The acting wear mechanisms on metal-on-metal hip joint bearings: in vitro results

Wear 250 (2001) 129–139

The acting wear mechanisms on metal-on-metal hip joint bearings: in vitro results M.A. Wimmer a , J. Loos b , R. Nassutt c , M. Heitkemper d , A. Fischer d,∗ b

a AO Research Institute, Davos, Switzerland Dutch Polymer Institute, Technical University Eindhoven Eindhoven, The Netherlands c Biomechanics Section, Technical University Hamburg, Harburg, Germany d Materials Science and Engineering, University of Essen, Essen, Germany

Abstract Metal-on-metal (MOM) hip joint bearings are currently under discussion as alternatives to metal-on-polymer (MOP) bearings. Some criteria under scrutiny are the wear resistance, the influence of wear particles on the surrounding tissue, as well as the frictional torque. In order to understand and control the wear behavior of such a bearing a close correlation between the microstructures of the alloys used and the acting wear mechanisms has to be found. Thus, commercially available CoCrMo-balls were tested against self mating concave pins in a physiological fluid at 37◦ C under reciprocating sliding wear (1 Hz). The compressive load was 750 N (body weight). For 2 × 106 cycles tests were carried out continuously and with periodically occurring resting periods. On the basis of the observed wear appearances the acting wear mechanisms are defined and evaluated as to their contribution to the wear behavior. Due to the high local contact stresses surface fatigue prevails initially. Cr– and Mo–carbides are fractured and torn off the surfaces bringing about additional surface fatigue by indentations and initiating abrasion. The weight loss can be predominately attributed to these mechanically dominated wear mechanisms. In a parallel occurring tribochemical reaction layers are generated from denatured proteins. These adhere rigidly to the surfaces and cover parts of the contacting surfaces avoiding adhesion. Thus, the wear behavior is mainly influenced by the alternating balance between surface fatigue and abrasion on the one side and by tribochemical reactions on the other side. © 2001 Elsevier Science B.V. All rights reserved. Keywords: Hip joint; Metal-on-metal bearing; Resting periods; Wear mechanisms

1. Introduction Metal-on-metal (MOM) hip joint bearings are gaining more and more acceptance as an alternative to conventional metal-on-polyethylene (MOP) bearings [1,2]. Some criteria in favor of MOM bearings are the excellent wear resistance [3], the influence of wear particles on the surrounding tissue, as well as the frictional torque [4,5]. Recent developments of MOM bearings indicate a three times less frictional torque during simulated gait than MOP bearings [6]. This is an interesting finding, because high friction was one of the reasons that the MOM prostheses lost their popularity at the end of the 1960s, after being quite fashionable for more than one decade. The idea of an all-metal joint was taken up again in the mid 1980s and led to the development of the so-called second generation metal articulation with improved alloy microstructure, surface finish and manufacturing tolerances. The family of cobalt–chromium–molybdenum alloys, which are suitable for self-bearing applications, include both ∗ Corresponding author. Tel.: +49-201-183-2655; fax: +49-201-183-2508. E-mail address: [email protected] (A. Fischer).

cast and wrought materials with either low or high carbon contents. The alloy composition for cast and wrought surgical implant applications is specified in ISO 5832-12 [7], ASTM F75-98 [8], and ASTM F1537-94 [9], respectively. Typically, the carbon content is about 0.2% for the high carbon and below 0.08% for low carbon alloys. Carbon is responsible for the generation of carbides, which strengthen the material and affect the wear resistance [10]. In order to understand and control the wear behavior of these materials, a close correlation between the microstructures and the acting wear mechanisms has to be found. 2. Experimental procedures 2.1. Materials, metallography and microstructure Wrought CoCrMo balls and pins with a chemical composition given in Table 1 were used. The microstructure of these specimens was obtained by sectioning pins and balls and grinding them using SiC paper with 1200 mesh size (Struers, Willich, Germany). Afterwards polishing was carried out using diamond (Struers, Willich, Germany) with a

0043-1648/01/$ – see front matter © 2001 Elsevier Science B.V. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 1 ) 0 0 6 5 4 - 8

130

M.A. Wimmer et al. / Wear 250 (2001) 129–139

Table 1 Chemical composition in wt.% of the tested materiala Element

Cr

Mo

Ni

Co

Fe

C

ISO 5832-12 As-measured

26.0–30.0 29.7

5.0–7.0 6.2

<1.0 <1.0

Bal Bal

<0.75 <0.7

<0.35 0.25

a The metallic elements were determined by EDX analysis, while C was measured wet-chemically. This composition fulfills the criteria of ISO 5832-12 [7].

grit size of 6 ␮m for 600 s and with 1 ␮m size for 900 s. The final polishing step was done using polishing suspension (OPS, Testa, Kempen, Germany) for 300 s. The specimens were electrolytically etched (50 ml distilled water + 50 ml

hydrochloric acid +5 g chromium-4-oxide) for 20 s with 2 V at 20◦ C. The forged microstructure shown in Fig. 1 consists of a CoCrMo solid solution, which solidifies primarily from the melt forming metal cells and eutectic carbides of M23 C6 – (M: Cr, Co) and M6 C– (M: Mo, Cr) type. While the eutectic M23 C6 carbides precipitate predominantly at grain boundaries as well as within the grains (Fig. 2a) thereby revealing a size of about 1–5 ␮m, M6 C precipitates only within the grains (Fig. 2b) and are much smaller. This microstructure still prevails after forging. The average hardness value is 518 ± 11 HV10, while the micro-hardness of the metal matrix is 434 ± 30 HV0.05. The hardness of the eutectic carbides is reportedly higher and scatter between 1170 and 1300 HV0.05 for M23 C6 [11].

Fig. 1. SEM secondary electrons contrast (SE) micrograph of the forged CoCr29 Mo6 . (a) Eutectic M23 C6 carbide precipitated at a grain boundary; (b) M6 C carbide precipitated within the metal matrix.

Fig. 2. TEM micrographs and diffraction patterns of carbides.

M.A. Wimmer et al. / Wear 250 (2001) 129–139

Fig. 3. Structure of the pin-on-ball tribological test system.

2.2. Tribological test system — reciprocating wear test A special testing device (“pin-on-ball”, POB) was built, to account for the bearing geometry as well as the contact zone kinematics. The theoretical concept behind the simulator and the technical details of the screening device are described elsewhere [12]. Briefly, a pneumatic short-stroke cylinder assists to press two opposed concave pins onto a commercially available femoral head (Fig. 3). The 2D interface motion is generated by axial oscillation of both the pins and ball. By adjusting a 90◦ phase shift between both amplitudes, elliptical displacement trajectories with crossing paths are generated. A total of 24 pins and 12 balls with nominally 28 mm head diameter were put together to 24 articulations with a radial clearance of 57±2 ␮m. Metal pins (having a cylindrical diameter of 12 mm) were worked out of commercially available cups (Mathys Orthopaedics, Switzerland) to maintain prosthetic-like surface conditions. The heads were taken from stock. The average roughness Ra , the maximum roughness Rmax , and the 10-point height Rz were evaluated before and after the test using a laser-profilometer (UBM Messtechnik GmbH, Ettlingen, Germany). Measurements were performed according to DIN 4768 with a minimum of four traces for each sample. Wear tests were carried out in two groups with six test stations each. In group I (continuous mode), 2 × 106 cycles of continuous oscillation were applied, in group II (resting mode), 2 × 106 cycles with resting periods were carried out. In the resting group, 30 s of oscillation were followed by 15 s of rest. This reflects just about the average duration and frequency of resting periods during daily activity of total hip replacement patients [13]. In both groups, the test force was set to constant 750 N, which is the equivalent of one body weight. The angular amplitudes of pin and ball were set to 30◦ . The frequency was adjusted to 1 Hz, which produced an average sliding velocity of 29.3 mm/s between head and pins. For lubrication, each testing station (i.e. one ball and two pins) was placed in a Plexiglas® container with 120 ml tempered fluid (37◦ C). The latter was a mixture of 80 ml distilled water containing 9 g/l sodium chloride and 40 ml ‘new born’ calf serum, such that the salt

131

(sodium, potassium, calcium, chloride) and the total protein constituents approximated healthy synovial fluid. Additives included antibiotics to retard bacterial degradation as well as Tris–hydroxy-methylamine to minimize precipitation of calcium phosphate and to buffer the pH in the region of 7.6. With respect to the necessary system analysis [14,15] the body is the ball, while the pins are designated as counter bodies. The interfacial medium is the synovial fluid, which had to be cooled in order to gain a constant temperature of 37◦ C (307.5 K) during the tests. The surrounding medium is laboratory air (24◦ C, 40% RH). The type of interaction of body and counterbodies is sliding. Thus, the type of wear can be defined as reciprocating sliding wear. On the basis of this macroscopic analysis all four major wear mechanisms (abrasion, surface fatigue, adhesion, tribochemical reactions) can govern the wear behavior of this tribosystem. The wear rates were calculated by the weight loss G of the ball and two pins, which were measured before and after testing, subsequent to thoroughly cleaning of the test bodies (see Section 2.3 for details). These values were then divided by the density of the material ρ, the nominal area of contact A, and the length of the wear path L. This brings about the linear wear rate with dimension m/m, which then was divided by two in order to achieve the wear rate per articulation. To more precisely evaluate wear of pins and head with cycling serum samples were taken after ten defined intervals (50, 100, 250, 500, . . . , 2 × 106 cycles). In order to retrieve a representative sample, the content of the 1200 ml chamber was homogenized prior to retrieval by fluid circulation and stirring. An amount of 2 ml of serum were then pipetted into a polypropylene tube and thoroughly mixed ultrasonically, immediately prior to digestion. Pressurized digestion of metal particles and serum constituents were performed in a quartz glass container with one part serum and two parts nitric acid using a microwave. The simultaneous measurement of cobalt and chromium was performed using graphite furnace atomic absorption spectrometry. Due to the 10-fold smaller concentration of molybdenum, this element was determined in a separate trial. Technical details, the calibration protocol, and the validation of method are reported in [16]. In addition, friction torque measurements were carried out using the same screening concept as for wear testing. The relative change in friction before and after the wear experiments was determined for continuous and interrupted motion. The dynamic and static friction coefficients were calculated from the measured torque. All friction torque measurements were performed under uniaxial oscillation of the ball at 1 Hz and ±20◦ angular amplitude. Friction torque was measured for 3 s during continuous oscillating motion and during motion initiation after a pre-defined period of rest using a torque transducer (30 ± 0.025 Nm) at 2.5 kHz sampling rate. Six pairings (three from each group I and II, respectively) were investigated. Three test repetitions were stored for each identical pairing under the same pre-defined condition. Statistical analyses were performed at a 95% confidence level.

132

M.A. Wimmer et al. / Wear 250 (2001) 129–139

2.3. Wear appearances and surface topography analyses After the wear test, samples were rinsed in tap water, soaked in acetone and then ultrasonically cleaned in deionized water. All components were inspected without further preparation using scanning electron microscopy (SEM). 2.3.1. Scanning electron microscopy The morphologies of the contact surfaces (pin and ball) were investigated using predominately a Philips environmental SEM XL-30 ESEM FEG (Phillips, Eindhoven, The Netherlands) equipped with energy dispersive X-ray spectrometer (EDX) for local and area distribution analyses of metallic and nonmetallic elements. Secondary electron (SE) imaging of the sample surfaces was performed using acceleration voltages of 1 or 20 kV, whereas backscattered electron imaging (BSE) and qualitative EDX analysis were carried out at accelerating voltages of 5 and 20 kV. For an acceleration voltage of 1 kV, the penetration of the incident electron beam is in the range of a few tens of nanometers for the investigated materials. Thus, in addition to standard high acceleration voltage SEM, SE images acquired at 1 kV acceleration voltage show surface features in more detail, even at high magnification. For additional surface topographical and chemical analyses a ZEISS DSM 962 with a Noran Instruments Voyager 2.6 EDX analyzer was used with 20 kV in SE and BSE mode. 2.3.2. Atomic force microscopy A Digital Instruments Inc., USA, Dimension 3100 was used in tapping mode (non-contact) for atomic force microscopy (AFM) of topographic features of the sample surfaces.

Fig. 4. Coefficient of friction measured during one cycle after 2 × 106 cycles under continuous mode. The relative velocity is 0 m/s at 20 and –20◦ , while it is at its maximum value at 0◦ .

ter the wear test, and there was no statistically significant difference in friction values. After a short period of running-in, with cycling, the weight loss increased linearly for both the resting and continuous motion groups (Fig. 6). There was a tendency for higher wear rate for the resting group with 6.84 × 10−10 l per articulation while it was 6.34 × 10−10 l per articulation for the continuous motion group. However, there was no significant difference for any measured time interval. During wear the ratios between the amounts of Cr, Mo, and Co metal ions within Ringers solution reflected the alloy’s composition within ±1 wt.%.

3. Results 3.1. Frictional forces and wear rates After each motion reversal a static friction peak occurred (Fig. 4). This static peak was highest after the period of rest (Fig. 5), where it reached up to µ = 0.2 compared to µ = 0.11 during continuous motion. This frictional effect disappeared after one cycle and there was no difference in friction between those tests running continuously and those including resting periods. Interestingly, dynamic friction was rather unaffected due to motion interruption (Fig. 5). In both test groups, the (average) dynamic friction coefficient was µ = 0.098, having a minimum at highest velocity. Hence, the difference in dissipated energy was negligible: immediately after start-up, during the first cycle, the increase in dissipated energy was approximately 2% for the resting group when compared to the continuous group. All observations described above were similar before and af-

Fig. 5. Coefficient of friction measured during one cycle after 2.03 × 106 cycles under resting mode. The frictional torque is strongly influenced by stick phenomena at the beginning of that cycle.

M.A. Wimmer et al. / Wear 250 (2001) 129–139

133

Fig. 6. Weight loss vs. number of cycles f or continuous and resting mode wear tests.

3.2. Roughness

Fig. 7. Grooves and pits on the worn surface of the ball. Resting mode, 2 × 106 cycles SEM-BSE mode, high kV.

The roughness parameters of both pins and balls increased during wear testing. While the pins’ peak surface parameters (Rz and Rmax ) increased approximately two-fold, the balls’ parameters exhibited a six-fold enlargement (Table 2). Table 2 Roughness parameters of 22 metal pins and 10 balls before and after 2 × 106 cycles of wear testing Ra (␮m) Before wear test After wear test

Rz (␮m)

Rmax (␮m)

Pin

0.077 ± 0.020

1.224 ± 0.562

1.989 ± 0.900

Ball

0.071 ± 0.024

0.882 ± 0.116

1.028 ± 0.431

Pin

0.171 ± 0.054

2.239 ± 0.745

3.418 ± 1.173

Ball

0.500 ± 0.213

4.772 ± 1.391

6.489 ± 1.741

3.3. Wear appearances 3.3.1. Body — ball In the contact area, the ball exhibits grooves, which are predominately parallel to the sliding direction of the pin (Fig. 7). Pits of different sizes are visible in between these grooves. In addition to shallow elongated pits, which appear like delaminations, deeper circular pits from torn-off carbides as well as smaller pointed ones are observed. The latter seem to originate from detached particles, which indented into the metal matrix during rolling, or represent those areas, in which carbides had been fragmented. The AFM scan shows that these pittings are accompanied by grooves (Fig. 8).

Fig. 8. AFM image, which reveals the topography of grooves and pits. Resting mode, 2 × 106 cycles.

134

M.A. Wimmer et al. / Wear 250 (2001) 129–139

Fig. 9. Pitting within the metal matrix and delamination at the interfaces between carbides and metal matrix on the surface of a pin. Resting mode, 2 × 106 cycles, SEM-SE-mode, high kV.

Fig. 10. Layers, grooves, and Indentations on the surface of a ball. Resting mode, 2 × 106 cycles, SEM-BSE mode, high kV.

3.3.2. Counterbody — pin Pitting and delaminations prevail within the metal matrix on the worn pin surface (Fig. 9). Obviously, the latter appear mainly in the vicinity of carbides. In particular, the interfaces between the carbides and the metal matrix are torn open, indicating the weakest constituent. 3.3.3. Ball and Pin On both bodies distinct layers, which adhere rigidly to the surfaces, have been observed within and — larger in number — at the boundaries of the contact zones (Figs. 10 and 11). With respect to the EDX analyses (Fig. 12) they consist mainly of C and contain small amounts of chemical elements from the interfacial (Na, Cl, K, Ca, P, O) and the surrounding (O) media. The shapes of these layers vary in a wide range between hilly and compact to flat and lengthy. Some of them show cracks (Fig. 11) and have a rim

Fig. 11. Compacted and segmented layer on the surface of a ball. Resting mode, 2 × 106 cycles, SEM-SE mode, 20 kV.

Fig. 12. EDX spectrum of a layer on the surface of a ball. Resting mode, 2 × 106 cycles, 20 kV. (a) SE mode, 1.0 kV, note the different contrast of the immediate vicinity of the layer. (b) BSE mode, 20 kV, note that the rim surrounding the layer is no longer visible.

M.A. Wimmer et al. / Wear 250 (2001) 129–139

135

Fig. 13. Compacted and segmented layer on the surface of a pin. Resting mode, 2 × 106 cycles.

Fig. 14. EDX spectrum at 10 kV of a compacted layer resting mode, 2 × 106 cycles.

of a thin oxide film (Fig. 13). The comparison of SE (Fig. 13a) and BSE (Fig. 13b) contrast of a flat and lengthy layer shows that according to the atomic weight there is a marked contrast between the C-containing layer and the base material and no contrast between the rim and the substrate (Fig. 13b). Taking this and the EDX analyses (Fig. 14) at 10 kV into account one can assume that the layer mainly consists of C and the rim of O and Cr.

4. Discussion The type of wear in this tribological test system is reciprocating sliding wear, and therefore, all known major wear mechanisms may act at the same time [14]. In order to find a close correlation between the microstructure and its wear properties, the acting wear mechanisms and the failure sequence have to be known [17]. In doing so, one should distinguish between mechanisms which are predominantly

of mechanical nature (surface fatigue, abrasion) and mechanisms which are of chemical and mechanical nature (adhesion, tribochemical reactions) [18]. Here, with respect to the wear appearances, we observed abrasion (Figs. 7 and 8), surface fatigue (Figs. 7–9), and tribochemical reactions (Figs. 10–13). Surprisingly, there were no signs of adhesion even though this is a self-mating MOM contact system. Abrasion is an often reported mechanism in MOM and other hip joint bearings, because scratches and grooves are always obvious [3,19–24]. Abrasion may be induced by foreign particles (contaminants from outside the system), or most likely, from system inherent particles, like fractured carbides, compacted wear debris, and plastically deformed parts of the metal matrix. Even though the main wear loss is correctly attributed to abrasion in many references it is still questionable, which mechanism initiates wear and what is the presumable failure sequence. In order to understand this the local mechanical and thermal stresses have been estimated.

136

M.A. Wimmer et al. / Wear 250 (2001) 129–139

4.1. Mechanical stresses The local mechanical stresses can be calculated according to Hertz [25]. The mismatch between the radii of ball and pin is 0.057 mm. With Young’s moduli E = 210 GPa and Poisson’s ratios ν = 0.3 for ball and pin and the contact force of 750 N the circular contact area A is 2.05×10−5 m2 and the maximum Hertzian pressure p0 is approximately 55 MPa. Due to the fact that hard phases like carbides protrude from the surfaces under sliding wear [26], the real area of contact changes steadily during the wear process between two limits. Limit one is theoretically given by the Hertzian contact area and limit two is brought about by direct contacts of the carbides of ball and pin surfaces. Since there are about 4 vol.% of M23 C6 carbides on both bodies the area fraction of carbide/carbide contacts is 0.04 × 0.04 = 0.0016. Taking this fraction and the Hertzian contact area into account the real area of contact scatters between 2.05 × 10−5 and 3.28×10−8 m2 . This would result theoretically in local contact stresses between p0 and 22 GPa, which is twice as much as the hardness of the carbides with roughly 11 GPa. This indicates that mixed local carbide/carbide, metal–metal, and carbide/metal contacts are necessary to achieve equilibrium. With respect to the observed fracturing of carbides (Fig. 9) the local contact stresses indeed become very high. In addition one has to consider that metal–metal contact will be rare, because the metallic surface of CoCrMo solid solution is instantaneously covered with a thin film of a metal-oxide of Cr2 O3 -type. These passive layers reach a thickness of approximately 10–30 nm. If the layers are cracked by high plastic deformation close to the surface, a direct metal–metal contact appears bringing about adhesion. However, no signs of adhesion have been observed, and therefore, the oxide layers or other features seem to have a beneficial influence on wear with a very strong influence on contact chemistry. The regeneration of these layers is accelerated by orders of magnitude under tribological contact situations because of the high energy stored in deformed metal and high defect density. This hinders adhesion and lowers the wear rate under sliding wear [27–29]. 4.2. Thermal stresses On the basis of the Hertzian contact stress calculation and the parameters chosen for the tests, the increase of the average temperature as well as of the flash temperatures can be derived according to Kuhlmann-Wilsdorf and coworkers [30–32]. With a hardness of the alloy of 5.18 GPa and the physical properties of a CoCrMo solid solution, the average temperature increase of dry sliding is about 5.9 K (see Appendix A). For the flash temperature calculations the contact stresses are limited to the hardness of the respective constituent. Due to the fact that the local contact stresses on protruding carbides may be as high as 11 GPa it can be assumed that dry sliding conditions apply. This leads to a steep increase of flash temperatures in a carbide/carbide contact of

about 34 K for a time period of about 43 ms (see Appendix B). Neither the properties of the metal matrix nor those of the carbides are influenced by these temperatures; however, the biological components of the lubricant may change. The following sequence and interaction of wear mechanisms are drawn based on the wear appearances investigated in this work, the estimations of the mechanical and thermal stresses, as well as the influence of the initial microstructure on the contact situation. 4.3. Sequence and interaction of wear mechanisms Starting with smooth polished surfaces the contact situation is governed by cyclic local stresses which scatter theoretically between 55 MPa and 22 GPa depending on the contact of the solid constituents. Despite recent findings [21], which have to be discussed, we consider, that the fluid does not bear mechanical contact stresses under this low relative speed and high normal force, if carbides are in contact with themselves or with the matrix. In carbide/carbide contact, the local pressure may become higher than their hardness, which is physically impossible. Thus, the maximum possible local pressure is limited to 11 GPa. Above this value protruding carbides are fractured or pressed into the metal matrix to achieve a mechanically balanced situation between the remaining matrix and carbide contact spots. A direct metal–metal contact would be possible in areas of high plastic deformation, and a metal-oxide/metal-oxide contact in areas of low plastic deformation. The oxides might spall off the surfaces by surface fatigue resulting in wear particles of compacted metal-oxides [28,33]. This type of surface fatigue can bring about two different routes of failure. One follows the assumption that an oxide layer is generated by tribochemical reactions and spalls off when reaching a critical thickness [27,29]. The other involves a wear particle that is generated by cyclic contact stresses that are so small that the particle oxidizes instantaneously when torn-off from the surface in order to reach a state of reduced free energy [33]. In both cases one finds wear particles, which have an oxide layer just at the surface or which are completely oxidized. The investigation of the wear particles does not lead to an unequivocal verification of which of these routes is valid and the analysis of the metal ions in the fluid does not help either. On a CoCrMo solid solution most likely Cr-oxides are generated together with metal-hydroxides whether there is mechanical activation or not. During the time span in which the metal surface is not covered by a repassivation layer shortly after being scratched or torn off all metal ions are dissolved by corrosion. Thus, very often the metal ions in the fluid just represent the chemical composition of the materials tested. These indirect testing methods, therefore, do not bring about a coherent view of what happens in the single contact spots and have to be taken very critically. Oxidized metal matrix wear debris is found after in vivo as well as in vitro tests [24,34–37]. One can assume that metallic or oxide particles are torn off the surfaces and act

M.A. Wimmer et al. / Wear 250 (2001) 129–139

afterwards as interfacial media. This leads to abrasion or surface fatigue depending on whether the particles stick to one surface and scratch the other or act as rolling particles bringing about indentations [38]. Both mechanisms can act at the same time if the normal loads are predominantly carried by these particles and not by the fluid. The effect of the much harder, torn off carbide fragments on the wear behavior is even more distinct. Again, small particles become loose and slide or roll between the contacting surfaces resulting in surface fatigue by indentation. The bigger particles, with a size exceeding the gap between the contacting bodies, become embedded within one of the bodies and scratch the counterface resulting in abrasion. The question appears which of the above mechanisms is the initiating one. With respect to the wear appearances observed in this tribosystem one can assume that even if the first particles are generated by tribooxidation, the most effective will be those, which are generated by surface fatigue leading to fractured or torn-off carbides (Fig. 9). The load is then carried by a very small number of contact points under high mechanical stress leading to additional abrasion (grooves) and surface fatigue (indentations). Both bring about loss of material and a steady removal of existing passive layers followed by repassivation. Putting all observations together, the stationary tribological behavior of this tribosystem can be attributed to a balance between mechanically acting mechanisms (abrasion and surface fatigue) and chemically and mechanically driven tribochemical reactions. Interestingly, we did not find any signs of excessive oxide layer formation but markedly compacted and partly segmented C-containing layers, which adhered strongly to the surfaces of balls and pins (Figs. 10–13). The only element of the tribosystem, which has enough carbon for the formation of these layers is the lubricant. The incorporated proteins decompose at temperatures above 330 K (60◦ C). Following the average (5.9 K) and flash temperature (34 K) estimations according to Appendices A and B, the authors showed that theoretically temperatures of about (273.5+37+5.9+34 = 350.4) 350 K can be attained in some contact spots by the present contact situation. Therefore, it is suggested that a solid layer of protein derived material will be generated between the contacting bodies. These layers act as solid lubricant and separate the bodies from direct metal–metal contact. This type of tribochemical reaction is accompanied by other wear mechanisms and brings about a distinct roughening of all surfaces (Table 2). The difference

in the roughening of pin and ball is most likely related to the dissimilar motion characteristics of pin and ball. The latter follows a more unidirectional movement, while the former experiences bi-directional movement with crossing motion trajectories. This enhances the described self-polishing or “self-healing” effect [19] of all-metal articulations. Nevertheless, the stick phenomenon after the resting periods may be attributed to the interlocking of the decomposed protein layers and other roughness features generated on both contacting surfaces.

5. Conclusions and outlook Extensive tribo-layers derived from proteins have been observed in MOM hip joint bearings after in vitro testing. On the basis of the observed wear appearances and the estimation of mechanical and thermal contact stresses a new failure sequence is presented and discussed. 1. The protein layers act as solid lubricant and have a beneficial effect on the wear behavior. They cover large areas of the surfaces, hinder surface fatigue, and prevent adhesion. 2. A possible detrimental effect may arise from the change of chemistry of the surrounding medium, which might have an influence on the in vivo performance. Up to now this is unknown and is the subject of future research. 3. In addition, further research work will include the analyses and comparison of explanted MOM hip joint bearings with respect to adhering protein layers. Acknowledgements M. Wimmer would like to thank the Robert Mathys Foundation, Bettlach, Switzerland for specimens and partial support of the study. A. Fischer and M. Heitkemper thank Prof. Dr. W. Dudzinski, TU-Wroclaw, Poland for assistance in analyzing the TEM diffraction patterns of carbides.

Appendix A. Determination of medium temperature in the contact area In addition to the Young’s modulus and the radii of the spheres the other variables have to be defined

Variables Sliding speed Friction coefficient Temperature of surrounding medium Ball radius, r1 Pin radius, r2

137

Materials constant v := 0.0461 m/s µ := 0.098 as-measured TU := 309 K 14.0 mm 14.05 mm

Thermal conductivity Density Heat capacity Hardness Young’s modulus

λ1 := 12.766 kg m/s3 K d1 := 8280 kg/m3 c1 := 452 m2 /s2 K H := 52 × 108 kg/s2 m E = 210 GPa

138

M.A. Wimmer et al. / Wear 250 (2001) 129–139

With these, the amount of heat q taken in can be determined and f is equal to the contact pressure, as this is lower than the hardness, otherwise f is equal to H q := µfv λr :=

λ2 λ1

The relation between the heat conductivities of the two counterparts is determined. The thermal diffusivity of the substrate and the velocity of the heat conduction in the material are determined as follows. The ratio between the velocity of macroscopic sliding and the velocity of heat conduction are calculated vr :=

v λ1 , v0 c1

v0 :=

λ1 , d1 c1 r

dT0 :=

π qr 4λ1

A reference temperature is calculated, which would appear under adiabatic conditions. From the ratio of velocities vr and the shape factor e, the constants Z, S and S0 (for details see [30–32]) are determined and with them the average contact temperature enhancement is found dT :=

dT0 (1/Z)S + (λr/S0 )

In this case, the average temperature increase is 5.9 K.

Appendix B. Determination of flash temperature in a carbide/carbide contact spot In contrast to Appendix A now the focus is on the single contact spot if it is brought about by protruding carbides. These have significantly different physical properties compared to the matrix. The friction coefficient is taken for dry sliding of carbide/carbide contact.

Variables Sliding speed Friction coefficient Temperature of surrounding

contact spots have to meet to be able to carry, so only 0.04 × 0.04 = 0.0016 of the macroscopic area is carrying. The remaining area of load is divided by the area of one basal plane of a carbide and results in the estimated number of contact spots Nk p. With the macroscopic load of P the radius of one contact can be estimated.
rk :=

P π Nk pHk

Similar to Appendix A with maximum f is equal to H of carbide the heat generation per contact qk := µk Hk v and the velocity of heat conduction are calculated v0k :=

√ κ1k π Nk pHk √ P

Also the ratios of heat conductivity (which is still = 1) and the ratio of velocities are determined. With these results and the constant Z0k (dependent on vrk ) (for details see [30–32]) flash temperatures for λrk :=

λ2k , λ1k

vrk :=

v v0k

different speed cases can be calculated. For the general case the equation √ µk (π PHk /Nk p)v T0 := 4λ1k ((1/Z0k ) + λrk ) leads to a temperature increase in the single carbide/carbide contact spot of 34.0 K.

Materials constant of carbides v := 0.0461 m/s µk := 0.2 TU := 309 K

For determination of the flash temperatures it is considered that the carbides are carrying the load up to a local contact pressure which is as high as their hardness. They have a volume fraction of about 4%. Two

Thermal conductivity Density Heat capacity Hardness

λ1k := 1.83 kg m/s3 K d1k := 7000 kg/m3 c1k := 543.8 m2 /s2 K Hk := 135 × 108 kg/s2 m

References [1] L.D. Dorr, Z. Wan, D.B. Longjohn, B. Dubois, R. Murken, Total hip arthroplasty with use of the metasul metal-on-metal articulation: 4–7-year results, J. Bone Joint Surg. Am. 82 (6) (2000) 789–798.

M.A. Wimmer et al. / Wear 250 (2001) 129–139 [2] C.B. Rieker, P. Köttig, R. Schön, M. Windler, U.P. Wyss, Clinical wear performance of metal-on-metal hip arthroplasties, in: J.J. Jacobs, T.L. Craig, (Eds.), Alternative Bearing Surfaces in Total Joint Replacement, ASTM 1346 STP, American Society for Testing and Materials, 1998, pp. 144–156. [3] H.P. Sieber, C.B. Rieker, P. Kottig, Analysis of 118 second generation metal-on-metal retrieved hip implants, J. Bone Joint Surg. Br. 81 (1) (1999) 46–50. [4] R.M. Streicher, R. Schön, M.F. Semlitsch, Investigation of the tribological behavior of metal-on-metal combinations for artificial hip joints, Biomed. Tech. 35 (5) (1990) 107–111. [5] S.C. Scholes, A. Unsworth, Comparison of friction and lubrication of different prostheses, Proc. Inst. Mech. Eng. H 214 (1) (2000) 49–57. [6] R.A. Poggie, R. Afflitto, K. St. John, The wear performance of precision Co–Cr–Mo alloy metal-on-metal hip bearings. in: Proceedings of the Conference Transactions of the 12th Annual International Symposium for Technology in Arthroplasty, Chicago, 23–25 September 1999, pp. XI:1–2. [7] ISO 5832-12, 1st Edition, Implants for Surgery, Metallic Materials. Part 12. Wrought CoCrMo Alloy. ISO, Casa Postale 56, 1211 Geneva 20, Switzerland, 1996, pp. 1–2. [8] ASTM F75-98, Standard Specification for Cobalt-28Chromium-6-Molybdenum Casting Alloy and Cast Products for Surgical Implants, American Society for Testing and Materials, 100 Barr Harbor Drive, Conshocken, PA, USA, 1998, 1 pp. [9] ASTM F1537-95 Standard specification for wrought cobalt-28-chromium-6-molybdenum alloy for surgical implants. American Society for Testing and Materials, 100 Barr Harbor Drive, Conshocken, PA, USA, 1995, 1 pp. [10] R. Varano, S. Yue, J.D. Bobyn, J.B. Medley Co–Cr–Mo alloys used in metal–metal bearing surfaces, in: J.J. Jacobs and T.L.Craig (Eds.), Alternative Bearing Surfaces in Total Joint Replacement, ASTM 1346 STP, American Society for Testing and Materials, 1998, pp. 55–68. [11] W. Theisen, Neue Hartlegierungen auf Ni- und Co-Basis für die Auftragschwei␤ung, Diss. Universität Bochum, 1988, s.a. Fortschr. Ber. VDI-Z, Reihe 5, Nr. 153, 1988, 1 pp. [12] M.A. Wimmer, R. Nassutt, F. Lampe, E. Schneider, M.M. Morlock, A new screening method designed for wear analysis of bearing surfaces used in total hip arthroplasty, in: J.J. Jacobs, T.L. Craig (Eds.), Alternative Bearing Surfaces in Total Joint Replacement, ASTM 1346 STP, American Society for Testing and Materials, 1998, pp. 30–43. [13] M.M. Morlock, R. Nassutt, A. Bluhm, M. Vollmer, M. Honl, V. Nüller, E. Hille, E. Schneider, Quantification of hip joint resting periods during daily activities, in: Proceedings of the Transactions of the 46th Annual Meeting of the Orthopaedic Research Society, Vol. 25, 12–15 March 2000, 497 pp. [14] DIN50320-1 Wear, Terms, System Analysis, Classification of Wear, Beuth, Berlin, Germany, 1979, pp. 1–8. [15] DIN50323-2 Tribology, Terms, Beuth, Berlin, Germany, 1995, pp. 1–13. [16] D.S. Artelt, J. Kunze, M.A. Wimmer, A new method of wear determination for metal-on-metal hip joints, in: Proceedings of the Transactions of the Swiss Society of Biomaterials, Davos, 8–9 June 2000, 1 pp. [17] K.-H. Zum Gahr, Microstructure and Wear of Materials. Tribology Series, Vol. 10, Elsevier, Amsterdam, 1987, 1 pp. [18] Czichos, H., Introduction to Friction and Wear, in: K. Friedrich (Ed.), Friction and Wear of Polymer Composits, Composite Material Series, Vol. 1, Elsevier, Amsterdam, 1986, pp. 1–23. [19] H. McKellop, S.-H. Park, R. Chiesa, P. Doorn, B. Lu, P. Normand, P. Grigoris, H. Amstutz, In vivo wear of three types of metal on metal hip prostheses during two decades of use, Clin. Orthop. 329 (1996) 128–140. [20] F.W. Chan, J.D. Bobyn, J.B. Medley, J.J. Krygier, S. Yue, M. Tanzer, Engineering issues and wear performance of metal on metal hip implants, Clin. Orthop. 333 (1996) 96–107.

139

[21] F.W. Chan, J.D. Bobyn, J.B. Medley, J.J. Krygier, M. Tanzer, Wear and lubrication of metal-on-metal hip implants, Clin. Orthop. 369 (1999) 10–24. [22] P.S. Walker, E. Salvati, R.K. Hotzler, The wear on removed McKee–Farrar total hip prostheses, J. Bone. Joint. Surg. Am 56 (1) (1974) 92–100. [23] M. Semlitsch, R.M. Streicher, H. Weber, Verschlei␤verhalten von Pfannen und Kugeln aus CoCrMo-Gu␤legierung bei langzeitig implantierten Ganzmetall Hüftprothesen, Orthopäde 18 (1989) 377– 381. [24] H.G. Willert, G.H. Buchhorn, D. Göbel, G. Köster, S. Schaffner, R. Schenk, M. Semlitsch, Wear behavior and histopathology of classic cemented metal on metal hip protheseses, Clin. Orthop. 32 (1996) 160–186. [25] H.R. Hertz, Über die Berührung fester elastischer Körper und über die Härte, Sitzungsbericht des Vereins zur Förderung des Gewerbeflei␤es, 1882, pp. 449–463. [26] A. Fischer, Einflu␤ der Temperatur auf das tribologische Verhalten metallischer Werkstoffe. Habilitationsschrift Ruhr Universität Bochum, 1992, Fortschr.-Ber. VDI-Z, Reihe 5, Band 378, VDI, Düsseldorf, 1994, pp. 115–118. [27] T.F.J. Quinn, NASA Interdisciplinary Collaboration in Tribology. A Review of Oxidational Wear, NASA Contractor Report 3686, 1983, 1 pp. [28] J.L. Sullivan, The role of oxides in the protection of tribological surfaces. Parts I and II. in: Procedings of the Conference on Tribology — Friction, Lubrication, and Wear. Fifty years On 1–3 July 1987, London, England, IMechE Publ. Ltd., P.O. Box 24, Northgate Ave., Bury St. Edmunds, Suffolk IP32 6BW, England, 1987, pp. 283–301. [29] R.Y. Lee, Z. Eliezer, On the critical film thickness of protective films at sliding interfaces, Wear 95 (1984) 165–175. [30] D. Kuhlmann-Wilsdorf, Demystifying flash temperatures. Part I. Analytical expressions based on a simple model, Mater. Sci. Eng. 93 (1987) 107–117. [31] D. Kuhlmann-Wilsdorf, Demystifying flash temperatures. Part II. First order approximation for plastic contact spots, Mater. Sci. Eng. 93 (1987) 119–133. [32] D. Kuhlmann-Wilsdorf, D.D. Makel, N.A. Sondergaard, Refinement of Flash Temperature Calculations, in: F.A. Smidt, P.J. Blau (Eds.), Proceedings of the Conference on Engineered Materials for Advanced Friction and Wear Applications, 1.-3.3.1988, Gaithersburg, MD, USA, ASM Int. (1988) 23–32. [33] W.J. Saleski, R.M. Fisher, R.O. Ritchie, G. Thomas, The Nature and Origin of Sliding Wear Debris from Steels, in: K.C. Ludema (Ed.), Proceedings of the Conference on Wear of Materials, Vol. 83, 11–14 April 1983, Reston, VA, USA, ASME, 345 East 47th Street, New York, NY 10017, USA, 1983, pp. 434–445. [34] I. Catelan, D. Bobyn, H. Vali, J.B. Medley, D. Zukor, A. Peut, O.L. Iluk, Characterization of metal–metal wear particles isolated from a hip joint simulator, in: Proceedings of the 6th Conference on World Biomaterial Congress Transaction, Society for Biomaterials, USA, 2000, 482 pp. [35] P.F. Doorn, P.A. Campbell, J. Worrall, P.D. Benya, H.A. McKellop, H.C. Amstutz, Metal wear particle characterization from metal on metal total hip replacements: transmission electron microscopy study of periprostehtic tissues and isolated particles, J. Biomed. Mater. Res. 42 (1998) 103–111. [36] P.F. Doorn, P.A. Campbell, H.C. Amstutz, Metal vs. polyethylene wear particles in total hip replacement, Rev Clin. Orthop. 329 (1996) 206–216. [37] H.L. Anissian, A. Stark, A. Gustafson, V. Good, I.C. Clark, Metal-on-metal bearing in hip prothesis generates 100-fold less wear debris than metal-on-polythylene, Acta Orthop. Scand. 70 (6) (1999) 578–582. [38] A. Fischer, Well-founded selection of materials for improved wear resistance, in: K.C. Ludema (Ed.), Proceedings of the Conference on Wear of Materials ‘95, 4.1995, Boston, MA, USA, Wear 194 (1995) 238–245