–6Al–4V (#060929) taken with optical microscope (a and b) and SEM (d and e). Note that the TiB2 phase appears white in optical micrographs and as dark phase in SEM micrographs. The TiB appears as light gray in optical micrographs and as dark gray in SEM micrographs.

Fig. 6. Micrographs of the largest wear spot of the borided Ti–6Al–4V (#061004) taken with optical microscope (a and b) and SEM (c). Note the changes in the contrast of the TiB2 /TiB phases upon going from optical to SEM micrography.


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Fig. 7. SEM images of a representative wear spot on an alumina ball (#061103) taken at different magnifications.

Grade 100 ball. Nonetheless, the surface texture was unexpectedly non-uniform, and this seems to have arisen from the increased asphericity of the balls after boriding. Sphericity is the radially measured envelope between the highest and lowest points on the sphere; thus, it measures spherical error. The average sphericity after boriding (∼20 ␮m) increased significantly from that before boriding (0.15 ␮m). Also, the sphericity was greater than the thickness of the monolithic portion of the coating. The ball polishing process works by progressively wearing the highest points more than the lowest, with the eventual result (ideally) of removing all spherical error. As such, some areas (high spots) of the surfaces were polished more (Fig. 3(c)), while other areas (low spots) were polished less (Fig. 3(b)). In the high spots, polishing apparently went through the TiB2 outer layer and ended within the TiB whisker zone. These areas macroscopically appeared duller due to polishing differences between the whiskers and the Ti-alloy substrate. In other areas, the polishing affected the outer TiB2 layer and finished leaving a highly reflective and polished surface. The Ra in such areas, selectively measured as 0.02 ␮m, was smoother than the overall Ra of the alumina balls (0.04 ␮m). The hardness results were consistent with the surface finish variation. Hardness indentations on wellpolished areas were smaller, indicating the presence of TiB2 and a sublayer of TiB whiskers. Larger indents were obtained in areas lacking the TiB2 layer. The spherical error resulting from the boriding process and its effects in polishing were causes of variation in surface roughness, hardness, and the coating thickness.

4.2. Wear The wear spot diameter, depth, and volume were substantially less with the borided balls (Table 4) relative to alumina balls. Although on average, the alumina balls were run for a slightly longer distance, the difference was not statistically significant. The wear difference between the borided Ti–6Al–4V balls and alumina balls, to some extent, appears to be due to the intrinsic differences in hardness. The hardness of the alumina ball was 17 GPa (HV ). The intrinsic hardness of TiB2 has been reported as 25 GPa (HV ) for a polycrystalline material [11]. A TiB2 + TiB coating has been reported to have a hardness of 32–44 GPa (KHN) [19]. PVD deposited TiB2 coatings have been reported to have hardness levels in the range of 52 ± 8 GPa [20]. TiB2 coatings made by fused salt electrolysis were reported to have hardness in the range 22–52 GPa (HV ) [21]. In the authors’ laboratory, the diffusion based TiB2 + TiB coatings, produced on flat Ti-alloy specimens, have typically exhibited hardness levels of 25–33 GPa (HV ), which is consistent with the range of hardness values reported by others. Further, flat specimens, mechanically polished through the TiB2 layer to the transition zone between TiB2 and TiB, have shown hardness values of 18–23 GPa (HV ). Fig. 8 illustrates the large variation in the sizes of Knoop hardness indentations from the outer TiB2 layer, through the TiB whisker layer, and into the substrate; there is about a factor of three difference between the indentation size in the TiB2 layer and that in the substrate. This should translate into about a factor of 10 difference in hardness values. From the cor-

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Fig. 8. Knoop microhardness indentations across the cross-section of the boride coating on Ti–6Al–4V alloy.

relation between the measurements of flat specimen hardness as well as metallographic examination of post-polish coating thickness on the borided balls, it is estimated that the hardness of the borided balls, in locations identified for testing, would have been in the range of 20–25 GPa (HV ). Consequently, on the basis of hardness alone, a significant reduction in wear could be anticipated for the borided balls compared to the alumina balls. In these tests, the specific wear rate of the TiB coated balls (6.55 × 10−8 mm3 /(N m)) was about two orders of magnitude less than that of the alumina balls (2.09 × 10−6 mm3 /(N m)), as shown in Table 5. Within limitations, the specific wear rate is useful for comparing these results against results reported in the literature. Saikko and Karenen examined the wear of a commercial medical implant alumina and reported a lower specific wear rate of 1.6 × 10−8 mm3 /(N m) [16]. This may be attributable to the higher density (3.98 g/cm3 ) and the significantly lower roughness of the disk (0.03–0.06 ␮m Ra ). Although there are few directly comparable wear tests of TiB and/or TiB2 , Pfohl et al. measured the wear of a ∼1 ␮m CVD TiB2 coating against alumina and reported a specific wear rate of 1.5 × 10−8 mm3 /(N m) [22]; however, their test conditions were not reported. For comparison of other coating systems on titanium, Met et al. reported wear results of a 2 ␮m thick and very smooth (15–35 nm Ra ) PVD diamond coating on Ti–6Al–4V [23]. The wear rate, in the range of 1.30 × 10−10 mm3 /(N m), is lower than the TiB2 results reported here. Clearly, further work to improve the surface finish of the boride coatings reported here are necessary to reduce the specific wear of the borided balls and to better compare the results against other reported systems. 4.3. Wear mechanisms Besides differences in hardness, differences in wear mechanisms affect wear rates [24]. The wear surface of the alumina balls appears to be relatively rough in comparison to the surrounding, unworn areas. The SEM images of the alumina wear spot clearly show that grains have been fractured and pulled out of the alumina ball (Fig. 7). Wear by grain fracture and grain removal has previously been reported in sliding wear of alumina


couples [25]. In contrast, the wear spots on the borided balls appear smooth, particularly in the boride layer regions. This is clearly observed on the smallest wear spot (Fig. 5). In this wear spot, the TiB2 layer is visible as the white material in the optical micrographs (it appears as the dark phase in the SEM images), located principally along the periphery. The beginning of the TiB zone is visible as the light gray material located principally in the interior of the spot. Darker gray material, representing the TiB whisker + Ti-alloy zone, is scattered throughout the interior. It is apparent that the TiB2 layer is worn through in the central area, while it is retained at the periphery, as anticipated in view of the ball’s curvature. Neither boride layer (TiB2 or TiB whisker layer) exhibited any sign of fracture or grain-fracturerelated wear mechanism. With the exception of some parallel wear striations, the boride surfaces appeared smooth. The fact that the TiB2 does not appear uniformly around the entire spot circumference may be attributable to asphericity of the ball. The largest wear spot on a borided ball exhibited a similar smoothness of the worn layers. Inward from the spot’s edge, the boride layers are identifiable, first as the white TiB2 , then by the light gray TiB (Fig. 6—colors are reversed for SEM image). The TiB2 layer is quite smooth, without any evidence of fracture or pull-out related wear processes. Further inward, within the darker gray interior of Fig. 6(b), there are many small white specs, which appear to be cross-sections of TiB whiskers. The increased wear rate of this specimen relative to the one having the smallest wear spot correlates well with the reduced proportion of boride material in the wear spot. It can be inferred that the variation in wear rates between the borided balls is linked to the varying degrees of exposure of the TiB whisker + Ti-alloy zone (e.g. Fig. 6) in different balls to the wear process. While the exposed inner region is undoubtedly softer than the boride annulus, the TiB whisker content should significantly reduce its wear rate compared to that of the untreated Ti-alloy. This is supported by previous observations of about 3× reductions in wear rate of a Ti-alloy + 20% TiB2 composite (the composite actually contained TiB particulates) compared to that of the Ti-alloy base [15]. Eventually, with increasing linear wear, the whisker layer would be surpassed, the area fraction of the boride layers would decrease, and the wear rate should become essentially close to that of a Ti-alloy ball. 4.4. Nature of wear tracks on disk The geometry and the depth of the wear tracks and the microstructure-scale damage on the alumina disk were dependent on the ball material. Fig. 9(a)–(d) illustrates the scanning electron microscopic analysis of a wear track made on an alumina disk by an alumina ball along with the details of Zygo profilometer measurements of wear track geometry. Fig. 10(a)–(d) illustrates similar data for a wear track made by a borided Ti–6Al–4V ball. The wear track created by the alumina ball is deeper and wider (Fig. 9(a) and (b)) compared to that made by a borided


C. Lee et al. / Wear 265 (2008) 375–386

Fig. 9. The nature of damage in the wear track of alumina disk created by alumina ball: (a) track width, (b) Zygo profilometric topography, (c) track depth profile at one location, and (d) SEM micrograph of worn region inside the track.

Fig. 10. The nature of damage in the wear track of alumina disk created by borided Ti–6Al–4V ball: (a) track width, (b) Zygo profilometric topography, (c) depth profile at one location, and (d) SEM micrograph of worn region inside the track.

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ball (Fig. 10(a) and (b)). The alumina ball wore into the alumina disk creating a wear track of about 1.13 mm wide and 1.75 ␮m deep (Fig. 9(c)). In contrast, even though the track made by the borided ball could be seen macroscopically as a faint, gray line, the profilometry data showed that the borided ball barely scratched the surface of the disk (Fig. 10(c)). Further, the microscopic damage from the alumina ball on the alumina disk is much worse than the damage created by the borided ball. In Fig. 9(d), the alumina appears to be damaged with the grain fracture features resembling those of a tensile fracture surface. In comparison, the microstructure of the wear track made by the borided ball (Fig. 10(d)) is relatively smooth and appeared similar to that of the unworn surface (Fig. 1(b)). It is therefore clear that the relatively reduced wear in the alumina disk + borided Ti–6Al–4V ball combination is attributable to an abrasive wear mechanism that produces a smooth surface both on the borided ball and on the disk. In contrast, the alumina disk + alumina ball combination seemed to result in a grain-fracture-and-removal wear mechanism, ultimately producing a much greater specific wear rate. 5. Conclusions 1. The volumetric wear rate of titanium alloy balls coated with a dual layer boride (TiB2 + TiB whisker) mated against alumina was 40 times less than that of alumina balls mated against alumina. This was found to be the case even in instances where a substantial amount of substrate Ti-alloy was exposed during the wear process. 2. The width and depth of the wear track on the alumina disk by the alumina ball was much larger than that made by the borided ball. The borided ball barely scratched the disk surface and showed much superior wear performance. 3. The boride layers wore by a different mechanism than the alumina. The alumina wear surfaces exhibited grainfracture-related wear mechanisms, producing a rough surface, whereas the boride material wore by an abrasive process, resulting in a relatively smooth wear surface. 4. The COF of the boride-coated Ti–6Al–4V balls was smaller than that of alumina balls by a narrow margin. This was observed in spite of the fact that the overall initial surface roughness of the alumina balls was better. 5. The boriding process produced surface roughening and geometric form changes that influenced the sphericity of the balls. The sphericity changes, although slight, affected the polishing process and resulted in a variable surface texture, hardness, coating thickness, and ultimately, wear. Acknowledgements The authors gratefully acknowledge the State of Utah Office of Economic Development and Ortho Development for financial support of this work. We are indebted to Dave Edwards


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