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Effect of Corrosion on the Wear Behavior of Passivating Metals in Aqueous Solutions
Thin Films in Tribology / D. Dowsm et al. (Editors) 1993 Elsevier Science Publishers B.V.
245
Effect of Corrosion on the Wear Behavior of Passivating Metals in Aqueous Solutions S. Mischler, E.A. Rosset, D. Landolt Materials Department, &ole Polytechnique Fkdkrale d e Lausanne, 1015Lausanne, Switzerland
The corrosive wear behavior under sliding condition of a Fe25Cr alloy - a model material for stainless steel - in 1M H,S04 has been studied by using electrochemical techniques. An attempt is made to correlate quantitatively corrosion parameters with material loss under corrosive wear conditions. Surface characterisation has been done by Auger Electron Spectroscopy (AES) and Atomic Force Microscopy (AFM)
1. INTRODUCTION Corrosion products play a very important technological role. For example stainless steels and titanium alloys owe their good corrosion resistance to the presence of a thin surface oxide film (passive film) formed by reaction of the metal with the corrosive medium. In many practical applications stainless steels are subject to combined degradation by aqueous corrosion and wear and it is therefore of interest to understand the behaviour of passive films under corrosive wear (tribocorrosion) conditions. The present study was initiated with the aim to elucidate the effect of passive films on the wear behaviour of a Fe-25% Cr alloy in 1 M sulphuric acid. Wear tests were carried out on samples polarised at a potential of 700 m V corresponding to the passive potential region (all potentials i n this paper are given with respect to the normal hydrogen reference electrode). The wear rate, the friction coefficient and the current were recorded during the wear test. The morphology and the composition of the worn surfaces were characterised by Atomic Force Microscopy (AFM) and Auger Electron Spectroscopy (AES). For reference, a wear test was carried out on a sample polarised cathodically (-800mV) where no
corrosion takes place. The passivation kinetics of the Fe25Cr alloy polarised at 700 mV was determinedelectrochemicallywith the aim to find a correlation with the current measured during the wear test.
2. ELECTROCHEMICAL ASPECTS I n acidic solutions one can describe the corrosion of a metal M by the following electrochemical reactions involving oxidation of
Figure 1. Polarisation curve of Fe25Cr in 1M H P 4
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the metal M and reduction of the protons via transfer of n electrons per metal atom oxidised.
oxidation
M=Mn++ne-
(1)
reduction
nH+ + ne- = n/2 H,
(2)
Electrochemical techniques are very helpful in studying corrosion mechanisms because they offer the possibility to control the chemical nature of the surface exposed to an electrolyte and to measure in-situ the material loss rate associated with corrosion. This possibilities are best illustrated by the polarisation curves. To establish a polarisation curve in addition to the electrode being studied (worhng electrode) a reference electrode and an inert auxiliary electrode are employed. A potentiostat automatically maintains the desired potential between the working and reference electrodes by passing an appropriate current between the working and counter electrode. Polarisation curves are established in the potentiostatic mode by plotting the measured current versus the applied potential. The polarisation curve of the Fe25Cr alloy in 1M H,SO,is illustrated in Fig. 1. Two regions can be distinguished: the cathodic region were the current is negative and the anodic region where the current is positive. At the corrosion potential (-200mV in Fig. 1) the current is zero and the reactions (1) and (2) occur at the same rate. In the cathodic region, at potentials well below the corrosion potential (about lOOmV), negligible metal dissolution occurs and the current is determined by the kinetics of reaction (2). In the same way in the anodic region well above the corrosion potential the reaction (2) becomes negligible and the current is determined by the metal oxidation kinetics (reaction 1). The relation between weight loss of the metal and current in the anodic region is given by Faraday's law:
-dm =-
dt
IM, nF
(3)
where m is the metal mass dissolved during the time t, I is the current, F is the Faraday's constant (96500 Clmole) and M, the atomic mass of the metal (55 glmole for Fe25Cr)
The anodic region is divided in three parts corresponding to three different surface states of the metal. In the active region (-200 to about 0 mV) the metal dissolves directly in contact with the solution. In the range 0 to 1100 mV (passive region) the surface is covered with an approximately 2 nm thick oxide film (mainly trivalent chromium oxide), which inhibits the metal dissomA/ lution and leads to very low currents ( crn,). Above 1100 mV (transpassive region) the current increases, the passive film being no more stable because of the oxidation of trivalent chromium oxide to soluble hexavalent chromium. For a more detailed discussion of the electrochemical techniques the reader is referred to the literature [ 11. 3. EXPERIMENTAL 3.1 Test Materials Sliding wear conditions were established by rubbing an alumina pin (chosen for its chemical inertness) against a high purity iron-chromium alloy (25%w Cr) plate. The hardness of the alloy was 160 HV30. The oscillating pins were prepared by machining the ends of 4 mm diameter alumina rods in the shape of truncated cones (120" included angle). The diameter of the flat end was 0.68 mm giving an apparent contact area of 0.36 mm2.The pin was oscillating at a frequency of 5 Hz. The stroke length of 3.6 mm corresponded to an average sliding velocity of 36 mmls. The applied normal load was 5 N resulting in a nominal contact pressure of 14 MPa. The metal samples were prepared by embedding 12x12 mm
247
Fe25Cr plates in Technovit mount. After polishing with 6 p m diamond paste and ultrasonic cleaning in ethanol the samples were mounted in the electrochemical cell of the wear test rig schematically presented in Fig. 2. The 1M sulphuric acid solution was prepared using doubly distilled water and analysis grade 96% sulphuric acid (Merck). The solutions were not deaerated. All tests were carried out at room temperature (20-23°C).
Figure 2. Top view of the electrochemical cell. 1 Working electrode. 2- Reference electrode. 3- Auxiliary electrodes. 4Moving arm with pin.
3.2 Electrochemical tests The polarisation curves were measured in the potentiodynamic mode by scanning the potential from -750 mV to l4OO mV at a scan rate of 5 mVIs. The passivation kinetics were determined by the potential step method: after 5 minutes of cathodic polarisation at -800 mV the potential was stepped to 700 mV. The resulting current transient was digitally recorded at a sampling rate of 1 kHz.
3.3. Wear Test The tribo-corrosion experiments were carried out in the reciprocating pin-on-plate rig described in more detail in [2]. Sliding wear conditions were established by rubbing the end of a vertically mounted pin against a fixed flat plate sample. The electrode potentials are applied using a HEKA PG285 potentiostat connected to a graphite counter electrode and a mercury sulphate reference electrode (Fig. 2). The electrochemical cell is located above a load cell allowing for the measurement of the normal force. The frictional force is measured with a piezoelectric force transducer. Reciprocal pin motion is provided by an electrodynamic vibration exciter (Bruel Kjaer 4809) driven by a triangular waveform signal. The pin motion was measured by a photoelectric sensor. During the wear test the frictional and the normal forces as well as the current were monitored using a Macintosh IIfx computer (Labview2 software from National Instruments) equipped with a National Instruments general purpose 16 bit IIO board.
The test conditions involved cathodic polarisation at -800 mV for 5 min followed by passivation at 700 mV. After 25 minutes passivation the rubbing was started. At the end of the test the metal samples and the alumina pins were removed from the solution, rinsed with distilled water and dried with an argon jet. The surface morphology resulting from tribocorrosion was characterised by optical microscopy and Atomic Force Microscopy (Park Scientific Instruments, SFM-BD2). Surface analysis and secondary electron imaging were carried out in a Perkin Elmer 660 Scanning Auger Microscope using a 3 keV (25 nA) electron beam. Depth profile acquisition was performed by rastering a 2 keV Ar+ beam over an
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area of 1.5x 1.5mm. Under these conditions the sputter rate of Ta,O, was 0.30 nmlmin. 4.
RESULTS
Figure 3 shows the evolution of the current with the time after switching the sample potential from the cathodic region to the passive region. At the beginning of the passivation the current is very high but decreases quickly as soon as the passive film is formed. The experimental curve plotted in Fig. 3 can be described by the empirical relation (4)
i = 103.exp (4.139.t) + 175.t 4-614 where t is expressed in ms and i in mA/cm2. The curve calculated using (4) is plotted in Fig. 3 together with the experimental points. The obtained friction and material loss parameters are listed in Table l. The tabulated coefficients of friction p correspond to the average values recorded during the rubbing. Variations in p were less than 10%. A linear decrease of the vertical pin position with sliding time was observed after a running-in period of 60 s. Since no pin geometry change was observed after the wear test, the measured decrease can be attributed to wear of the metal. The
i7 E
9E
!i
Y
Im
E 2 5
0 0
50
rm
150
203
250
I 300
Time [ms]
Figure 3. Experimental and fitted passivation current transients.
wear rates listed in table 1 correspond to the slope of the curve indicating the vertical pin position versus time. The average wear rate measured on passivated surfaces is 7.3 nmls. T h e start o f rubbing o n passivated samples corresponds t o a n increase of the current from lO~’mA to 0.65 mA. After the running-in period the current remained within a scatter band of 5%.The rubbing did not affect the current of the sample polarised in the cathodic region. Current densities were calculated by dividing the current by the wear scar surface. It is assumed that the current is due only to the dissolution occumng in the wear scar. The influence on the measured current of oxygen reduction is considered to be negligible. In order to calculate the wear scar surface a rectangular geometry of the wear scar was assumed. Length and width of the scar were determined by optical microscopy after the test. The averaged wall area was estimated by taking half the depth measured after sliding. The current densities were converted into material loss using (3). This yielded an average dissolution rate of (14.3in) nmls. The variation of current and of pin horizontal position recorded during a stroke are shown in Fig. 4. At the end of each stroke, when the direction of motion changes, the pin remains motionless for about 25 ms. This is because the static frictional force exceeds the force transmitted by the exciter to the pin in order to generate the alternate motion. During the dead time the current drops to about 25% of its maximum value. A low magnification image of the wear track obtained on a passivated sample is shown in Fig. 5. The geometry of the wear track exhibits a regular and well defined shape corresponding to the pin geometry. More detailed three dimensional images of the bottom of the wear tracks are obtained by
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Table I Wear tests results Applied Slidmg potential distance [VI [ml 0.7 I26 0.7 146 0.7 355
Sliding time [s1 3500 3840 10140
P
-.8
rn
225
using Atomic Force Microscopy (Fig. 6). Note that the relief in Figure 6 has been enhanced by using a vertical scale magnification factor of 50 with respect to the horizontal scale. Area roughness statistics can be extracted from AFM images by calculating the mean deviation from the average height. This value corresponds in principle to the R, value determined in conventional two dimensional profilometry. In this way we obtain roughness values of 61 and 39 nm for the wear track formed at cathodic potential and passive potential respectively. For comparison the roughness outside the wear track on a passivated sample corresponds to 28 nm.
.24 .24 .25
Total wear rate [nmls] 7.5 8.1 6.4
Current density [mA/cm2] 21 19 17
.22
1.5
--
ples during rubbing nor after filtering the solution or the water used for washing the samples. Some wear debris were found accumulating at the edges of the flat end of the pin however. Wear debris apparently did not play a significant role in the wear process.
Wear debris were found neither on the sam-
0
100
200
300
Time [ms]
Figure 4. Variation of current and horizontal pin position during sliding.
Figure 5. SEM image of the wear scar on passivated sample.
250
The grooves observed by AFM are oriented along the pin motion axis and are probably the result of the microploughing action of asperities present on the alumina surface. Abrasive wear is expected when one of the two moving surfaces is considerably harder than the other, as was the case in the present experiment. The geometry of the pin as observed by optical microscopy did not change during the experiment. Under the conditions of this study alumina, therefore, was not subject to corrosion or wear. The Auger depth profile measured on a area outside the wear track of a passivated sample is shown in Fig. 7. The oxygen profile exhibits a well defined oxide-metal interface at about 1.5
nm. Carbon is present as a surface contamination only, the carbon signal goes to zero at an apparent depth of 0.3 nm. Sulphur is found in the passive film probably as sulphate from the solution. A detailed surface analytical study of the composition of a passive film formed in similar conditions can be found in the literature [3]. The oxygen profile measured in the wear track (Fig. 8) does not show a sharp interface probably because the surface roughness induces shadowing effects of the ion beam used for sputtering. This implies that locally the passive oxide film is riot removed by sputtering and still contributes to the overall Auger signal. Ion beam shadowing effects render very difficult the interpretation of depth profiles and therefore no quantification of the profiles of Fig. 7 and 8 were camed out in the present study. It is however possible to conclude that the passive film formed in the wear track has about the same thickness as the one formed outside the track but contains less sulphur. 5) DISCUSSION Uhlig [4] has proposed a simple mechanistic model based on a chemical and a mechanical factor t o describe the weight loss of a metal surface undergoing fretting corrosion in a gaseous environment. Uhlig considered that an asperity rubbing on a metal surface produces a track of clean metal which immedately oxidizes and therefore leads to a chemically controlled surface damage. The mechanical factor correspond to a certain amount of wear caused by the asperities digging below the surface.
Figure 6. AFM images of the wear tracks formed at the cathodic potential (above) and at the passive potential (below).
In principle, the results of the present study can be interpreted by a similar model. Under sliding conditions the passive film is periodically removed from the contact area exposing the nascent metal surface to the acid solution. A considerable amount of dissolution is required in order to repassivate the exposed surface as shown by the current transient of Fig. 3.
251
hence the valence by (3.
n= v)
Lu
< 0
5
10 15 Sputter Time [min]
20
25
Figure 7. Auger depth profile measured outside the track on passivated sample.
0
5
10 15 Sputter Time [min]
20
25
Figure 8. Auger depth profile measured inside the wear track on passivated sample. Since at each stroke of the pin the passive film is removed and rebuilt the material loss due to corrosion is higher under sliding conditions than during static conditions. This can explain the increase of two orders of magnitude of the current observed on passivated samples when sliding starts. An increased corrosion rate compared to static conditions would also be expected if rubbing leads only to film thinning. To periodically restore the original film thickness metal must be oxidised. In this case the total material loss should correspond to the amount of metal oxidized and
ti
i n equation 3 can be obtained
W.L
MI l F p
Where wt is the total wear rate listed in table 1 and p the density (7.69 g/ml). Equation (3, for an average value of wt=2.3 nmls, yields a dissolution valence n= 1.96.This value is to low for the considered passive potential, where chromium is expected to dissolve at a valence of 3 and iron at valences ranging from 2 to 3. This suggest that passive film thinning is not the predominent mechanism of material removal. However, a more exact determination of the dissolution valence is required in order to confirm this conclusion. The presence of a passive film on the Fe25Cr surface can explain the smoother morphology of the wear track surfaces formed at the passive potential compared to the surfaces formed under cathodic conditions where no film is present. During the sliding, and at each stroke, the film covering the rough surface is removed together with some underlying metal from the contact areas, i.e. from the asperities. After the film removal the fresh metal surface exposed to the corrosive solution at the tips of the asperities undergoes dissolution until the passive film is rebuilt. On the non-contact areas, i.e. valleys, the passive film remains and thus no dissolution occurs. The sliding process thus causes a chemical selective material loss from the asperities which leads to smoothing. In absence of corrosion the relief is determined only by the ploughing action of the asperities on the alumina surface and no chemical induced smoothing occurs. Assuming the validity of the wear mechanism involving passive film removal from the surface with subsequent repassivation it should be possible to calculate the current transient of Fig. 4 on the basis of the passivation kinetic.
252
For this purpose the length of the wear track can be divided in N segments of equal length. The total current I measured as a function of the time t is then given by (6) N
where S is the wear scar area and t,(k) corresponds to the time when the segment k was last activated by rubbing. The activation time of each segment is determined by the displacement profile of the pin. Under the present conditions the motion of the pin can be divided into four steps involving a forward displacement lasting 75 ms, a dead time of 25 ms with no motion, a backwards displacement lasting again 75 ms and finally again a dead time of 2S ms. Thus for forward displacement the activation time of each segment k is defined by:
where L is the length of a segment and v is the displacement velocity (30 mmls i n our case). For the backwards displacement t,(k) is given by
tA(k) = 75 + 25 + (N-k+ 1) . Uv
[ms]
(8)
Although the conditions i n the wear scars may differ from the passivation conditions used to establish (4) this relation can describe the function i(t-t,) of (6) . In particular slightly higher temperatures due to friction and changes in solution composition due to corrosion may play a role.
The values calculated numerically using (6) with 75 segments are compared to experimental results in Fig. 9. A reasonable good agreement IS found i n the shape of the curves. The absolute value of the calculated current is however about 150 times higher than the measured current. This
can be interpreted by assuming than the real area of contact corresponds to 0.5 - 1 5% of the apparent area of contact. One may assume that the mechanical factor of corrosive wear of passivated samples corresponds to the total wear found on the sample polarised cathodically, where no corrosion occurs. In this case an average chemical wear rate of
I ---1O 8
0
0
50
Irn
150
200
: 1 250
0
300
Time [ms]
Figure 9. Calculated and measured current transients during sliding on passivated sample. 5.8 nm/s can be obtained by subtracting the wear rate of the cathodic sample (1.5 nmls) to the average wear rate of passivated samples (7.3 nmi s). Comparing this value with the average current density leads to a valence of dissolution of 2.47. This value is reasonable and implies that, assuming chromium dissolves at valence 3 . 3 0 % of the iron atoms dissolve with a valence of 3 and 70% with a valence 2. However, because of the presence of the passive film, it is not likely that mechanical wear at the passive potential corresponds to that at the cathodic potential. The observed differences in surface morphology and in the frictional coefficient should also influence the wear behaviour. If the valence of dissolution were known one could calculate exactly the chemical material removal rate by using Faraday’s law and therefore the mechanical wear rate by comparison with the
253
total wear rate. 6) CONCLUSIONS
A simple mechanistic model considering a mechanical and a chemical factor may be used to describe weight loss and surface morphology of Fe25Cr alloy surfaces sliding against alumina in 1M sulphuric acid under applied passive potential of 700mV. The current measured during sliding can be calculated by known displacement of the pin on the basis of the passivation kinetics measured under static conditions. Knowledge of the exact dissolution valence of the metal during sliding is required i n order to quantitatively interpretate corrosive wear measurements.
REFERENCES [ 11H.H. Uhlig and R.W. Revie Corrosion and Corrosion Control J. Wiley & Sons, New York (1985) [2] E.A. Rosset, S. Mischler, D. Landolt To be published i n Thin Films in Tribology Proceedings o f the 19th Leeds-Lyon Symposium on Tribology [3]S. Mischler, H.J. Mathieu and D. Landolt Surface and Interface Analysis 11,182 (1988) [4]H.H. Uhlig Journal of Applied Mechanics 21,401 (1954)
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Effect of Corrosion on the Wear Behavior of Passivating Metals in Aqueous Solutions S. Mischler, E.A. Rosset, D. Landolt Materials Department, &ole Polytechnique Fkdkrale d e Lausanne, 1015Lausanne, Switzerland
The corrosive wear behavior under sliding condition of a Fe25Cr alloy - a model material for stainless steel - in 1M H,S04 has been studied by using electrochemical techniques. An attempt is made to correlate quantitatively corrosion parameters with material loss under corrosive wear conditions. Surface characterisation has been done by Auger Electron Spectroscopy (AES) and Atomic Force Microscopy (AFM)
1. INTRODUCTION Corrosion products play a very important technological role. For example stainless steels and titanium alloys owe their good corrosion resistance to the presence of a thin surface oxide film (passive film) formed by reaction of the metal with the corrosive medium. In many practical applications stainless steels are subject to combined degradation by aqueous corrosion and wear and it is therefore of interest to understand the behaviour of passive films under corrosive wear (tribocorrosion) conditions. The present study was initiated with the aim to elucidate the effect of passive films on the wear behaviour of a Fe-25% Cr alloy in 1 M sulphuric acid. Wear tests were carried out on samples polarised at a potential of 700 m V corresponding to the passive potential region (all potentials i n this paper are given with respect to the normal hydrogen reference electrode). The wear rate, the friction coefficient and the current were recorded during the wear test. The morphology and the composition of the worn surfaces were characterised by Atomic Force Microscopy (AFM) and Auger Electron Spectroscopy (AES). For reference, a wear test was carried out on a sample polarised cathodically (-800mV) where no
corrosion takes place. The passivation kinetics of the Fe25Cr alloy polarised at 700 mV was determinedelectrochemicallywith the aim to find a correlation with the current measured during the wear test.
2. ELECTROCHEMICAL ASPECTS I n acidic solutions one can describe the corrosion of a metal M by the following electrochemical reactions involving oxidation of
Figure 1. Polarisation curve of Fe25Cr in 1M H P 4
246
the metal M and reduction of the protons via transfer of n electrons per metal atom oxidised.
oxidation
M=Mn++ne-
(1)
reduction
nH+ + ne- = n/2 H,
(2)
Electrochemical techniques are very helpful in studying corrosion mechanisms because they offer the possibility to control the chemical nature of the surface exposed to an electrolyte and to measure in-situ the material loss rate associated with corrosion. This possibilities are best illustrated by the polarisation curves. To establish a polarisation curve in addition to the electrode being studied (worhng electrode) a reference electrode and an inert auxiliary electrode are employed. A potentiostat automatically maintains the desired potential between the working and reference electrodes by passing an appropriate current between the working and counter electrode. Polarisation curves are established in the potentiostatic mode by plotting the measured current versus the applied potential. The polarisation curve of the Fe25Cr alloy in 1M H,SO,is illustrated in Fig. 1. Two regions can be distinguished: the cathodic region were the current is negative and the anodic region where the current is positive. At the corrosion potential (-200mV in Fig. 1) the current is zero and the reactions (1) and (2) occur at the same rate. In the cathodic region, at potentials well below the corrosion potential (about lOOmV), negligible metal dissolution occurs and the current is determined by the kinetics of reaction (2). In the same way in the anodic region well above the corrosion potential the reaction (2) becomes negligible and the current is determined by the metal oxidation kinetics (reaction 1). The relation between weight loss of the metal and current in the anodic region is given by Faraday's law:
-dm =-
dt
IM, nF
(3)
where m is the metal mass dissolved during the time t, I is the current, F is the Faraday's constant (96500 Clmole) and M, the atomic mass of the metal (55 glmole for Fe25Cr)
The anodic region is divided in three parts corresponding to three different surface states of the metal. In the active region (-200 to about 0 mV) the metal dissolves directly in contact with the solution. In the range 0 to 1100 mV (passive region) the surface is covered with an approximately 2 nm thick oxide film (mainly trivalent chromium oxide), which inhibits the metal dissomA/ lution and leads to very low currents ( crn,). Above 1100 mV (transpassive region) the current increases, the passive film being no more stable because of the oxidation of trivalent chromium oxide to soluble hexavalent chromium. For a more detailed discussion of the electrochemical techniques the reader is referred to the literature [ 11. 3. EXPERIMENTAL 3.1 Test Materials Sliding wear conditions were established by rubbing an alumina pin (chosen for its chemical inertness) against a high purity iron-chromium alloy (25%w Cr) plate. The hardness of the alloy was 160 HV30. The oscillating pins were prepared by machining the ends of 4 mm diameter alumina rods in the shape of truncated cones (120" included angle). The diameter of the flat end was 0.68 mm giving an apparent contact area of 0.36 mm2.The pin was oscillating at a frequency of 5 Hz. The stroke length of 3.6 mm corresponded to an average sliding velocity of 36 mmls. The applied normal load was 5 N resulting in a nominal contact pressure of 14 MPa. The metal samples were prepared by embedding 12x12 mm
247
Fe25Cr plates in Technovit mount. After polishing with 6 p m diamond paste and ultrasonic cleaning in ethanol the samples were mounted in the electrochemical cell of the wear test rig schematically presented in Fig. 2. The 1M sulphuric acid solution was prepared using doubly distilled water and analysis grade 96% sulphuric acid (Merck). The solutions were not deaerated. All tests were carried out at room temperature (20-23°C).
Figure 2. Top view of the electrochemical cell. 1 Working electrode. 2- Reference electrode. 3- Auxiliary electrodes. 4Moving arm with pin.
3.2 Electrochemical tests The polarisation curves were measured in the potentiodynamic mode by scanning the potential from -750 mV to l4OO mV at a scan rate of 5 mVIs. The passivation kinetics were determined by the potential step method: after 5 minutes of cathodic polarisation at -800 mV the potential was stepped to 700 mV. The resulting current transient was digitally recorded at a sampling rate of 1 kHz.
3.3. Wear Test The tribo-corrosion experiments were carried out in the reciprocating pin-on-plate rig described in more detail in [2]. Sliding wear conditions were established by rubbing the end of a vertically mounted pin against a fixed flat plate sample. The electrode potentials are applied using a HEKA PG285 potentiostat connected to a graphite counter electrode and a mercury sulphate reference electrode (Fig. 2). The electrochemical cell is located above a load cell allowing for the measurement of the normal force. The frictional force is measured with a piezoelectric force transducer. Reciprocal pin motion is provided by an electrodynamic vibration exciter (Bruel Kjaer 4809) driven by a triangular waveform signal. The pin motion was measured by a photoelectric sensor. During the wear test the frictional and the normal forces as well as the current were monitored using a Macintosh IIfx computer (Labview2 software from National Instruments) equipped with a National Instruments general purpose 16 bit IIO board.
The test conditions involved cathodic polarisation at -800 mV for 5 min followed by passivation at 700 mV. After 25 minutes passivation the rubbing was started. At the end of the test the metal samples and the alumina pins were removed from the solution, rinsed with distilled water and dried with an argon jet. The surface morphology resulting from tribocorrosion was characterised by optical microscopy and Atomic Force Microscopy (Park Scientific Instruments, SFM-BD2). Surface analysis and secondary electron imaging were carried out in a Perkin Elmer 660 Scanning Auger Microscope using a 3 keV (25 nA) electron beam. Depth profile acquisition was performed by rastering a 2 keV Ar+ beam over an
248
area of 1.5x 1.5mm. Under these conditions the sputter rate of Ta,O, was 0.30 nmlmin. 4.
RESULTS
Figure 3 shows the evolution of the current with the time after switching the sample potential from the cathodic region to the passive region. At the beginning of the passivation the current is very high but decreases quickly as soon as the passive film is formed. The experimental curve plotted in Fig. 3 can be described by the empirical relation (4)
i = 103.exp (4.139.t) + 175.t 4-614 where t is expressed in ms and i in mA/cm2. The curve calculated using (4) is plotted in Fig. 3 together with the experimental points. The obtained friction and material loss parameters are listed in Table l. The tabulated coefficients of friction p correspond to the average values recorded during the rubbing. Variations in p were less than 10%. A linear decrease of the vertical pin position with sliding time was observed after a running-in period of 60 s. Since no pin geometry change was observed after the wear test, the measured decrease can be attributed to wear of the metal. The
i7 E
9E
!i
Y
Im
E 2 5
0 0
50
rm
150
203
250
I 300
Time [ms]
Figure 3. Experimental and fitted passivation current transients.
wear rates listed in table 1 correspond to the slope of the curve indicating the vertical pin position versus time. The average wear rate measured on passivated surfaces is 7.3 nmls. T h e start o f rubbing o n passivated samples corresponds t o a n increase of the current from lO~’mA to 0.65 mA. After the running-in period the current remained within a scatter band of 5%.The rubbing did not affect the current of the sample polarised in the cathodic region. Current densities were calculated by dividing the current by the wear scar surface. It is assumed that the current is due only to the dissolution occumng in the wear scar. The influence on the measured current of oxygen reduction is considered to be negligible. In order to calculate the wear scar surface a rectangular geometry of the wear scar was assumed. Length and width of the scar were determined by optical microscopy after the test. The averaged wall area was estimated by taking half the depth measured after sliding. The current densities were converted into material loss using (3). This yielded an average dissolution rate of (14.3in) nmls. The variation of current and of pin horizontal position recorded during a stroke are shown in Fig. 4. At the end of each stroke, when the direction of motion changes, the pin remains motionless for about 25 ms. This is because the static frictional force exceeds the force transmitted by the exciter to the pin in order to generate the alternate motion. During the dead time the current drops to about 25% of its maximum value. A low magnification image of the wear track obtained on a passivated sample is shown in Fig. 5. The geometry of the wear track exhibits a regular and well defined shape corresponding to the pin geometry. More detailed three dimensional images of the bottom of the wear tracks are obtained by
249
Table I Wear tests results Applied Slidmg potential distance [VI [ml 0.7 I26 0.7 146 0.7 355
Sliding time [s1 3500 3840 10140
P
-.8
rn
225
using Atomic Force Microscopy (Fig. 6). Note that the relief in Figure 6 has been enhanced by using a vertical scale magnification factor of 50 with respect to the horizontal scale. Area roughness statistics can be extracted from AFM images by calculating the mean deviation from the average height. This value corresponds in principle to the R, value determined in conventional two dimensional profilometry. In this way we obtain roughness values of 61 and 39 nm for the wear track formed at cathodic potential and passive potential respectively. For comparison the roughness outside the wear track on a passivated sample corresponds to 28 nm.
.24 .24 .25
Total wear rate [nmls] 7.5 8.1 6.4
Current density [mA/cm2] 21 19 17
.22
1.5
--
ples during rubbing nor after filtering the solution or the water used for washing the samples. Some wear debris were found accumulating at the edges of the flat end of the pin however. Wear debris apparently did not play a significant role in the wear process.
Wear debris were found neither on the sam-
0
100
200
300
Time [ms]
Figure 4. Variation of current and horizontal pin position during sliding.
Figure 5. SEM image of the wear scar on passivated sample.
250
The grooves observed by AFM are oriented along the pin motion axis and are probably the result of the microploughing action of asperities present on the alumina surface. Abrasive wear is expected when one of the two moving surfaces is considerably harder than the other, as was the case in the present experiment. The geometry of the pin as observed by optical microscopy did not change during the experiment. Under the conditions of this study alumina, therefore, was not subject to corrosion or wear. The Auger depth profile measured on a area outside the wear track of a passivated sample is shown in Fig. 7. The oxygen profile exhibits a well defined oxide-metal interface at about 1.5
nm. Carbon is present as a surface contamination only, the carbon signal goes to zero at an apparent depth of 0.3 nm. Sulphur is found in the passive film probably as sulphate from the solution. A detailed surface analytical study of the composition of a passive film formed in similar conditions can be found in the literature [3]. The oxygen profile measured in the wear track (Fig. 8) does not show a sharp interface probably because the surface roughness induces shadowing effects of the ion beam used for sputtering. This implies that locally the passive oxide film is riot removed by sputtering and still contributes to the overall Auger signal. Ion beam shadowing effects render very difficult the interpretation of depth profiles and therefore no quantification of the profiles of Fig. 7 and 8 were camed out in the present study. It is however possible to conclude that the passive film formed in the wear track has about the same thickness as the one formed outside the track but contains less sulphur. 5) DISCUSSION Uhlig [4] has proposed a simple mechanistic model based on a chemical and a mechanical factor t o describe the weight loss of a metal surface undergoing fretting corrosion in a gaseous environment. Uhlig considered that an asperity rubbing on a metal surface produces a track of clean metal which immedately oxidizes and therefore leads to a chemically controlled surface damage. The mechanical factor correspond to a certain amount of wear caused by the asperities digging below the surface.
Figure 6. AFM images of the wear tracks formed at the cathodic potential (above) and at the passive potential (below).
In principle, the results of the present study can be interpreted by a similar model. Under sliding conditions the passive film is periodically removed from the contact area exposing the nascent metal surface to the acid solution. A considerable amount of dissolution is required in order to repassivate the exposed surface as shown by the current transient of Fig. 3.
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hence the valence by (3.
n= v)
Lu
< 0
5
10 15 Sputter Time [min]
20
25
Figure 7. Auger depth profile measured outside the track on passivated sample.
0
5
10 15 Sputter Time [min]
20
25
Figure 8. Auger depth profile measured inside the wear track on passivated sample. Since at each stroke of the pin the passive film is removed and rebuilt the material loss due to corrosion is higher under sliding conditions than during static conditions. This can explain the increase of two orders of magnitude of the current observed on passivated samples when sliding starts. An increased corrosion rate compared to static conditions would also be expected if rubbing leads only to film thinning. To periodically restore the original film thickness metal must be oxidised. In this case the total material loss should correspond to the amount of metal oxidized and
ti
i n equation 3 can be obtained
W.L
MI l F p
Where wt is the total wear rate listed in table 1 and p the density (7.69 g/ml). Equation (3, for an average value of wt=2.3 nmls, yields a dissolution valence n= 1.96.This value is to low for the considered passive potential, where chromium is expected to dissolve at a valence of 3 and iron at valences ranging from 2 to 3. This suggest that passive film thinning is not the predominent mechanism of material removal. However, a more exact determination of the dissolution valence is required in order to confirm this conclusion. The presence of a passive film on the Fe25Cr surface can explain the smoother morphology of the wear track surfaces formed at the passive potential compared to the surfaces formed under cathodic conditions where no film is present. During the sliding, and at each stroke, the film covering the rough surface is removed together with some underlying metal from the contact areas, i.e. from the asperities. After the film removal the fresh metal surface exposed to the corrosive solution at the tips of the asperities undergoes dissolution until the passive film is rebuilt. On the non-contact areas, i.e. valleys, the passive film remains and thus no dissolution occurs. The sliding process thus causes a chemical selective material loss from the asperities which leads to smoothing. In absence of corrosion the relief is determined only by the ploughing action of the asperities on the alumina surface and no chemical induced smoothing occurs. Assuming the validity of the wear mechanism involving passive film removal from the surface with subsequent repassivation it should be possible to calculate the current transient of Fig. 4 on the basis of the passivation kinetic.
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For this purpose the length of the wear track can be divided in N segments of equal length. The total current I measured as a function of the time t is then given by (6) N
where S is the wear scar area and t,(k) corresponds to the time when the segment k was last activated by rubbing. The activation time of each segment is determined by the displacement profile of the pin. Under the present conditions the motion of the pin can be divided into four steps involving a forward displacement lasting 75 ms, a dead time of 25 ms with no motion, a backwards displacement lasting again 75 ms and finally again a dead time of 2S ms. Thus for forward displacement the activation time of each segment k is defined by:
where L is the length of a segment and v is the displacement velocity (30 mmls i n our case). For the backwards displacement t,(k) is given by
tA(k) = 75 + 25 + (N-k+ 1) . Uv
[ms]
(8)
Although the conditions i n the wear scars may differ from the passivation conditions used to establish (4) this relation can describe the function i(t-t,) of (6) . In particular slightly higher temperatures due to friction and changes in solution composition due to corrosion may play a role.
The values calculated numerically using (6) with 75 segments are compared to experimental results in Fig. 9. A reasonable good agreement IS found i n the shape of the curves. The absolute value of the calculated current is however about 150 times higher than the measured current. This
can be interpreted by assuming than the real area of contact corresponds to 0.5 - 1 5% of the apparent area of contact. One may assume that the mechanical factor of corrosive wear of passivated samples corresponds to the total wear found on the sample polarised cathodically, where no corrosion occurs. In this case an average chemical wear rate of
I ---1O 8
0
0
50
Irn
150
200
: 1 250
0
300
Time [ms]
Figure 9. Calculated and measured current transients during sliding on passivated sample. 5.8 nm/s can be obtained by subtracting the wear rate of the cathodic sample (1.5 nmls) to the average wear rate of passivated samples (7.3 nmi s). Comparing this value with the average current density leads to a valence of dissolution of 2.47. This value is reasonable and implies that, assuming chromium dissolves at valence 3 . 3 0 % of the iron atoms dissolve with a valence of 3 and 70% with a valence 2. However, because of the presence of the passive film, it is not likely that mechanical wear at the passive potential corresponds to that at the cathodic potential. The observed differences in surface morphology and in the frictional coefficient should also influence the wear behaviour. If the valence of dissolution were known one could calculate exactly the chemical material removal rate by using Faraday’s law and therefore the mechanical wear rate by comparison with the
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total wear rate. 6) CONCLUSIONS
A simple mechanistic model considering a mechanical and a chemical factor may be used to describe weight loss and surface morphology of Fe25Cr alloy surfaces sliding against alumina in 1M sulphuric acid under applied passive potential of 700mV. The current measured during sliding can be calculated by known displacement of the pin on the basis of the passivation kinetics measured under static conditions. Knowledge of the exact dissolution valence of the metal during sliding is required i n order to quantitatively interpretate corrosive wear measurements.
REFERENCES [ 11H.H. Uhlig and R.W. Revie Corrosion and Corrosion Control J. Wiley & Sons, New York (1985) [2] E.A. Rosset, S. Mischler, D. Landolt To be published i n Thin Films in Tribology Proceedings o f the 19th Leeds-Lyon Symposium on Tribology [3]S. Mischler, H.J. Mathieu and D. Landolt Surface and Interface Analysis 11,182 (1988) [4]H.H. Uhlig Journal of Applied Mechanics 21,401 (1954)