Influence of fretting regimes on the tribocorrosion behaviour of Ti6Al4V in 0.9 wt.% sodium chloride solution

Wear 256 (2004) 963–972

Influence of fretting regimes on the tribocorrosion behaviour of Ti6Al4V in 0.9 wt.% sodium chloride solution S. Barril, S. Mischler∗ , D. Landolt Laboratoire de Métallurgie Chimique, Institut des Matériaux, Ecole Polytechnique Fédérale de Lausanne (EPFL), CH-1015 Lausanne, Switzerland

Abstract The effect of displacement amplitude, normal force and tribometer stiffness on the tribocorrosion response of a Ti6Al4V alloy in contact with a smooth alumina ball was investigated using a fretting corrosion rig equipped with an electrochemical cell. The roughness of the alloy plates was 0.6 ␮m, typical for the surface of hip joint stems. All experiments were carried out in 0.9 wt.% NaCl under anodic polarisation at a potential of 0.5 V versus the silver–silver chloride reference electrode, corresponding to the passive potential region. Anodic current, load, friction force and displacement amplitude were monitored in real time. Wear volume and surface morphology were evaluated at the end of fretting experiments. Results show that the prevailing fretting regime critically affects the overall system behaviour. Only fretting regimes involving slip led to measurable wear and to an enhancement of the anodic current. Previously developed tribocorrosion models for passivating metals were found to well describe the effect of normal force and displacement amplitude on the anodic current. © 2003 Elsevier B.V. All rights reserved. Keywords: Fretting regimes; Tribocorrosion; Titanium alloy; Stiffness

1. Introduction Total hip arthroplasty is a common surgical intervention. However, in certain cases post-operative loosening of the implants has been observed. The detachment of the anchoring of the femoral stem has been attributed to a rapid degradation of the interface between the stem and the bone cement due to micro-movements at the bone–implant interface and to fretting corrosion. Windler and Klabunde [1] reported that relative displacements up to 30 ␮m do not disturb bone ingrowth at the interface between the implant and the osseous tissue. However, in case of relative movement above 150 ␮m, fibrous tissue appears between the bone and the metallic surface and weakens the interface between bone and stem surface [2]. Kovacs et al. [3] suggested that fretting corrosion may occur at the interface between the femoral stem and the bone. Release of metallic ions and wear fragments by fretting corrosion lead to allergy and to tissues inflammation in the vicinity of the stem. This lowers the adhesion at the interface and eventually leads to loosening of the implant [4]. Material degradation in fretting is affected by the motion amplitude and the corresponding fretting regime. Vingsbo and Söderberg [5] first described fretting regimes. At low ∗ Corresponding author. Tel.: +41-21-693-2954; fax: +41-21-693-3946. E-mail address: [email protected].

0043-1648/$ – see front matter © 2003 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2003.11.003

amplitude (stick regime) the elastic deformation of the contact system accommodates the relative displacement between the two counter parts and only very limited surface damage occurs. At higher amplitudes (gross slip regime), slip between the counterparts and plastic deformation result in severe surface damage by wear. At intermediate amplitudes (partial slip regime) partial slip occurs leading to slight wear usually associated with a fatigue mechanism. The observation of transients of tangential force versus relative displacement of the two counter parts permits one to identify the prevailing fretting regime in the contact: under stick conditions the transient exhibits a closed loop. Under gross slip conditions, sliding is clearly identified by a plateau of the tangential force in the transient. Under partial slip conditions no plateau is observed, but the transient loop exhibits an elliptic shape attributed to crack nucleation and growth underneath the surface [6]. The fretting corrosion of titanium and its alloys has been investigated previously in aqueous solutions simulating body fluids [7]. Generally a synergistic effect between wear and corrosion was found to increase metal degradation rate. The effect of amplitude and fretting regime on the overall degradation of Ti alloys in aqueous electrolytes is still unexplored, however. The goal of the present paper is to experimentally evaluate the effect of fretting regimes on the tribocorrosion of a Ti6Al4V alloy in a saline so-

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lution. The investigation was carried out using a fretting rig developed for the purpose [8]. Through the use of an electrochemical cell with a potentiostatic set-up, this apparatus allows one to impose well-defined potentials to the metal under investigation. Further, the corrosion rate can be determined by measuring the current flowing through the metal necessary to maintain the applied potential. The continuous measurement of the frictional force and of the relative displacement allows the establishment of friction loops (i.e. plots of the frictional force as a function of the displacement) for the identification of the fretting regime. The amplitude of the displacement as well as the load were varied to establish different regimes. The investigated model system consisted in a Ti6Al4V alloy flat held at a passive potential and loaded against an alumina ball (chosen for its chemical and mechanical inertness) in a 0.9 wt.% NaCl solution. Experiments were carried out under two different values of the contact stiffness along the vertical axis. Secondary electron spectroscopy, optical microscopy and laser profilometry were used to characterise the wear morphology for different experimental conditions.

2. Experimental Fretting corrosion experiments were carried out using the apparatus described earlier [8]. The contact configuration consisted in a stationary alumina ball (SWIP AG Brügg, 5 mm radius, G10 AFBMA finish) rubbing against an underlying Ti6Al4V (Siber&Hegner, Zürich, 4 mm radius) flat sample. The contact was immersed in a 0.9 wt.% NaCl electrolyte and its potential was controlled using a three-electrode set-up including a platinum coil (inert counter electrode), a reference electrode (commercial Ag/AgCl electrode) and the working electrode (Ti6Al4V sample). For all the experiments the applied potential was +0.5 V with respect to the Ag/AgCl reference electrode (all potentials given here are referred to this reference electrode). The Ti6Al4V samples were ground according to a well-defined procedure to achieve reproducible surface finish that is essential to obtain good reproducibility [9]. The sample roughness was 0.6 ␮m Ra to reproduce the typical roughness range of hip joint stems. The motion of the flat sample was achieved through a piezoelectric actuator driven by a triangular waveform signal with an oscillation frequency of 1 Hz. Experiments were performed at room temperature and lasted 3600 s. In the first set of experiments, the normal load was fixed to 10 N (910 MPa maximum Hertzian stress) and the stiffness in the vertical direction to 9.81 N mm−1 . The displacement amplitude varied between 3, 10, 30, 50 and 100 ␮m. A second series of experiments was carried out at different displacement amplitudes (10, 30, 50 and 100 ␮m) under a normal force of 30 N (1314 MPa maximum Hertzian stress). Finally, under an applied normal force of 10 N, an other set

of experiments was carried out at different displacement amplitudes of 10, 30, 50 and 100 ␮m with a vertical stiffness in the contact of 98.1 N mm−1 . After each experiment the samples were extracted from the cell and rinsed with distilled water. The wear volume was estimated by means of non-contact laser profilometry (UBM, Telefokus UBC 14). The dedicated software routine was used to calculate the wear volume from the acquired surface scans by 80 line scans per mm in the Y-direction. Each line scan had a point density of 1000 points per mm in the X-direction. Surface scans were typically performed on squares of 1.5 mm per 1.5 mm. SEM observations of worn surfaces were carried out on different samples with a Philips XL 30 scanning electron microscope. The working parameters were an accelerating voltage of 10 kV and a working distance of 10 mm. Moreover, in order to estimate the plastic deformation of the bulk material, cross-sections of the wear trace were studied. After an experiment performed at a displacement amplitude of 100 ␮m, a disc of approximately 1 mm thick was cut from the cylindrical sample using a diamond wire saw (Well-type 3032-4). This disc was again cut along its diameter close to the wear trace using the same wire saw. The side of the cut sample was polished with diamond paste up to grade 1 ␮m. Before observation with a JEOL 6300F scanning microscope, the material was chemically etched (2% HF, 10% HNO3 solution) in order to reveal the alloy microstructure. 3. Results 3.1. Anodic current Fig. 1 displays the evolution with time of the current averaged over a time interval of 1 s (1 oscillation cycle). Current decreases from 0 to 1800 s thus showing the progressive passivation of the metallic sample due to the application of a passive potential through the potentiostatic set-up. At 1800 s, fretting starts and the current increases abruptly. At the end of fretting the current decreases again. The enhancement of anodic current by rubbing is commonly observed in tribocorrosion of passive metals, where abrasion of the passive film covering the alloy leads to the exposure to the solution of metal that undergoes anodic oxidation until the passive film forms again. From these data, the current enhancement due to fretting Ir can be extrapolated by subtracting the base current calculated with values before and after fretting starts to the average value during fretting. Figs. 2 and 3 show the evolution of the current Ir with the displacement amplitude for different applied loads and tribometer stiffness values. The displacement amplitude clearly affects the anodic current Ir . Indeed, at amplitudes of 3 and 10 ␮m no increase in anodic current is observed under an applied force of 10 N, whereas, above 30 ␮m the Ir current increases with increasing displacement amplitude. Under a load of 30 N significant increase of current due to fretting is

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Fig. 1. Evolution of mean current during a fretting corrosion experiment. Test conditions are: normal force 10 N; displacement amplitude 100 ␮m; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution. The current enhancement Ir due to fretting is equal to the difference between the measured current and the base line current.

observed only at higher displacement amplitudes, i.e. 50 and 100 ␮m. At higher displacement amplitudes the mechanical stiffness significantly affects the anodic current (Fig. 3): indeed Ir is lower at higher stiffness. 3.2. Fretting regimes Fig. 4 illustrates the transient values of displacement, tangential force and current measured during a single back and forth cycle. The displacement corresponds well with the triangular motion imposed by the waveform generator. The tangential force assumes fairly constant values of approximately +5 or −5 N depending on motion direction. The presence of a plateau indicates that the relative displacement imposed to the contact effectively results in relative sliding between the two counter bodies. However, the inversion of

the tangential force is not instantaneous with the change in direction. At the start of each stroke, the slope in the tangential force transient reflects the elastic deformation of the samples and of the tribometer to accommodate the relative displacement between the two counter bodies. The current exhibits a plateau between 2 and 2.5 ␮A interrupted in one cycle by two marked minima below 1 ␮A corresponding to each change in direction. During elastic accommodation of the imposed displacement the two counterparts are relatively at rest and passive film removal stops. During this time the previously depassivated metal surface has time to repassivate and therefore the current drops. The prevailing fretting regime is expected to play an important role for the electrochemical response. Fretting regimes can be identified by plotting the transient values of the tangential force as a function of the displacement during a single oscillation cycle. Fig. 5a shows force–displacement

4

4

10 N

3.5

Current from the wear trace / µA

Current from the wear trace / µA

30 N

3 2.5 2 1.5 1 0.5 0 0

20

40

60

80

100

120

Displacement amplitude / µm Fig. 2. Current Ir at different displacement amplitudes and under two different normal forces. Test conditions are: normal force 10 and 30 N; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

9.81 N.mm -1 98.1 N.mm -1

3.5 3 2.5 2 1.5 1 0.5 0

0

20

40

60

80

100

120

Displacement amplitude / µm Fig. 3. Current Ir at different displacement amplitudes and under two vertical contact stiffness of 9.81 and 98.1 N mm−1 . Test conditions are: normal force 10 N; applied potential +0.5 V Ag/AgCl; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

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(a)

(b)

(c)

Fig. 4. Transients of lateral position (a), tangential force (b) and current (c). Data are taken 30 min after the onset on fretting. Test conditions are: normal force 10 N; amplitude 100 ␮m; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

transients for different amplitudes for tests carried out at 10 N normal force. At 3 ␮m, a closed fretting loop is observed. It illustrates that the stick regime prevails in the contact where elastic deformation of the two contacting bodies accommodates most of the relative motion. For an amplitude of 10 ␮m the loop presents a slightly elliptic shape indicating that partial slip starts to appear in the contact and that cracks start to nucleate under the contact surface [6]. At 30 ␮m, the tangential force presents an almost constant value of approximately ±5 N during more than the half of the imposed displacement amplitude. Under these conditions the elastic deformation of the alloy sample and of the apparatus can no longer fully accommodate the imposed displacement and an effective relative motion takes place in the contact between the two counter parts. Friction occurs between the two contacting surfaces and yields the constant value of the tangential force. For displacement amplitudes above 30 ␮m the contact is in the gross slip regime. At 50 and 100 ␮m slip dominates in the contact as shown by the extended plateau observed in the transients. Fig. 5b shows the relevant effect of the normal force on the distribution of the fretting regime prevailing in the contact. At 30 ␮m, no plateau is visible in Fig. 5b although it was clearly observed under a normal force of 10 N (Fig. 5a). Modification of the

Fig. 5. Tangential force transients for different displacement amplitudes taken 30 min after the onset on fretting. Test conditions are: applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution—(a) 10 N load; (b) 30 N load.

normal force influences the respective elastic and plastic accommodation of the contact and thus the slip distance in the contact. Force–displacement transients measured under an applied vertical stiffness of 98.1 N mm−1 are almost identical to those measured at lower vertical stiffness. Apparently, in this range of values, the vertical stiffness has no noticeable influence on the prevailing fretting regime. All the transients shown in Fig. 3 were taken at a precise instant of the experiment (i.e. at 1800 cycles). No significant change in transient shape could be observed during the test. In case of slip regime, the plateau value of the tangential force corresponds to the frictional force and, after normalisation by the normal force, to the coefficient of friction µ. Independently on displacement amplitude, normal force or stiffness, all measured µ values were in the range of 0.46–0.52.

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Fig. 6. SEM micrographs showing wear morphology after 3600 s fretting at a displacement amplitude of 50 ␮m and under an applied potential of +0.5 V Ag/AgCl; duration 3600 s in 0.9 wt.% NaCl: (a) stiffness 9.81 N mm−1 , 10 N load; (b) stiffness 98.1 N mm−1 , 10 N load; (c) stiffness 9.81 N mm−1 , 30 N load.

3.3. Wear Fig. 6 shows SEM pictures of worn surfaces after 3600 cycles of fretting under different load and stiffness values. Under all conditions, two zones of the wear trace can be identified. The central part is characterised by relative severe material damage and the presence of debris particles. The elliptical dimension corresponds well with the dimensions

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of the Hertzian contact area summed, for the axis parallel to the motion direction, with the effective displacement [8]. Accordingly, the dimensions of the central part increase with increasing load. The surrounding external area of the wear trace is smoother and exhibits large material smearing. The stiffness does not seem to significantly affect wear morphology, no significant differences appearing between Fig. 6a and b. At equivalent displacement amplitude, the smeared area is smaller at higher load (Fig. 6c) because the effective slip amplitude is lower. To understand the origin of this particular wear morphology, SEM investigations were carried out on samples rubbed for 1, 50, 100 cycles (Fig. 7) under a normal load of 10 N and displacement amplitude of 50 ␮m. At the beginning damage is characterised by an increasing crushing and smearing of asperities left by grinding operations. After 50 cycles most of the surface concerned by fretting was covered with smeared material. Profilometry analysis did not show any evidence of material detachment from the titanium sample but only smoothening of the initial profile. At 100 cycles the central zone appeared for the first time in the SEM images and profilometry shows that some material detachment has occurred over the entire fretted area. After 3600 cycles the smeared area is considerably extended because of increasing ball penetration due to wear of titanium. Fig. 8 displays SEM observations of worn surfaces after fretting by one hour at displacement amplitudes of 10, 30, 50, 100 ␮m (magnification 200× and 500×). Strong differences in the wear morphology are observed with increasing amplitude. At 10 ␮m, the contact surface exhibits little plastic deformation after 1 h of fretting. This is related to the little relative displacement in the contact because of a stick regime. These deformed zones are likely to be upper parts of contacting asperities plastically deformed by the pressure field. At 30, 50 and 100 ␮m the surfaces exhibit typical structures with an internal zone of severe damage and an external ring of spread material. Fig. 9 shows a cross-sectional view of the wear trace after chemical etching. The sample is tilted by 27◦ to show in the upper part the wear track surface. The alloy microstructure is clearly visible with isolated ␤-grains embedded in the ␣-phase. This structure is homogeneous and corresponds well to the structure found in the bulk alloy at a depth of 200 ␮m below the surface. The grey layer close to the surface was also observed in other parts of the sample where fretting did not occur and is probably related to an enhanced etching at the sample edge. Thus, under the present conditions, no permanent plastic deformation or transformation occurred during fretting in the Ti6Al4V alloy. In fretting a so-called tribologically transformed structure (TTS) often forms below the surface after the accumulation of a critical amount of deformation [10]. However, the present experiments were carried out for shorter durations and at lower loads so that accumulation of damage was probably not sufficient to transform the sub-surface region. Rather, deformation occurred only at the very surface as evidenced by the

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Fig. 7. SEM wear morphologies (left column) after different number of fretting cycles together with corresponding topographical profile across the wear trace (right column): (a) after 1 cycle; (b) after 10 cycles; (c) after 50 cycles; (d) after 100 cycles; (d) after 3600 cycles. Test conditions are: normal force 10 N; amplitude 50 ␮m; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

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spread material observed on the wear track surface. Indeed, the coefficient of friction being greater than 0.3 contact mechanics theory [11] predicts that the maximum shear stress moves from below the surface (as predicted by Hertz’s theory in absence of tangential force) to the surface. No significant wear or damage of the alumina ball could be detected by three-dimensional laser profilometry or optical microscopy. 3.4. Wear volumes Fig. 10 shows the normal load influence on wear volumes measured with three-dimensional non-contact laser profilometry versus displacement amplitude. At amplitudes up to 30 ␮m the measured wear volumes are inferior to 0.1×10−3 mm3 for both 10 and 30 N. Wear volume remains low probably because stick conditions prevail for both normal forces. At displacement amplitude of 50 ␮m, wear volume increases between 0.1 × 10−3 and 0.2 × 10−3 mm3 for an applied normal force of 10 N. However, wear volumes for 30 N at this amplitude are still inferior to 0.1 × 10−3 mm3 . This confirms that an increase in normal force affects the prevailing fretting regime and thus the tribological conditions in the contact. At a higher displacement amplitude of 100 ␮m, wear volumes are more important, between 0.28 and 0.38 × 10−3 mm3 for 10 N and between 0.46 and 0.54 × 10−3 mm3 for 30 N. Experiments performed at 9.81 N mm−1 and at a 10-fold higher stiffness resulted in comparable wear volumes. Thus, the extent of wear seems not to be affected by contact stiffness.

4. Discussion

Fig. 8. SEM micrographs showing the effect of displacement amplitude on wear morphology: (a) 10 ␮m; (b) 30 ␮m; (c) 50 ␮m; (d) 100 ␮m. Test conditions are: normal force 10 N; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

The obtained results indicate that a significant electrochemical response of the Ti6Al4V is observed only when appreciable slip occurs in the contact. Indeed, Figs. 2 and 3 indicate a threshold value of the displacement below which current increase is negligible. This threshold value corresponds to the minimum amplitude required to achieve slip as shown by transients shown in Fig. 5. Further, on the transient scale, one observes a current decrease at the stroke edges, when displacement is accommodated by elastic deformation and slip is negligible (Fig. 4). The increase in current due to rubbing of passive metals is generally attributed to the exposure of native metal to the solution. Because the applied potential lies in the passive range, the metal oxidises until the passive film is formed again. Metal can be either oxidised in form of dissolved ions or in form of solid oxide forming the passive film. Repeated depassivation–repassivation cycles in different contact spots yield the observed increase in current. Several mechanisms can yield depassivation: plastic deformation of asperities or metal particles, detachment of metal particles, spalling off of the passive film. In the present case plastic deformation likely determines the current during rubbing. In effect, the

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Fig. 9. Cross-section of the material approximately in the middle of the wear trace. Tilt of the sample: 27◦ . Test conditions are: normal force 10 N; amplitude 100 ␮m; applied potential +0.5 V Ag/AgCl; stiffness 9.81 N mm−1 ; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.

initial current peak (Fig. 1) corresponds well with the period (>100 cycles) when smearing of the asperities but no wear is observed (Fig. 7). Models, based on asperity junction and plastic behaviour, were previously developed to describe the anodic current in sliding tribocorrosion of passive metals [12]. For a smooth hard indenter sliding against a soft passive metal the measured current Ir is given by

0.6

Wear volume / .10

-3

mm3

10 N

0.5

30 N

0.4

0.3

0.2

0.1

0

0

20

40

60

80

100

120

Displacement amplitude / µm Fig. 10. Measured wear volumes under normal forces of 10 and 30 N for different amplitudes. Test conditions are: stiffness 9.81 N mm−1 ; applied potential +0.5 V Ag/AgCl; duration 3600 s; oscillation frequency 1 Hz in 0.9 wt.% NaCl solution.



 Fn 1 Ir = k vs Qp λ H

(1)

where k is the probability factor taking into account that not all the asperity contacts lead to depassivations, λ the length of the smooth body in the sliding direction, vs the sliding velocity, Fn the normal force, H the metal hardness and Qp the passivation charge density, i.e. electrochemical charge per surface area leading to the reformation of the original passive film on the exposed metal surface. For a ball on flat contact, the dimension λ can be defined by the radius rH of the Hertzian contact (in the present conditions 72 and 105 ␮m for 10 and 30 N normal force, respectively). Thus, according to (1) the current Ir should be proportional to the factor (1/rH )vs Fn . The sliding velocity is determined by dividing the effective slip distance (the width of the force–displacement hysteresis shown in Fig. 3) by the stroke duration (0.5 s at 1 Hz). Fig. 11 displays the current values as a function of the factor (1/rH )vs Fn measured for two different normal forces and two stiffness values at identical load. For a stiffness of 9.81 N mm−1 all data points lie on a straight line independently on the applied load. This shows that Eq. (1) correctly describes the current resulting from fretting under slip conditions. However, the model does not explain the effect of the stiffness on the current. The SEM images suggest that less smearing and thus less depassivation occurred at higher stiffness and therefore, according to the model, less current is expected. Although this prediction is supported qualitatively by the experimental observations, the exact mechanism by which the stiffness affects the deformation behaviour of the

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4.0

Stiffness 9.81 N/mm Stiffness 98.1 N/mm

Ir [mA]

3.0

2.0

1.0

0.0 0

10

20

30

40

v s F n / λ [N/s] Fig. 11. Plot of the current Ir vs. the factor (1/rH )vs Fn (rH radius of the Hertzian area of contact, vs sliding velocity, Fn normal load). Data for stiffness of 9.81 N mm−2 include results obtained under a load of 10 or 30 N.

metal is not clear. Possibly, the stiffness imposed between the two counter parts affects particles ability to be ejected from the contact. At low stiffness, the slider encountering a particle can more easily move up and thus trap and deform the particle within in the contact. In a stiff contact the particle might just be pushed out of the contact by the slider. The probability of the particle to generate anodic current due to plastic deformation should therefore be higher at lower stiffness. Plastic deformation of asperities was observed even in situations dominated by the stick regime (Fig. 8). However, in such conditions, no current enhancement occurred probably because the depassivated areas were not accessible by the solution and thus could not repassivate. The model mentioned above considers that the displacement of the Hertzian contact area determines alone the depassivation rate. However, the SEM images show that the external area of the wear trace is also subject to plastic deformation and, according to the surface profiles, contribute to the overall wear. Thus the external area is expected to contribute to the anodic current. However, the distinct degradation mechanisms observed in the two areas may differ in their depassivation efficiency, i.e. in the factor k. According to this, the fracture and particle detachment observed in the central part should be much more efficient in depassivating the metal than the crushing of particles and asperities (external area) and therefore determine the overall current. The Ir values can be converted in anodic volume, i.e. the equivalent metal volume that was oxidised during rubbing, by using Faraday’s equation as described in [8]. In Fig. 12, the wear volume was plotted against the anodic volume Van determined by assuming that the oxidation of titanium to Ti4+ is the only electrochemical reaction contributing to the current Ir . For simplification the oxidation of Al and V

was neglected. A linear relationship between wear and Van is observed. If one assumes that the total wear volume is the sum of anodic wear and mechanical wear and the latter is a constant the data of Fig. 12 should exhibit a slope of one. The measured value is clearly higher, however. Two interpretations of this observation are possible. On one hand, if one would assume a valence of three instead of four for the conversion of the current into anodic wear volume the slope in Fig. 12 would be exactly one. Indeed, Hanawa et al. [13] observed a significant amount of Ti3+ and Ti2+ in the surface film after scratching titanium in different simulated body fluids. Another explanation could be that a substantial part of the current results from oxidation of particles mechanically detached from the first body. If particle oxidation is not

Fig. 12. Plot of the wear trace volume as a function of the anodic volume Van (all data at different amplitudes, loads and stiffness values are represented).

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complete the total amount of material loss from the contact would be larger than that calculated from anodic oxidation only. At present it is not possible to distinguish between the two hypotheses.

5. Conclusions The fretting corrosion behaviour of a rough Ti6Al4V alloy flat in contact with an alumina ball immersed in 0.9 wt.% NaCl was investigated under an applied passive potential of 0.5 V SSE using a dedicated triboelectrochemical fretting rig. Different stick and slip regimes (stick, mixed, slip) could be established by applying different displacement amplitude and loads. The obtained results clearly show that the electrochemical response of the alloy to tribocorrosion is critically affected by the prevailing fretting regime. In particular the following conclusions can be drawn: • Significant wear and enhancement of the anodic current occurred only in presence of slip. • The enhancement of the anodic current can be described by existing models developed for sliding contacts and based on cyclic depassivation–repassivation of the metal surface. • The stiffness of the tribometer in the load direction was found to affect the anodic current but not overall wear. The difference in anodic behaviour was attributed to the behaviour of third body particles.

References [1] M. Windler, R. Klabunde, in: D.M. Brunette, et al. (Ed.), Titanium in Medicine, vol. 1, Springer-Verlag, Berlin, 2001, pp. 703– 746. [2] R. Pilliar, D. Deporter, P. Watson, in: P. Vincenzini (Ed.), Materials in Clinical Applications, Techna Srl., 1995. [3] P. Kovacs, J.A. Davidson, K. Daigle, in: K.R. St. John (Ed.), Particulate Debris from Medical Implants: Mechanisms of Formation and Biological Consequences, American Society for Testing and Materials, ASTM STP 1144, Philadelphia, 1992, pp. 160– 176. [4] J. Fisher, in: B. Bhushan (Ed.), Modern Tribology Handbook, vol. 2, CRC Press, New York, 2001, pp. 1593–1609. [5] O. Vingsbo, S. Söderberg, Wear 126 (1988) 131–147. [6] G. Zambelli, L. Vincent, in: Matériaux et Contacts, PPUR, Lausanne, 1998, pp. 285–298. [7] R.B. Waterhouse, in: ASM Handbook: Friction, Lubrication and Wear Technology, vol. 18, ASM International, 1992, pp. 242– 256. [8] S. Barril, N. Debaud, S. Mischler, D. Landolt, Wear 252 (2002) 744–754. [9] S. Barril, Fretting corrosion of Ti6Al4V: contribution to the in-vitro simulation of the femoral-bone cement interface, EPFL Thesis, 2003, p. 2751 [10] E. Sauger, S. Fouvry, L. Ponsonnet, P. Kapsa, J.M. Martin, L. Vincent, Wear 245 (2000) 39–52. [11] K.L. Johnson, Wear 190 (1995) 162–170. [12] D. Landolt, S. Mischler, M. Stemp, Electrochim. Acta 46 (2001) 3913–3929. [13] T. Hanawa, K. Asami, K. Asaoka, J. Biomed. Mater. Res. 40 (1998) 530–538.