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Investigation into wear of Ti–6Al–4V under reciprocating sliding conditions
Wear 267 (2009) 368–373
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Wear journal homepage: www.elsevier.com/locate/wear
Investigation into wear of Ti–6Al–4V under reciprocating sliding conditions R.S. Magaziner a , V.K. Jain a , S. Mall b,∗ a b
University of Dayton, Dayton, OH 45469-0238, USA Air Force Institute of Technology, Wright-Patterson Air Force Base, OH 45433-7817, USA
a r t i c l e
i n f o
Article history: Received 27 August 2008 Received in revised form 8 December 2008 Accepted 9 December 2008 Available online 24 March 2009 Keywords: Wear Fretting Fatigue Titanium alloy Relative slip
a b s t r a c t Wear behaviour of titanium alloy, Ti–6Al–4V under fretting-reciprocating condition was characterized under different loading conditions, contact surface conditions and contact configurations. Six series of tests were analyzed to investigate the effects of relative slip, contact load, number of cycles, bulk cyclic stress, contact geometry, and a lubricant. Dry and lubricated surface conditions produced “W” (scar with multiple valleys) and “U” (scar with single valley) shaped scars on both specimen and pad, respectively. WV (wear volume) was linearly related to the cumulative product of contact load and relative slip and to the total dissipated energy under different loading and sliding conditions for the cylinder-on-flat as well as flat-on-flat contacts for dry and lubricated conditions. The microscopy of wear scars showed that the wear was caused by both the adhesion and abrasion processes. Published by Elsevier B.V.
1. Introduction Wear occurs due to the relative displacement between two bodies mating under clamping force, and with or without cyclic load on either body. It is referred to as the fretting wear in the presence of cyclic load. Further, the surface damage is produced due to fretting action, which in turn results in premature initiation and subsequent growth of fatigue cracks, leading to shorter fatigue life of a component or failure at the stress well below the fatigue strength of the material. Fretting induced surface damage is maximum under the partial slip condition while the maximum wear damage is found to occur under the gross slip condition [1]. Therefore, there is practical need to predict/estimate the geometry of fretting and reciprocating wear behaviour to design mating mechanical components reliably and efficiently since it could result in undesirable vibration as well as misalignment. Several studies have been conducted to characterize the wear behaviour of various materials under different fretting and sliding conditions [1–18]. Several fretting maps concepts have been proposed to identify the wear or fretting wear behavior where test conditions and damage evolutions are interrelated. Vingsbo and Soederberg introduced this approach by correlating the displacement and normal contact force to identify the partial slip, gross slip or mixed fretting regimes, which was referred to as the running condition fretting map [1]. Later, Fouvry and co-workers expanded this approach by including several other tests as well as material
∗ Corresponding author. Tel.: +1 937 255 3636; fax: +1 937 656 7053. E-mail address: Shankar.Mall@afit.edu (S. Mall). 0043-1648/$ – see front matter. Published by Elsevier B.V. doi:10.1016/j.wear.2008.12.083
response parameters [2–6]. They also conducted a series of wear and fretting wear studies in a detailed way to develop fretting maps as well as to investigate the wear extensively without any applied load on the substrate [7–13]. The fretting wear behavior of SC 65-2 uncoated and TiN coated steel against an alumina ball was investigated by this group of researchers [7–9]. They used gross sliding condition. The WV (wear volume) was obtained from the wear scars produced on the flat surface without any applied load. It was noted that the maximum wear occurred at the contact borders and minimum at the center. This led to a typical W-shaped scar, which changed with number of cycles or slip amplitude, and the wear was found to have the characteristic U-shape for large sliding amplitudes. Here, the wear volume was measured using three-dimensional as well as two-dimensional topographies of wear scars. The modeling of wear indicated that calculation of wear using a two-dimensional topography is much less time consuming and sufficiently accurate for modeling work. They confirmed that wear volume could be correlated with the cumulative dissipated energy with a linear relationship. The TiN coating was found to be about nine times wear resistant as the plain HSS substrate. Working with coated and uncoated HSS steel against an alumina ball, Fourvy et al. [10] noted that Archard’s wear model does not work well since friction coefficient is not constant. They suggested that use of interfacial shear work is more appropriate. The wear volume was linearly related to the interfacial shear work. The wear energy coefficient determined by this method accurately classified the wear resistance of various tribosystems. Fouvry et al. [11] also studied the fretting behavior of Ti–6Al–4V for a plain cylinder/plain contact configurations under ambient laboratory condition again without any applied load on the substrate.
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Fretting wear response of the system was studied under gross slip condition. They noted that to correctly quantify the wear behavior, the total wear volume of both the sliding bodies must be considered. They developed a fretting wear parameter in terms of pressure, sliding amplitude, test duration, and contact dimensions to quantify the wear kinetics. Fouvry et al. [12] also studied the fretting wear behavior of Ti–V–C and Ti–V–N unloaded systems sliding against an alumina ball. It was observed that for the Ti–V–C system, the wear debris was ejected from the interface; the wear mechanism is driven by a two-body abrasion displaying conventional U-shaped wear scar morphology. The wear rate decreased with hardness. For the Ti–C–N system, the three body abrasion process was modified. Oxide debris agglomerated on the alumina counter body, shifting the interface to a three body abrasion promoting typical W-shaped scar morphology. The wear rate dependence with coating hardness could not be verified for this case. Fridrici et al. studied the fretting wear of shot-peened and un-peened Ti–6Al–4V also without any applied load on the substrate [13]. They used cylinder-on-flat configuration and varied both the relative slip amplitude and contact load in their tests. They found that shot-peening has no effect on fretting wear behavior of Ti–6Al–4V at a given relative slip amplitude and contact load using the dissipated energy method. However, the energy wear factor was found to be dependent on relative slip amplitude. Huq and Celis analyzed the wear of several hard coatings using the dissipated energy [14]. They concluded that energy dissipation is a useful tool to quantify the wear damage, and its empirical parameters are different for unidirectional and reciprocating sliding conditions. McColl et al. [15] analyzed the fretting wear using finite element analysis based on Archard’s equation for a cylinder-on flatcontact. They determined surface and sub-surface stresses in the contact zone. It was shown that for gross slip condition, the high wear rate leads to contact edges to move rapidly outwards, leaving the material previously at the contact edges in a permanently compressive state, which retards the fretting-fatigue crack initiation. Thus, they suggested that wear becomes the predominant feature of the gross slip condition in the absence of fretting-fatigue. Jin and Mall designed and developed a fretting test set up which was capable of applying independent pad movement under a normal force and stress amplitude condition [16]. They investigated the fretting wear behavior of Ti–6Al–4V under gross slip condition for cylinder-on-flat contact configuration, but they did not measure the amount of wear [16,17]. The fretting-fatigue life was found to be larger for the gross slip condition than that for the partial slip. The morphology of scar profile was also studied, which was found to be W-shaped. The large wear at the edge of contact was attributed to abrasion caused by the detached and fragmented wear particles. Lee et al. investigated the fretting behavior of soft Cu–Al coating on Ti–6Al–4V substrate with and without fatigue load [18]. Energy approach of wear analysis showed a linear relationship between wear volume and accumulated dissipated energy. This relationship was independent of fatigue loading condition and extended from partial slip to gross slip regimes. The above review of the previous studies clearly shows that although a large amount of research has been conducted in the area of fretting wear, but none of these previous studies were conducted where substrate was subjected to an cyclic axial bulk load except by Lee et al. on coating. This is where the present work differs from the previous investigations [1–17]. In the present study, a cyclic axial load was applied on the flat specimen to investigate the fretting wear behavior of Ti–6Al–4V unlike the previous studies where substrate was without any applied load. The present authors have also investigated the frettingreciprocating wear behaviour of a titanium alloy in contact with the same titanium alloy (i.e. both Ti–6Al–4V) [19], since this material is widely used in aerospace and biomedical industry due to
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its superior mechanical properties and relatively low density. For this purpose, four series of tests were conducted to characterize the effects of different sliding and loading parameters on the wear behaviour under cylinder-on-flat contact configuration; and these parameters were relative slip, normal contact load, number of cycles, and bulk cyclic stress. The present study is an extension of this previous study [19] where wear behaviour of same material (Ti–6Al–4V) was characterized with new contact geometry (flaton-flat) and under the lubricated surface condition. These new data along with data from the different conditions from a previous study were combined and analyzed to characterize the wear behaviour of Ti–6Al–4V. As mentioned earlier, it should be reiterated that the focus of this study is fretting-reciprocating sliding condition unlike the conventional plain reciprocating sliding condition in the previous studies [1–17]. 2. Experiments 2.1. Test set up A test machine was used that could apply alternating load on the specimen and reciprocating motion to the pads independent of each other. It was a modified servo-controlled MTS fatigue test machine with two actuators. The details of this test machine are provided in [16–19]. A lubricating system was also set up to get lubricant between the specimen and pads. Lubricant was brought from a reservoir, which was placed above the test frame. Small capillary tubes were used so that lubricant could be dropped in the region where the pads contacted the specimen and ensured that lubricant entered into the contact area such that both the pads and specimen were lubricated well. SAE 15W-40 Heavy Duty Motor Oil was used in this study as the lubricant. Though it is not exactly the same lubricant as is used in the gas turbine engines, its lubricating properties are very similar. This arrangement simulated well the practical conditions of dovetail joint in the gas turbine engines. Two extensometers were mounted on the fretting fixture. One extensometer was attached to the specimen and the pad holder to measure their relative displacement. The second one was attached to the pad holder and a fixed point, which measured the absolute displacement of the former. These extensometer measurements were then used to calculate the structural compliance of the test rig, which was then used to compute the relative slip (ı) between the specimen and pad from the applied displacement, ıapp from the test machine to the pad [16–21]. This provided the relative slip, ı by incorporating the effects of both the structural compliance and specimen as elaborated in previous studies [16–21]. The normal contact load (P) on the pad as the function of applied cycles was continuously monitored by the load cell in each test. The P decreased in the tests with increasing number of cycles, especially in the longer running tests (i.e. with number of cycles). This decrease in P was caused by the wear of the pad and specimen [19]. This decrease in normal contact load was duly accounted for in the analysis of all test results. The tangential force produced at the contact was monitored from the loads measured at the both sides of the specimen. The measured tangential force (Q) for each of the two fretting contacts (one on each side of the specimen) was one-half of the difference of these loads. 2.2. Material Titanium alloy, Ti–6Al–4V was the test material because this material is widely used in aerospace applications, experiencing fretting wear damage [2,6,17–19]. The dog-bone shaped specimens having a width of 6.35 mm and thickness of 3.81 mm were used (Fig. 1a). The cylindrical fretting pads had a radius of 50.8 mm
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men. In test series #5, all test conditions were same as that “baseline series (#1) and also a lubricant was used in this case which was 15W-40 oil to examine the effect of lubrication on the wear mechanism, development of scar shape and the wear volume. In test series #6, a flat pad was used with loading conditions similar to “baseline” series except that ıapp = 1 mm. Several tests were conducted in each test series where specimens were subjected to different number of cycles in laboratory air at ambient room temperature at a frequency of 2 Hz. 3. Results and discussion 3.1. Wear scar details Fig. 1. Schematic of: (a) specimen, (b) cylindrical, and (c) flat fretting pad.
(Fig. 1b) and the flat had an edge radius of 2.54 mm with a central flat length of 4.45 mm (Fig. 1c). The material had an elastic modulus of 126 GPa, yield strength 930 MPa, ultimate tensile strength of 1 GPa and Brinell hardness of 302. 2.3. Wear Taylor Hobson “Ultra” profilometer was used to measure the wear volume. The profiles were taken over the scar surface in the direction of the fretting-reciprocating motion at 250 m intervals, and the above process was repeated until the complete scar was scanned. All the profiles from a surface were then assembled using MATLAB software developed to generate a complete topography of the wear scar. Wear volumes were computed using the trapezoid rule of numerical integration process. Each test in this study produced four wear scars: one on each side of the specimen and two on the corresponding fretting pads. The combined pad and specimen wear volumes were considered as the wear volume as opposed to the specimen or pad wear volume only [2,6]. It is because the material transfer between the pad and specimen can distort the perception of the actual wear damage if only either the pad or specimen wear volume is used, the reported wear volume was calculated by dividing the sum of four wear volumes (two pads and two sides of the flat specimen) by two. Finally, the wear scars on the flat specimen were examined by optical and scanning electron microscopes to elucidate the wear mechanisms operating in fretting-reciprocating sliding conditions. The specimens were ultrasonically cleaned to ensure that no surface features are hidden by the wear debris.
Fig. 2a shows the typical W-shaped wear scar on the specimen from a test in the “baseline” series after an application of 22,000 fretting cycles. Fig. 2b shows a scar with one broad valley in the middle. This scar is from the “lub” series test also after again 22,000 fretting cycles. This type of scar is termed as U-shaped scar. Thus, the fretting tests yielded in two types of wear scars: scars with multiple peaks and valleys (W-shaped) and scars with one wide valley (Ushaped). The evolution of two-dimensional view of the scar profiles of “baseline”, “flat”, and “lub” series tests are compared in Fig. 3. These profiles are in the longitudinal direction (i.e. sliding direction) at the middle of the specimen which provides the development and progression of wear scar with increasing number of fretting cycles. As the pad wore down deeper into the specimen surface, the
2.4. Test plan A total of six test series were considered; four from a previous study [19] and two from the present study to investigate the effects of relative slip amplitude, normal contact load, contact geometry, lubrication, loading cycles and bulk axial stress on fretting wear behaviour of Ti–6Al–4V. These were (1) the “baseline” series, (2) “baseline” series with ıapp = 1 mm (will be referred to as “low-ı” series), (3) “baseline” series with P = 4003 N (will be referred to as “high-P” series), (4) “baseline” series with = 0 (will be referred to as “no-” series), (5) “lubricated” series with 15W-40 lubricant (will be referred to as “lub” series), and (6) “flat pad” series with loading conditions similar to “baseline” series except that pad was flat (unlike the cylindrical in previous five series of tests) and ıapp = 1 mm (will be referred to as “flat” series). Test series #1 to #5 used the cylindrical pads. In the “baseline” series test (#1), the cylindrical pad was loaded against the flat specimen with a load, P of 1334 N, and subjected to an applied slip (ıapp ) of 1.5 mm, and an alternating bulk stress = 17–570 MPa was applied on the flat speci-
Fig. 2. Three-dimensional profile of (a) “baseline” series and (b) “lub” series; 22,000cycle specimen scar, ıapp = 1.5 mm; initial P = 1334 N, = 17–570 MPa.
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the figure and flaking of the contact surface in the upper portion. The sliding is in the vertical direction. Severe plastic deformation, which was noted for all cylindrical pads as shown in Fig. 4c, was absent in this case. For the sake of comparison, Fig. 4c shows a typical SEM micrograph of the middle region of scar from test series #1 to 4. This difference may be due to the reason that the pressure and shear stress at the surface were much lower in the flat case. However, a number of wear particles can be seen lying over the surface, Fig. 4a. It is these particles which are responsible for the abrasion grooves. Thus, wear mechanism was the adhesion, breakage of adhesive bonds, which form wear particles, and finally the abrasion caused by the wear particles. Fig. 4b is another micrograph obtained from the flat pad at the edge of the scar. This figure shows light deformation with material movement in the direction of the sliding. The deformation is mainly occurring at the asperity peaks only. Few wear particles can be seen over the surface but they were unable to form deep grooves. This may be due to the rounded edge of the pad, which did not produce sufficient pressure and also because the particles were not sufficiently strain hardened to penetrate the material. Fig. 4d shows a SEM micrograph obtained from the middle of the lubricated scar. This scar surface is smooth indicating the absence of severe plastic deformation and adhesion, which was expected for a lubricated contact. The surface is covered with circular-to-elliptical regions with dimensions of 40 m × 60 m. These are the material grains with very faint sliding marks. Thus, the wear mechanism for the lubricated contact was plastic deformation of high asperities followed by their removal due to polishing action. 3.2. Wear volume
Fig. 3. Evolution of wear scar with cycling for (a) “baseline” series, (b) “flat” series and (c) “lub” series; ıapp = 1.5 mm for (a) and (c), ıapp = 1.0 mm for (b), initial P = 1334 N, = 17–570 MPa.
contact area increased initially and then remained constant, which occurred around 22,000th cycle for all the series tests. Note that the “baseline” series scar has three valleys and two peaks; “flat” series profiles have one peak and two valleys; whereas the “lub” series has only one wide valley. For rest of the series tests, the scar shape was similar to that of the “baseline” series. The progression of scar lengths shows initial high rate of wear due to peak contact pressure, which decreases and the shear stresses are redistributed over the wider contact surfaces [15,22]. Therefore, the wear rate started decreasing with increasing number of cycles. Fig. 4a and b shows SEM micrographs obtained from the middle and edge of the wear scar produced on the specimen by the flat pad with rounded edges, respectively after 60,000 fretting cycles. Fig. 4a shows a number of abrasion grooves in the lower portion of
3.2.1. Wear versus fretting cycles Relationships between wear volumes and number of cycles (N) for all six series of tests are shown in Fig. 5, where each data corresponds to an independent test. As expected the wear volume increased with numbers of cycles. The wear volume is clearly dependent on the applied loading condition, i.e. P and ı, and the interface condition. When compared to the “baseline” series tests, the increased contact load in the “high-P” series tests increased the wear volume for the same number of cycles, i.e. it had the maximum wear volume out of the six test series. Decreasing the applied relative slip in the “low-ı” series had the opposite effect, i.e. it had the minimum wear volume. The effect of no-bulk stress on the specimen (“no-” test series) resulted in tests with roughly 200 m shorter relative slip than the “baseline” series tests. Differences in the effective relative slip explain the reason for the “no-” series wear volume curve being between the “baseline” and “lowı” series curves. The wear volume for the flat pad is smaller than the “baseline” series but larger than the “no-” series. This may be attributed to the lower contact pressure between the flat pad and the fretting specimen and lower ıappl (=1 mm). For the lubricated condition, the wear volume was expected to be lowest of all the cases. However, it was higher than the “low-ı” series. This may be attributed to several reasons: (a) the larger slip between the pad and the specimen due to lower coefficient of friction, (b) rapid rate of wear particle removal from the contact zone, and (c) corrosion of the contact zone by the extreme pressure additives in the lubricant, because 15W-40 oil is not formulated for a titanium alloy. Furthermore, the slope of each curve decreases with the increasing number of cycles, indicating a decreasing wear rate. This reduction in wear rate can be attributed to the decrease in normal load with number of cycles and the contact stresses due to increased area of contact. 3.2.2. Cumulative product of normal contact load and relative slip In this approach, the sum of the product of normal contact load and relative slip per cycle, (Pın ) were correlated with the wear
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Fig. 4. SEM micrograph of scar (a) at the centre of a “flat” series, (b) at the edge of a “flat” series, (c) at centre of “no-” series and (d) at centre of “lub” series.
volume as shown in Fig. 6. This shows a single linear correlation between wear volume and cumulative Pın for all the six series of tests of this study. Thus, it appears that the linear relationship between the wear volume and product of normal load and relative slip data is independent of the test parameters, i.e. contact load, relative slip, applied cyclic load on the substrate (specimen), contact geometry, contact load and surface condition. The slope of linear relationship between wear volume and cumulative product of contact load and relative slip ( Pın ) was 3.709 × 10−4 mm3 /(N m) with an R2 value of 0.982. The corresponding value for this material for unidirectional sliding is reported as 3.18 × 10−4 mm3 /(N m) [23]. Thus, the value of slope of this study is comparable to its counterpart as obtained from the classical unidirectional sliding method. The difference between these two values may be attributed to the
Fig. 5. Wear volume versus number of fretting cycles relationships.
several factors, such as sliding condition, test apparatus, material heat-treatment and microstructure, and the operating wear mechanisms. 3.2.3. Dissipated energy This is an alternate approach which could be used to characterize the wear behaviour. In this approach, wear volume is assumed to be proportional to dissipated energy ( Ed ) and is given by: WV = aV
Ed
(1)
where ˛V is a constant of proportionality and is called the energy wear factor. Here, the dissipated energy, Ed for a single cycle is the
Fig. 6. Wear volume versus sum of product of contact load and relative slip relationship.
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cyclic load on the substrate (specimen). For these different fretting test parameters, the proportionality constant of linear relationship between wear volume and cumulative product of contact load and relative slip (Pın ) was 3.709 × 10−4 mm3 /(N m), and for the linear relationship between wear volume and dissipated energy was 7.019 × 10−4 mm3 /(N m) in the case of the Ti–6Al–4V. However, further tests are needed to confirm and generalize for other fretting conditions and materials. References
Fig. 7. Wear volume versus total dissipated energy relationship.
area within a hysteresis loop between tangential force and relative Ed is deterslip (Q–ı plot). The accumulated dissipated energy, mined by summing the areas of all loops of a test [18,19]. This study used the actual area within the fretting loops. For this purpose, a MATLAB based program was written to numerically integrate the Q–ı loops. The dissipated energy values were then summed together to determine the total energy dissipated for a specific num ber of cycles ( Ed ). Fig. 7 shows the variation of wear volume as a function of the total dissipation energy. The plot shows that wear volume is linearly proportional to total dissipated energy for all six test series. The least squares fit to the data resulted in a slope of the line, ˛V = 7.019 × 10−4 mm3 /J with an R2 value of 0.964. It may be noted that the correlation between the wear volume and dissipated energy data is excellent. Further, it is interesting to note that the linear relationship between the wear volume and dissipated energy data is also independent of the test parameters, i.e. contact load, relative slip, contact geometry, lubrication, and applied cyclic load on the substrate (specimen) as in the case of Archard’s approach. 4. Conclusions The wear behaviour of Ti–6Al–4V subjected to frettingreciprocating sliding condition was investigated to analyze the effects of contact load, relative slip, contact geometry, lubrication, and applied cyclic load on the substrate (specimen). Dry and lubricated surface conditions produced “W” (scar with multiple valleys) and “U” (scar with single valley) shaped scars on both specimen and pad, respectively. The wear volume was linearly proportional to cumulative product of contact load and relative slip (Pın ) as well as to the total dissipated energy (Ed ). These correlations suggest that linear relationships exist to characterize the fretting wear of a material which are independent of the contact load, relative slip, contact geometry, lubrication, and applied
[1] O. Vingsbo, S. Soederberg, On fretting maps, Wear 126 (1988) 131–147. [2] S. Fouvry, P. Duo, P. Perruchaut, A quantitative approach to Ti–6Al–4V fretting damage: friction, wear and crack nucleation, Wear 257 (2004) 916–929. [3] T. Liskiewicz, S. Fouvry, Development of a friction energy capacity approach to predict the surface coating endurance under complex oscillating sliding conditions, Tribol. Int. 38 (2005) 69–79. [4] T. Liskiewicz, S. Fouvry, B. Wendler, Impact of variable loading conditions on fretting wear, Surf. Coat. Technol. 163–164 (2003) 465–471. [5] S. Fouvry, P. Kapsa, L. Vincent, in: Y. Mutoh, S. Kinyon, D. Hoeppner (Eds.), A global Methodology to Quantify Fretting Damage. Fretting-Fatigue: Advances in Basic Understanding and Applications, ASTM STP 1425, ASTM International, West Conshohocken, PA, 2003, pp. 17–31. [6] C. Paulin, S. Fouvry, S. Deyber, Wear kinetics of Ti–6Al–4V under constant and variable fretting sliding conditions, Wear 259 (2005) 292–299. [7] S. Fouvry, P. Kapsa, H. Zahouani, L. Vincent, Wear analysis in fretting of hard coatings through a dissipated energy concept, Wear 203–204 (1997) 393–403. [8] S. Fouvry, P. Kapsa, An energy description of hard coating wear mechanisms, Surf. Coat. Technol. 138 (2001) 141–148. [9] S. Fouvry, P. Kapsa, L. Vincent, Quantification of fretting damage, Wear 200 (1996) 186–205. [10] S. Fouvry, T. Liskiewicz, P. Kapsa, S. Hannel, E. Sauger, An energy description of wear mechanisms and its applications to oscillating sliding contacts, Wear 255 (2003) 287–298. [11] S. Fouvry, P. Duo, P. Perruchaut, A quantitative approach of Ti–6Al–4V fretting damage: friction, wear and crack nucleation, Wear 257 (2004) 916–929. [12] S. Fouvry, T. Liskiewicz, C. Paulin, Global-local wear approach to quantify the contact endurance under reciprocating-fretting sliding conditions, Wear 263 (2007) 518–531. [13] V. Fridrici, S. Fouvry, P. Kapsa, Effect of shot peening on the fretting wear of Ti–6Al–4V, Wear 250 (2001) 642–649. [14] M.Z. Huq, J.-P. Celis, Expressing wear rate in sliding contacts based on dissipated energy, Wear 252 (2002) 375–383. [15] I.R. McColl, J. Ding, S.B. Leen, Finite element simulation and experimental validation of fretting wear, Wear 256 (2004) 1114–1127. [16] O. Jin, Mall.F S., Effects of independent pad displacement on fretting-fatigue behavior of Ti–6Al–4V, Wear 253 (2002) 585–596. [17] O. Jin, S. Mall, Influence of contact configuration on fretting-fatigue behavior of Ti–6Al–4V under independent pad displacement condition, Int. J. Fatigue 24 (2002) 1243–1253. [18] H. Lee, S. Mall, J. Sanders, S. Sharma, Wear analysis of Cu–Al Coating on Ti–6Al–4V substrate under fretting condition, Tribol. Lett. 19 (2005) 239–248. [19] R.S. Magaziner, V.K. Jain, S. Mall, Wear characterization of Ti–6Al–4V under fretting-reciprocating sliding conditions, Wear 264 (2008) 1002–1014. [20] B.U. Wittkowsky, P.R. Birch, J. Dominguez, S. Suresh, An apparatus for quantitative fretting-fatigue testing, Fatigue Fract. Eng. Mater. Struct. 22 (1999) 307–320. [21] R. Magaziner, O. Jin, S. Mall, Slip regime explanation of observed size effects in fretting, Wear 257 (2004) 190–197. [22] O. Jin, S. Mall, Effects of slip on fretting behavior: experiments and analyses, Wear 256 (2003) 671–684. [23] M.B. Peterson, W.O. Winer (Eds.), Wear Control Handbook, The American Society of Mechanical Engineers, USA, 1980.
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Wear journal homepage: www.elsevier.com/locate/wear
Investigation into wear of Ti–6Al–4V under reciprocating sliding conditions R.S. Magaziner a , V.K. Jain a , S. Mall b,∗ a b
University of Dayton, Dayton, OH 45469-0238, USA Air Force Institute of Technology, Wright-Patterson Air Force Base, OH 45433-7817, USA
a r t i c l e
i n f o
Article history: Received 27 August 2008 Received in revised form 8 December 2008 Accepted 9 December 2008 Available online 24 March 2009 Keywords: Wear Fretting Fatigue Titanium alloy Relative slip
a b s t r a c t Wear behaviour of titanium alloy, Ti–6Al–4V under fretting-reciprocating condition was characterized under different loading conditions, contact surface conditions and contact configurations. Six series of tests were analyzed to investigate the effects of relative slip, contact load, number of cycles, bulk cyclic stress, contact geometry, and a lubricant. Dry and lubricated surface conditions produced “W” (scar with multiple valleys) and “U” (scar with single valley) shaped scars on both specimen and pad, respectively. WV (wear volume) was linearly related to the cumulative product of contact load and relative slip and to the total dissipated energy under different loading and sliding conditions for the cylinder-on-flat as well as flat-on-flat contacts for dry and lubricated conditions. The microscopy of wear scars showed that the wear was caused by both the adhesion and abrasion processes. Published by Elsevier B.V.
1. Introduction Wear occurs due to the relative displacement between two bodies mating under clamping force, and with or without cyclic load on either body. It is referred to as the fretting wear in the presence of cyclic load. Further, the surface damage is produced due to fretting action, which in turn results in premature initiation and subsequent growth of fatigue cracks, leading to shorter fatigue life of a component or failure at the stress well below the fatigue strength of the material. Fretting induced surface damage is maximum under the partial slip condition while the maximum wear damage is found to occur under the gross slip condition [1]. Therefore, there is practical need to predict/estimate the geometry of fretting and reciprocating wear behaviour to design mating mechanical components reliably and efficiently since it could result in undesirable vibration as well as misalignment. Several studies have been conducted to characterize the wear behaviour of various materials under different fretting and sliding conditions [1–18]. Several fretting maps concepts have been proposed to identify the wear or fretting wear behavior where test conditions and damage evolutions are interrelated. Vingsbo and Soederberg introduced this approach by correlating the displacement and normal contact force to identify the partial slip, gross slip or mixed fretting regimes, which was referred to as the running condition fretting map [1]. Later, Fouvry and co-workers expanded this approach by including several other tests as well as material
∗ Corresponding author. Tel.: +1 937 255 3636; fax: +1 937 656 7053. E-mail address: Shankar.Mall@afit.edu (S. Mall). 0043-1648/$ – see front matter. Published by Elsevier B.V. doi:10.1016/j.wear.2008.12.083
response parameters [2–6]. They also conducted a series of wear and fretting wear studies in a detailed way to develop fretting maps as well as to investigate the wear extensively without any applied load on the substrate [7–13]. The fretting wear behavior of SC 65-2 uncoated and TiN coated steel against an alumina ball was investigated by this group of researchers [7–9]. They used gross sliding condition. The WV (wear volume) was obtained from the wear scars produced on the flat surface without any applied load. It was noted that the maximum wear occurred at the contact borders and minimum at the center. This led to a typical W-shaped scar, which changed with number of cycles or slip amplitude, and the wear was found to have the characteristic U-shape for large sliding amplitudes. Here, the wear volume was measured using three-dimensional as well as two-dimensional topographies of wear scars. The modeling of wear indicated that calculation of wear using a two-dimensional topography is much less time consuming and sufficiently accurate for modeling work. They confirmed that wear volume could be correlated with the cumulative dissipated energy with a linear relationship. The TiN coating was found to be about nine times wear resistant as the plain HSS substrate. Working with coated and uncoated HSS steel against an alumina ball, Fourvy et al. [10] noted that Archard’s wear model does not work well since friction coefficient is not constant. They suggested that use of interfacial shear work is more appropriate. The wear volume was linearly related to the interfacial shear work. The wear energy coefficient determined by this method accurately classified the wear resistance of various tribosystems. Fouvry et al. [11] also studied the fretting behavior of Ti–6Al–4V for a plain cylinder/plain contact configurations under ambient laboratory condition again without any applied load on the substrate.
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Fretting wear response of the system was studied under gross slip condition. They noted that to correctly quantify the wear behavior, the total wear volume of both the sliding bodies must be considered. They developed a fretting wear parameter in terms of pressure, sliding amplitude, test duration, and contact dimensions to quantify the wear kinetics. Fouvry et al. [12] also studied the fretting wear behavior of Ti–V–C and Ti–V–N unloaded systems sliding against an alumina ball. It was observed that for the Ti–V–C system, the wear debris was ejected from the interface; the wear mechanism is driven by a two-body abrasion displaying conventional U-shaped wear scar morphology. The wear rate decreased with hardness. For the Ti–C–N system, the three body abrasion process was modified. Oxide debris agglomerated on the alumina counter body, shifting the interface to a three body abrasion promoting typical W-shaped scar morphology. The wear rate dependence with coating hardness could not be verified for this case. Fridrici et al. studied the fretting wear of shot-peened and un-peened Ti–6Al–4V also without any applied load on the substrate [13]. They used cylinder-on-flat configuration and varied both the relative slip amplitude and contact load in their tests. They found that shot-peening has no effect on fretting wear behavior of Ti–6Al–4V at a given relative slip amplitude and contact load using the dissipated energy method. However, the energy wear factor was found to be dependent on relative slip amplitude. Huq and Celis analyzed the wear of several hard coatings using the dissipated energy [14]. They concluded that energy dissipation is a useful tool to quantify the wear damage, and its empirical parameters are different for unidirectional and reciprocating sliding conditions. McColl et al. [15] analyzed the fretting wear using finite element analysis based on Archard’s equation for a cylinder-on flatcontact. They determined surface and sub-surface stresses in the contact zone. It was shown that for gross slip condition, the high wear rate leads to contact edges to move rapidly outwards, leaving the material previously at the contact edges in a permanently compressive state, which retards the fretting-fatigue crack initiation. Thus, they suggested that wear becomes the predominant feature of the gross slip condition in the absence of fretting-fatigue. Jin and Mall designed and developed a fretting test set up which was capable of applying independent pad movement under a normal force and stress amplitude condition [16]. They investigated the fretting wear behavior of Ti–6Al–4V under gross slip condition for cylinder-on-flat contact configuration, but they did not measure the amount of wear [16,17]. The fretting-fatigue life was found to be larger for the gross slip condition than that for the partial slip. The morphology of scar profile was also studied, which was found to be W-shaped. The large wear at the edge of contact was attributed to abrasion caused by the detached and fragmented wear particles. Lee et al. investigated the fretting behavior of soft Cu–Al coating on Ti–6Al–4V substrate with and without fatigue load [18]. Energy approach of wear analysis showed a linear relationship between wear volume and accumulated dissipated energy. This relationship was independent of fatigue loading condition and extended from partial slip to gross slip regimes. The above review of the previous studies clearly shows that although a large amount of research has been conducted in the area of fretting wear, but none of these previous studies were conducted where substrate was subjected to an cyclic axial bulk load except by Lee et al. on coating. This is where the present work differs from the previous investigations [1–17]. In the present study, a cyclic axial load was applied on the flat specimen to investigate the fretting wear behavior of Ti–6Al–4V unlike the previous studies where substrate was without any applied load. The present authors have also investigated the frettingreciprocating wear behaviour of a titanium alloy in contact with the same titanium alloy (i.e. both Ti–6Al–4V) [19], since this material is widely used in aerospace and biomedical industry due to
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its superior mechanical properties and relatively low density. For this purpose, four series of tests were conducted to characterize the effects of different sliding and loading parameters on the wear behaviour under cylinder-on-flat contact configuration; and these parameters were relative slip, normal contact load, number of cycles, and bulk cyclic stress. The present study is an extension of this previous study [19] where wear behaviour of same material (Ti–6Al–4V) was characterized with new contact geometry (flaton-flat) and under the lubricated surface condition. These new data along with data from the different conditions from a previous study were combined and analyzed to characterize the wear behaviour of Ti–6Al–4V. As mentioned earlier, it should be reiterated that the focus of this study is fretting-reciprocating sliding condition unlike the conventional plain reciprocating sliding condition in the previous studies [1–17]. 2. Experiments 2.1. Test set up A test machine was used that could apply alternating load on the specimen and reciprocating motion to the pads independent of each other. It was a modified servo-controlled MTS fatigue test machine with two actuators. The details of this test machine are provided in [16–19]. A lubricating system was also set up to get lubricant between the specimen and pads. Lubricant was brought from a reservoir, which was placed above the test frame. Small capillary tubes were used so that lubricant could be dropped in the region where the pads contacted the specimen and ensured that lubricant entered into the contact area such that both the pads and specimen were lubricated well. SAE 15W-40 Heavy Duty Motor Oil was used in this study as the lubricant. Though it is not exactly the same lubricant as is used in the gas turbine engines, its lubricating properties are very similar. This arrangement simulated well the practical conditions of dovetail joint in the gas turbine engines. Two extensometers were mounted on the fretting fixture. One extensometer was attached to the specimen and the pad holder to measure their relative displacement. The second one was attached to the pad holder and a fixed point, which measured the absolute displacement of the former. These extensometer measurements were then used to calculate the structural compliance of the test rig, which was then used to compute the relative slip (ı) between the specimen and pad from the applied displacement, ıapp from the test machine to the pad [16–21]. This provided the relative slip, ı by incorporating the effects of both the structural compliance and specimen as elaborated in previous studies [16–21]. The normal contact load (P) on the pad as the function of applied cycles was continuously monitored by the load cell in each test. The P decreased in the tests with increasing number of cycles, especially in the longer running tests (i.e. with number of cycles). This decrease in P was caused by the wear of the pad and specimen [19]. This decrease in normal contact load was duly accounted for in the analysis of all test results. The tangential force produced at the contact was monitored from the loads measured at the both sides of the specimen. The measured tangential force (Q) for each of the two fretting contacts (one on each side of the specimen) was one-half of the difference of these loads. 2.2. Material Titanium alloy, Ti–6Al–4V was the test material because this material is widely used in aerospace applications, experiencing fretting wear damage [2,6,17–19]. The dog-bone shaped specimens having a width of 6.35 mm and thickness of 3.81 mm were used (Fig. 1a). The cylindrical fretting pads had a radius of 50.8 mm
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men. In test series #5, all test conditions were same as that “baseline series (#1) and also a lubricant was used in this case which was 15W-40 oil to examine the effect of lubrication on the wear mechanism, development of scar shape and the wear volume. In test series #6, a flat pad was used with loading conditions similar to “baseline” series except that ıapp = 1 mm. Several tests were conducted in each test series where specimens were subjected to different number of cycles in laboratory air at ambient room temperature at a frequency of 2 Hz. 3. Results and discussion 3.1. Wear scar details Fig. 1. Schematic of: (a) specimen, (b) cylindrical, and (c) flat fretting pad.
(Fig. 1b) and the flat had an edge radius of 2.54 mm with a central flat length of 4.45 mm (Fig. 1c). The material had an elastic modulus of 126 GPa, yield strength 930 MPa, ultimate tensile strength of 1 GPa and Brinell hardness of 302. 2.3. Wear Taylor Hobson “Ultra” profilometer was used to measure the wear volume. The profiles were taken over the scar surface in the direction of the fretting-reciprocating motion at 250 m intervals, and the above process was repeated until the complete scar was scanned. All the profiles from a surface were then assembled using MATLAB software developed to generate a complete topography of the wear scar. Wear volumes were computed using the trapezoid rule of numerical integration process. Each test in this study produced four wear scars: one on each side of the specimen and two on the corresponding fretting pads. The combined pad and specimen wear volumes were considered as the wear volume as opposed to the specimen or pad wear volume only [2,6]. It is because the material transfer between the pad and specimen can distort the perception of the actual wear damage if only either the pad or specimen wear volume is used, the reported wear volume was calculated by dividing the sum of four wear volumes (two pads and two sides of the flat specimen) by two. Finally, the wear scars on the flat specimen were examined by optical and scanning electron microscopes to elucidate the wear mechanisms operating in fretting-reciprocating sliding conditions. The specimens were ultrasonically cleaned to ensure that no surface features are hidden by the wear debris.
Fig. 2a shows the typical W-shaped wear scar on the specimen from a test in the “baseline” series after an application of 22,000 fretting cycles. Fig. 2b shows a scar with one broad valley in the middle. This scar is from the “lub” series test also after again 22,000 fretting cycles. This type of scar is termed as U-shaped scar. Thus, the fretting tests yielded in two types of wear scars: scars with multiple peaks and valleys (W-shaped) and scars with one wide valley (Ushaped). The evolution of two-dimensional view of the scar profiles of “baseline”, “flat”, and “lub” series tests are compared in Fig. 3. These profiles are in the longitudinal direction (i.e. sliding direction) at the middle of the specimen which provides the development and progression of wear scar with increasing number of fretting cycles. As the pad wore down deeper into the specimen surface, the
2.4. Test plan A total of six test series were considered; four from a previous study [19] and two from the present study to investigate the effects of relative slip amplitude, normal contact load, contact geometry, lubrication, loading cycles and bulk axial stress on fretting wear behaviour of Ti–6Al–4V. These were (1) the “baseline” series, (2) “baseline” series with ıapp = 1 mm (will be referred to as “low-ı” series), (3) “baseline” series with P = 4003 N (will be referred to as “high-P” series), (4) “baseline” series with = 0 (will be referred to as “no-” series), (5) “lubricated” series with 15W-40 lubricant (will be referred to as “lub” series), and (6) “flat pad” series with loading conditions similar to “baseline” series except that pad was flat (unlike the cylindrical in previous five series of tests) and ıapp = 1 mm (will be referred to as “flat” series). Test series #1 to #5 used the cylindrical pads. In the “baseline” series test (#1), the cylindrical pad was loaded against the flat specimen with a load, P of 1334 N, and subjected to an applied slip (ıapp ) of 1.5 mm, and an alternating bulk stress = 17–570 MPa was applied on the flat speci-
Fig. 2. Three-dimensional profile of (a) “baseline” series and (b) “lub” series; 22,000cycle specimen scar, ıapp = 1.5 mm; initial P = 1334 N, = 17–570 MPa.
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the figure and flaking of the contact surface in the upper portion. The sliding is in the vertical direction. Severe plastic deformation, which was noted for all cylindrical pads as shown in Fig. 4c, was absent in this case. For the sake of comparison, Fig. 4c shows a typical SEM micrograph of the middle region of scar from test series #1 to 4. This difference may be due to the reason that the pressure and shear stress at the surface were much lower in the flat case. However, a number of wear particles can be seen lying over the surface, Fig. 4a. It is these particles which are responsible for the abrasion grooves. Thus, wear mechanism was the adhesion, breakage of adhesive bonds, which form wear particles, and finally the abrasion caused by the wear particles. Fig. 4b is another micrograph obtained from the flat pad at the edge of the scar. This figure shows light deformation with material movement in the direction of the sliding. The deformation is mainly occurring at the asperity peaks only. Few wear particles can be seen over the surface but they were unable to form deep grooves. This may be due to the rounded edge of the pad, which did not produce sufficient pressure and also because the particles were not sufficiently strain hardened to penetrate the material. Fig. 4d shows a SEM micrograph obtained from the middle of the lubricated scar. This scar surface is smooth indicating the absence of severe plastic deformation and adhesion, which was expected for a lubricated contact. The surface is covered with circular-to-elliptical regions with dimensions of 40 m × 60 m. These are the material grains with very faint sliding marks. Thus, the wear mechanism for the lubricated contact was plastic deformation of high asperities followed by their removal due to polishing action. 3.2. Wear volume
Fig. 3. Evolution of wear scar with cycling for (a) “baseline” series, (b) “flat” series and (c) “lub” series; ıapp = 1.5 mm for (a) and (c), ıapp = 1.0 mm for (b), initial P = 1334 N, = 17–570 MPa.
contact area increased initially and then remained constant, which occurred around 22,000th cycle for all the series tests. Note that the “baseline” series scar has three valleys and two peaks; “flat” series profiles have one peak and two valleys; whereas the “lub” series has only one wide valley. For rest of the series tests, the scar shape was similar to that of the “baseline” series. The progression of scar lengths shows initial high rate of wear due to peak contact pressure, which decreases and the shear stresses are redistributed over the wider contact surfaces [15,22]. Therefore, the wear rate started decreasing with increasing number of cycles. Fig. 4a and b shows SEM micrographs obtained from the middle and edge of the wear scar produced on the specimen by the flat pad with rounded edges, respectively after 60,000 fretting cycles. Fig. 4a shows a number of abrasion grooves in the lower portion of
3.2.1. Wear versus fretting cycles Relationships between wear volumes and number of cycles (N) for all six series of tests are shown in Fig. 5, where each data corresponds to an independent test. As expected the wear volume increased with numbers of cycles. The wear volume is clearly dependent on the applied loading condition, i.e. P and ı, and the interface condition. When compared to the “baseline” series tests, the increased contact load in the “high-P” series tests increased the wear volume for the same number of cycles, i.e. it had the maximum wear volume out of the six test series. Decreasing the applied relative slip in the “low-ı” series had the opposite effect, i.e. it had the minimum wear volume. The effect of no-bulk stress on the specimen (“no-” test series) resulted in tests with roughly 200 m shorter relative slip than the “baseline” series tests. Differences in the effective relative slip explain the reason for the “no-” series wear volume curve being between the “baseline” and “lowı” series curves. The wear volume for the flat pad is smaller than the “baseline” series but larger than the “no-” series. This may be attributed to the lower contact pressure between the flat pad and the fretting specimen and lower ıappl (=1 mm). For the lubricated condition, the wear volume was expected to be lowest of all the cases. However, it was higher than the “low-ı” series. This may be attributed to several reasons: (a) the larger slip between the pad and the specimen due to lower coefficient of friction, (b) rapid rate of wear particle removal from the contact zone, and (c) corrosion of the contact zone by the extreme pressure additives in the lubricant, because 15W-40 oil is not formulated for a titanium alloy. Furthermore, the slope of each curve decreases with the increasing number of cycles, indicating a decreasing wear rate. This reduction in wear rate can be attributed to the decrease in normal load with number of cycles and the contact stresses due to increased area of contact. 3.2.2. Cumulative product of normal contact load and relative slip In this approach, the sum of the product of normal contact load and relative slip per cycle, (Pın ) were correlated with the wear
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Fig. 4. SEM micrograph of scar (a) at the centre of a “flat” series, (b) at the edge of a “flat” series, (c) at centre of “no-” series and (d) at centre of “lub” series.
volume as shown in Fig. 6. This shows a single linear correlation between wear volume and cumulative Pın for all the six series of tests of this study. Thus, it appears that the linear relationship between the wear volume and product of normal load and relative slip data is independent of the test parameters, i.e. contact load, relative slip, applied cyclic load on the substrate (specimen), contact geometry, contact load and surface condition. The slope of linear relationship between wear volume and cumulative product of contact load and relative slip ( Pın ) was 3.709 × 10−4 mm3 /(N m) with an R2 value of 0.982. The corresponding value for this material for unidirectional sliding is reported as 3.18 × 10−4 mm3 /(N m) [23]. Thus, the value of slope of this study is comparable to its counterpart as obtained from the classical unidirectional sliding method. The difference between these two values may be attributed to the
Fig. 5. Wear volume versus number of fretting cycles relationships.
several factors, such as sliding condition, test apparatus, material heat-treatment and microstructure, and the operating wear mechanisms. 3.2.3. Dissipated energy This is an alternate approach which could be used to characterize the wear behaviour. In this approach, wear volume is assumed to be proportional to dissipated energy ( Ed ) and is given by: WV = aV
Ed
(1)
where ˛V is a constant of proportionality and is called the energy wear factor. Here, the dissipated energy, Ed for a single cycle is the
Fig. 6. Wear volume versus sum of product of contact load and relative slip relationship.
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cyclic load on the substrate (specimen). For these different fretting test parameters, the proportionality constant of linear relationship between wear volume and cumulative product of contact load and relative slip (Pın ) was 3.709 × 10−4 mm3 /(N m), and for the linear relationship between wear volume and dissipated energy was 7.019 × 10−4 mm3 /(N m) in the case of the Ti–6Al–4V. However, further tests are needed to confirm and generalize for other fretting conditions and materials. References
Fig. 7. Wear volume versus total dissipated energy relationship.
area within a hysteresis loop between tangential force and relative Ed is deterslip (Q–ı plot). The accumulated dissipated energy, mined by summing the areas of all loops of a test [18,19]. This study used the actual area within the fretting loops. For this purpose, a MATLAB based program was written to numerically integrate the Q–ı loops. The dissipated energy values were then summed together to determine the total energy dissipated for a specific num ber of cycles ( Ed ). Fig. 7 shows the variation of wear volume as a function of the total dissipation energy. The plot shows that wear volume is linearly proportional to total dissipated energy for all six test series. The least squares fit to the data resulted in a slope of the line, ˛V = 7.019 × 10−4 mm3 /J with an R2 value of 0.964. It may be noted that the correlation between the wear volume and dissipated energy data is excellent. Further, it is interesting to note that the linear relationship between the wear volume and dissipated energy data is also independent of the test parameters, i.e. contact load, relative slip, contact geometry, lubrication, and applied cyclic load on the substrate (specimen) as in the case of Archard’s approach. 4. Conclusions The wear behaviour of Ti–6Al–4V subjected to frettingreciprocating sliding condition was investigated to analyze the effects of contact load, relative slip, contact geometry, lubrication, and applied cyclic load on the substrate (specimen). Dry and lubricated surface conditions produced “W” (scar with multiple valleys) and “U” (scar with single valley) shaped scars on both specimen and pad, respectively. The wear volume was linearly proportional to cumulative product of contact load and relative slip (Pın ) as well as to the total dissipated energy (Ed ). These correlations suggest that linear relationships exist to characterize the fretting wear of a material which are independent of the contact load, relative slip, contact geometry, lubrication, and applied
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