- Email: [email protected]
Mapping sliding wear of steels in aqueous conditions
Wear 255 (2003) 456–465
Mapping sliding wear of steels in aqueous conditions M.M. Stack∗ , K. Chi Department of Mechanical Engineering, University of Strathclyde, James Weir Building, 75 Montrose St., Glasgow G1 1XJ, UK
Abstract In studies of sliding wear in dry environments, there have been many attempts at mapping the processes. For steels, the classical “wear map” has been developed and various other studies have extended the approach to composites, ceramics and coated materials. Despite this work, there have been few attempts to extend the methodology to wet conditions, where the wear process interacts with solutions that are defined by pH and electrochemical potential. The object of this work was to study the sliding wear–corrosion behaviour of steels in a pin-on-disc apparatus in aqueous conditions. The effects of applied load and velocity were evaluated at various electrochemical potentials in carbonate/bicarbonate solution pH 9.8). The results were analysed using weight loss and scanning electron microscopy techniques. Wear mechanisms were identified in the various environments and a method of identifying the wear–corrosion transitions, in aqueous conditions, was proposed in the work. These regimes were superimposed on wear–corrosion maps, where the change in wear–corrosion regime was identified as a function of velocity and electrochemical potential. Possible reasons for the differences in the boundaries of the map at various applied loads are discussed in this paper. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Sliding wear corrosion; Maps; Wear mechanisms; Aqueous environments
1. Introduction Significant advances in recent years have been made in the study of tribo-corrosion in aqueous conditions. In particular, there has been much work carried out in the field of erosion–corrosion by solid particles [1–5]. Regimes of interaction have been identified, ranging from “erosion-dominated” where the material loss is predominately from the substrate to “corrosion-dominated” where the losses are as a result of dissolution or formation of corrosion product [6–8]. Various methodologies have now been developed for characterising the tribo-corrosion interaction in erosion–corrosion environments, and in particular, in dry conditions [6]. There many “schools of thought” on the best method of describing such interactions and these have been the subject of a recent review [9]. However, such mechanistic descriptions are important because they enable the construction of a useful engineering tool for application to such environments, i.e. the engineering map.
∗ Corresponding author. Tel.: +44-141-5483754; fax: +44-141-5525105. E-mail address: [email protected] (M.M. Stack).
Erosion [10] and wear maps [11–14] are now part of mainstream tribology but it is worth noting that there are still many “territories” in this area that remain uncharted by such mapping approaches. In particular, although maps have been constructed for aqueous-corrosion (e.g. Pourbaix diagrams [15]) and erosion in such conditions [7,8], there has been no similar approach to characterising sliding wear in aqueous conditions. This is despite the fact that a map for dry sliding wear of steels (based on experimental results and constructed using analytical tools) [11] has been developed for over a decade. Some recent work by Jiang et al. [16] has provided a basis for construction of a simple map, based on mathematical models of wear–aqueous–corrosion in sliding contacts. However, there has been no laboratory work carried out to develop the map from raw data. In this work, the sliding wear behaviour in steels in an aqueous solution of buffered pH (carbonate/bicarbonate) was investigated at a range of velocities and applied loads. The results were used to establish regimes of wear–corrosion. Wear–corrosion maps were constructed as a function of sliding velocity and applied potential and the effects of increasing applied load on the wear–corrosion boundaries are discussed in terms of the competition between the frictional heating effects in the aqueous solutions versus the enhancement of mechanical wear.
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00203-5
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
2. Experimental description A pin-on-disc apparatus was used in this study (Fig. 1). The asymmetric test rig employed a pin pressed against a rotating ring both making conformal contact when the track rotates. Both the pin and ring were made of carbon steel of composition: 0.21% C, 0.38% Mn, 0.46% Si, 0.005% P and <0.004% S. The pre-weighed ring was fixed into a perspex trough acting as the container for the electrolyte (0.5 M NaHCO3 + 0.5 M Na2 CO3 solution prepared using de-ionised water). Importantly, the rear face and edges of the ring, and the cylindrical circumference of the pin, were coated with a lacquer (Lacomit) so as to limit the electrochemically exposed area solely to the abraded track (and also, necessarily, the pin face). The pin holder made it possible to apply variable force using a total weight of either 500 or 1000 g and also to connect the pin as the working electrode. A graphite auxiliary electrode ring was concentric to, but insulated from the pin. The rig was connected to a reference electrode (standard calomel electrode, SCE) by a salt bridge with a capillary placed close to, but not in contact with, the ring. The whole system was software controlled using a GillAC electrochemical interface from ACM Instruments (UK). Following an initial stabilisation period at −1000 mV (SCE) for 120 s, polarisation curves during simultaneous sliding wear were measured from −1000 to +500 mV (SCE) at a sweep rate of 1 mV s−1 . Necessarily, in this apparatus, the solution was exposed to air. Thus, the measured
457
current densities were corrected, at all potentials, for the presence of residual cathodic current by subtraction of the oxygen reduction current at −920 mV (SCE). The polarisation curves shown in this paper are presented after this correction. Prior to this process, the currents at −920 mV were plotted as a function of the square root of the rotation speed; the expected linear relationship confirmed that diffusion controlled kinetics dominated at this potential and that the correction was justified. Although a resistive (iR) drop may also have been present, it was assumed, due to the relatively high conductivity of the 0.5 M electrolyte, that this was minimal and no further corrections to the polarisation data were carried out. Wear–corrosion tests were carried out rotation speeds equivalent to linear velocities at the pin of 0.157, 0.314, 0.471, 0.628 m s−1 . The corrosion rates (Kc ) during sliding wear–corrosion were obtained from the current densities at potentials of −600, −400, −200, 0, 200 mV (SCE). Thus, the potential range covered active corrosion (−600 and −400 mV), pre-passivation (−400 and −200 mV) and passivation (−200 to 200 mV), depending on the sliding wear conditions. The total mass losses due to sliding wear and corrosion (Kwc ) at the above potentials after 1 h were obtained from the weight change of the ring; this was measured using a Mettler College 150 electronic balance, with an error of ±0.1 mg (it should be noted that for these experiments, the corrosion and wear rates were not separated; the corrosion rate refers to the corrosion of the pin and ring. The total area of the wear track on the ring was 31.42 cm2 ; the area of the pin was 0.5 cm2 . The wear and wear–corrosion
Fig. 1. Schematic diagram of experimental apparatus used for the wear–corrosion tests.
458
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
rates refer to the ring only; this was because tests on the pin showed that the wear and wear–corrosion losses were very low). Thus, the mass loss due to sliding wear during wear–corrosion (Kw ) can be obtained from subtraction of Kwc and Kc . The background corrosion rate in the absence of wear (Kco ) was obtained under static conditions. Finally, the background mass losses due to sliding wear in the absence of corrosion (Kwo ) were obtained during cathodic protection at −800 mV (SCE).
3. Results 3.1. Polarisation curves The polarisation curves for wear at 500 and 1000 g loads demonstrate the change in behaviour at various velocities (Fig. 2(a) and (b)). For loads of 500 g (Fig. 2(a)), the steel shows the expected active-to-passive transition between −700 and −500 mV
Fig. 2. Polarisation curves for steel at various sliding speeds and applied loads: (a) 500 g; (b) 1000 g.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
SCE. As the sliding wear rate increases, the maximum current density in the active loop tends to increase while the passive current density also increases. Although the corrosion (zero current) potential values are obscured by the high noise in the data, there is evidence that it tends to become more negative with increasing sliding wear. For 1000 g loads (Fig. 2(b)), the trends are similar to those at 500 g. Thus, the passive current densities increase, and the zero current density potentials become more negative, with increasing rotational velocity. Significantly, the currents were generally higher at all potentials with the higher load, this being particularly evident in the passive regime. Also, the potential at which the active-to-passive transition occurred was significantly delayed to more positive values. These effects are presumably due to the considerably more effective removal of the passive film under the higher load conditions. 3.2. Weight change data If
Kwc = Kw + Kc
459
Table 2 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.314 m s−1 and at 500 g load Potential Kwc
Kco
Kc
Kc
Kwo
Kw
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.963 0.203 0.244 0.243 0.246
0.830 0.151 0.200 0.201 0.203
1.100 1.100 1.100 1.100 1.100
0.737 −0.363 1.305 1.297 0.197 0.156 0.856 −0.244 0.286 0.657 −0.443 0.370 1.354 0.254 0.181
1.7 1.5 1.1 0.9 1.6
Kw
Kc /Kw
Table 3 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.471 m s−1 and at 500 g load Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
1.3 3 1.1 0.9 1.1
0.133 0.052 0.045 0.042 0.042
0.620 0.224 0.270 0.263 0.263
0.487 0.172 0.226 0.221 0.220
0.400 0.400 0.400 0.400 0.400
0.680 2.776 0.830 0.637 0.837
0.280 2.376 0.430 0.237 0.437
0.912 0.081 0.326 0.413 0.313
(i)
where, Kwc is the total wear–corrosion rate (measured as described above); Kw the total wear rate; Kc is the total corrosion rate. Kw can be written as Kwo + Kw
(ii)
where Kwo is the wear rate in the absence of corrosion; Kw is the “synergistic” effect of corrosion on the wear rate. Kc can be written as Kco + Kc where Kco is the corrosion rate in the absence of wear; Kc is the enhancement of corrosion due to wear, the so-called “additive” effect. The above analysis is based on an earlier approach used to characterise the various components of the wear–corrosion process. The results for the various contributions to the weight change are given in Tables 1–8 and Figs. 3–4. The corrosion rate data, Kc and Kco , were derived using Faraday’s law, e.g. Q Kc = (iii) ZF Mw it Kc = (iv) ZF
Table 1 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.157 m s−1 and at 500 g load
Table 4 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.628 m s−1 and at 500 g load Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.9 1.7 0.9 2.2 2.1
0.137 0.052 0.045 0.043 0.042
0.588 0.259 0.323 0.314 0.312
0.455 0.207 0.279 0.272 0.269
2.3 2.3 2.3 2.3 2.3
0.312 1.44 0.577 1.886 1.789
−1.988 −0.858 −1.723 −0.414 −0.512
1.89 0.180 0.560 0.167 0.174
Table 5 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.157 m s−1 and at 1000 g load Potential Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.551 0.181 0.214 0.212 0.206
0.419 0.129 0.169 0.170 0.164
1.200 1.200 1.200 1.200 1.200
0.449 0.919 0.486 0.588 1.094
−0.751 −0.281 −0.714 −0.612 −0.106
1.229 0.197 0.439 0.361 0.189
1.0 1.1 0.7 0.8 1.3
Table 6 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.314 m s−1 and at 1000 g load
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
1.1 0.7 0.9 1.2 2.0
0.133 0.052 0.045 0.043 0.042
0.587 0.159 0.201 0.198 0.199
0.454 0.108 0.156 0.156 0.156
1 1 1 1 1
0.514 0.541 0.699 1.001 1.801
−0.487 −0.469 −0.301 0.002 0.801
1.142 0.299 0.288 0.198 0.110
−0.6 −0.4 −0.2 0 0.2
1.7 1.5 1.2 2.0 1.5
0.133 0.052 0.045 0.042 0.042
0.506 0.197 0.275 0.256 0.258
0.374 0.146 0.230 0.214 0.216
0.700 0.700 0.700 0.700 0.700
1.194 1.303 0.925 1.744 1.242
0.494 0.603 0.225 1.044 0.542
0.424 0.152 0.297 0.147 0.208
460
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Table 7 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.471 m s−1 and at 1000 g load
Table 8 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.628 m s−1 and at 1000 g load
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
Potential Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.9 2.2 1.7 0.8 0.9
0.900 2.200 1.700 0.800 0.900
0.133 0.052 0.045 0.042 0.042
0.738 0.266 0.283 0.277 0.280
0.606 0.215 0.238 0.235 0.238
0.500 0.500 0.500 0.500 0.500
0.162 1.934 1.417 0.523 0.620
−0.338 1.434 0.917 0.023 0.120
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.738 0.266 0.283 0.277 0.280
0.606 0.215 0.238 0.235 0.238
2.500 2.500 2.500 2.500 2.500
0.962 0.734 1.617 0.423 0.920
−1.538 −1.766 −0.883 −2.077 −1.580
0.768 0.363 0.175 0.655 0.305
1.7 1.0 1.9 0.7 1.2
Fig. 3. Variation of weight change loss due to corrosion, Kc , due to wear, Kw , and due to wear–corrosion, Kwc , with velocity at applied loads of 500 g: (a) −0.6 V; (b) −0.4 V; (c) −0.2 V; (d) 0 V; (e) 0.2 V.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
461
Fig. 3. (Continued ).
where Q is the charge passed; F the Faraday’s constant (96494); n the number of electrons involved in corrosion process; i the current density; t the exposure time; Mw is the molecular weight of material The weight loss due to wear in the absence of corrosion, Kwo , was estimated by measuring the weight change in cathodic conditions. The results at the lower load of 500 g (Fig. 3) shows that at −0.6 V (Fig. 3(a)), the corrosion rate, Kc , exhibited a peak at intermediate velocities. There also appeared to be a peak in the wear rate, Kw and Kwc , at intermediate velocities. At higher potentials (Fig. 3(b)) −0.4 V, the corrosion rate was significantly lower with very little difference between the values of Kw and Kwc , with the peak in these values occurring at intermediate speeds. At higher potentials Fig. 3(c) −0.2 V, the values of Kc were higher than at −0.4 V (Fig. 3(b)). At potentials of −0.2 V (Fig. 3(d)), there was also evidence of a peak in the weight loss as a function of increasing velocity. At higher potentials of 0 and 0.2 V (Fig. 3(e) and (f)), there was an initial reduction in the Kw and Kwc values with increasing velocity, with these values increasing again at the highest rotational speeds, where the corrosion rates were at a maximum. The results at the higher load 1000 g (Fig. 4(a)), show a largely similar pattern to that observed at lower loads, at
potentials of −0.6 V, with a peak in the wear rate recorded as a function of increasing velocity; however in this case the values of Kw and Kwc attained minimum values at intermediate speeds and subsequently increased at higher values. The Kc value was high; it approached the value of Kw at the maximum rotational velocity. At higher potential values, −0.4 V, the characteristic peaks in the values of Kw and Kwc were recorded. At higher values, however (Fig. 4(c)), the peak disappeared and there was a continuous increase in these values with increasing rotational speed. The results at potentials of 0 and 0.2 V (Fig. 4(d) and (e)) showed that peaks in the values of Kw and Kwc as a function of velocity were also recorded; however, at the highest potential the wear rate appeared to attain a minimum value at intermediate velocities; further increases in rotational velocity caused an increase in the recorded weight changes. 3.3. Interactions of corrosion and wear on the steel surface The results show that there was a positive enhancement of corrosion rate due to wear, Kc , at the exposure potentials (Tables 1–8), for each of the applied loads. This was consistent with the polarisation data (Fig. 2), which showed that increasing the sliding speed shifted the passive current
462
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Fig. 4. Variation of weight change due to corrosion, Kc , due to wear, Kw , and due to wear–corrosion, Kwc with velocity at applied loads of 1000 g: (a) −0.6 V; (b) −0.4 V; (c) −0.2 V; (d) 0 V; (e) 0.2 V.
densities to higher values. The corrosion enhancement was highest at the active-to-passive transition, −0.6 V. Above this value, where the passive film formed, the values of Kc were lower. The Kc is the so-called “additive” effect of wear on corrosion and can be measured in situ from the Faradaic conversion of current density to weight change as described above. The values of Kw , the so-called “true” “synergistic” effect, varied as a function of potential, for the conditions tested (Tables 1–8). This indicated that for tests carried out
at 500 g loads, enhanced corrosion during the wear process, impeded wear at intermediate potentials (Tables 1–3), and at low applied loads and velocities, i.e. less than 0.471 m s−1 . At this velocity, corrosion appeared to enhance wear at the higher potentials, thus indicating that, at least for these conditions, corrosion was not effective in reducing wear. At the higher applied load of 1000 g, the values of Kw were negative at high velocities and at high potentials, i.e. 0.628 m s−1 . This was attributed to the formation of a protective tribo-chemical film in these conditions.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
463
Fig. 4. (Continued ).
3.4. Scanning electron microscopy of worn surfaces The micrographs (Fig. 5), show typical examples of the worn morphologies at various applied potentials. The results (at loads of 500 g) (Fig. 5(a)) showed that at low potentials of −0.8 V, at 0.314 m s−1 , there was extensive plastic deformation on the surface. As the potential was increased to −0.4 V, there was some evidence of oxide flakes on the surface consistent with the increase in corrosion at this potentials (Fig. 5(b)). At higher potentials of 0.2 V, the surface was heavily deformed (Fig. 5(c)). The corrosion on the surface appeared to be confined to the deformed surface proud of the wear track; the fact that there appeared to be no increase in corrosion rate compared to that at lower potentials (Fig. 5(b)) was consistent with the polarisation data (Fig. 2), and mass loss results (Table 3). 4. Discussion 4.1. Trends in wear–corrosion data as a function of increasing velocity and applied potential The results showed that the trends in wear rate data as a function of increasing velocity were similar to that in dry conditions, in which the classical peak in the wear rate has
been observed for steels [17]. This peak may be due to the effect of frictional heating, leading to formation of an oxide which provides some protection against wear (Figs. 3 and 4). That this peak changes at higher potentials at the lower load (Fig. 3(c)-(e)) may be indicative of the deleterious effects of aqueous-corrosion on the surface of this normally “protective” tribo-chemical film. The results at the higher load of 1000 g (Fig. 4(a) and (b)) at the lower potential values are largely similar to those recorded at 500 g (Fig. 3(a) and (b)) with the classical peak in the wear rate recorded with increasing rotational speed of the pin. The patterns at the higher loads (Fig. 4(c)-(e)) differ to those observed at the lower loads (Fig. 3(c)-(e)). The continuous increase in wear rate with increasing velocity, at −0.2 V (Fig. 4(c)) is not observed at high potentials (Fig. 4(d) and (e)) where the peak in the wear–corrosion rates is again observed. In these conditions (Fig. 4(d)), when the corrosion rates at high velocities, Kc , approach the wear rates, Kw , a reduction in the wear–corrosion rate, Kwc , is recorded. At 0.2 V (Fig. 4(e)), there appears to be range of rotational velocities in which the wear–corrosion rate is at a minimum indicating the ability of the corrosion film to provide protection against wear; at higher velocities, this film is removed again due mechanical wear. This indicates that the overall wear–corrosion rate Kwc can be reduced at higher velocities due to the protective nature of the tribo-chemical
464
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Fig. 5. Scanning electron micrographs of worn surfaces at 500 g loads and at 0.314 m s−1 : (a) −0.8 V; (b) −0.4 V; (c) 0.2 V.
film formed; wear occurring at higher loads tends to promote the development of this layer. Hence, the results indicate that a significant difference between the wear process in these passivating carbonate/bicarbonate conditions and wear in dry conditions, is the competition between frictional heating and corrosion of the iron oxide film which provides protection at high velocities. In these conditions, the iron oxide film appears to loose protection in aqueous conditions at low loads and high velocities and in the highest potential range studied. At high loads and velocities, the film which may be considerably thicker as indicated by the higher corrosion rates, tends to provide better protection during sliding, albeit in a narrow range of conditions. 4.2. Wear–corrosion regimes and wear maps in aqueous conditions The concept of wear–corrosion regimes can be applied to these conditions to distinguish between wear- and corrosion-dominated conditions.
The regimes can be defined as follows: Kc ≤ 0.1 Kw 0.1 < 1<
Kc ≤1 Kw
(wear) (wear − corrosion)
(v) (vi)
Kc ≤ 10 Kw
(corrosion − wear)
(vii)
Kc > 10 Kw
(corrosion)
(viii)
The corrosion process is further sub-divided to indicate “dissolution” (active corrosion) or “passivation” (film formation). The active-to-passive transition at various velocities is estimated from polarisation data (Fig. 2) (however, the corrosion process on the surface may involve processes of dissolution and film formation, even at constant potential due to the competition between the local oxidation rate due to frictional heating at the sliding contact, and the steady state electrochemical reaction defined by potential and pH). Clearly, the results at 500 g loads (Fig. 6(a)) shows that the
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
465
5. Conclusions (i) A study has been carried out of the sliding wear of steels in a carbonate/bicarbonate solution. (ii) While no direct interfacial temperature measurements were made in these experiments, the results indicate that frictional heating may have played a part in the wear–corrosion mechanism. (iii) Wear–corrosion maps have been constructed based on the results and show the differences in tribo-corrosion regimes as a function of electrochemical potential, applied load and sliding speed. References
Fig. 6. Velocity-potential wear–corrosion maps for steel at various applied loads: (a) 500 g; (b) 1000 g.
dissolution-wear-dominated regime occurs at low potentials. As the potential is increased, “wear-passivation” commences to dominate, due to the transition from active-to-passive corrosion behaviour (Fig. 2). The fact that there is a transition between “wear-dissolution” to “dissolution-wear”, and subsequently to “wear-dissolution” behaviour as a function of increasing velocity, at low potentials, is due to the distinctive effect of velocity in sliding wear. The predominance of the “wear-dissolution” regime at intermediate velocities is probably due to enhancement of mechanical damage on the surface as a function of increasing velocity. That the “dissolution-wear” regime extends to higher velocities is due to tribo-chemical effects as a result of increased frictional heating at such speeds leading to an enhanced role of the corrosion process. At 1000 g (Fig. 6(b)), the shift of the “dissolution-wear” regime on the map to higher velocities compared with that observed at 500 g (Fig. 6(a)), at low potentials, is probably due to the enhancement of mechanical wear. It is interesting that the “dissolution-wear” regime dominates the map to higher potential values, compared with the results at 500 g loads (Fig. 5(a)). This is probably due to the higher rates of frictional heating at the higher loads (such flash temperatures may increase significantly as a result of increasing load and this is well established in the literature [18]). This and the above observations indicate that in aqueous environments, the extent of tribo-corrosion on the surface is significant. Further work will be to investigate the wear–corrosion behaviour at other loads and pH values in order to understand more fully the characteristics of such sliding wear–corrosion maps in aqueous environments.
[1] Z. Yue, P. Zhou, J. Shi, Some factors influencing erosion–corrosion performance of materials, in: K.C. Ludema (Ed.), Proceedings of the Conference on Wear of Materials, Pub. ASME, 1987, p. 763. [2] M. Matsumura, Y. Oka, Hiura, M. Yano, The role of the passivating film in preventing slurry erosion–corrosion of austenitic stainless steel, ISIJ Int. 31(2) (1991) 168. [3] M. Bjordal, E. Bardal, T. Rogne, Erosion and corrosion properties of WC coatings and duplex stainless steel in sand containing synthetic sea water, Wear 186 (2) (1995) 508. [4] Y. Li, G.T. Burstein, I.M. Hutchings, Influence of environmental composition and electrochemical potential on the slurry erosion of aluminium, Wear 186–187 (1995) 515. [5] I. Garcia, D. Drees, J.P. Celis, Corrosion–wear of passivating materials in sliding contacts based on a concept of active wear track area, Wear 249 (2001) 452. [6] M.M. Stack, S. Zhou, R.C. Newman, Effects of particle velocity and applied potential on erosion of mild steel in carbonate/bicarbonate slurry, Mater. Sci. Tech. 12 (1996) 26. [7] M.M. Stack, N. Corlett, S. Zhou, A methodology for the construction of the erosion–corrosion map in aqueous environments, Wear 203–204 (1997) 474. [8] M.M. Stack, S. Zhou, R.C. Newman, Identification of transitions in erosion–corrosion regimes in aqueous environments, Wear 186–187 (1995) 523. [9] M.M. Stack, S. Lekatos, F.H. Stott, Erosion–corrosion regimes: number, nomenclature and justification, Tribol. Int. 28 (7) (1995) 445. [10] I.M. Hutchings, Ductile-brittle transitions and wear maps for the erosion and abrasion of brittle materials, J. Phys. D: App. Phys. 25 (1992) A121–A221. [11] S.C. Lim, M.F. Ashby, Wear mechanism maps, Acta. Metall. 35 (1) (1987) 1. [12] S.M. Hsu, Y.S. Yang, R.G. Munro, Quantitative wear maps as a visualization of wear mechanism transitions in ceramic materials, Wear 134 (1989) 1. [13] T.H.C. Childs, The sliding wear mechanisms of metals, mainly steels, Tribol. Int. 13 (1980) 285. [14] K. Adachi, K. Kato, N. Chen, Wear map of ceramics, Wear 203 (1987) 291. [15] M. Pourbaix, Atlas of Electrochemical Equilibria in Aqueous Solutions, Section 2, NACE, Houston, 1974. [16] J. Jiang, M.M. Stack, A. Neville, Modelling the tribo-corrosion interaction in aqueous sliding conditions, Tribol. Int. (2002). [17] W. Hirst, J.K. Lancaster, The influence of speed on metallic wear, Proc. R. Soc. London A259 (1960) 228. [18] J.F. Archard, Contact and rubbing of flat surfaces, J. Appl. Phys. 24 (1953) 98.
Mapping sliding wear of steels in aqueous conditions M.M. Stack∗ , K. Chi Department of Mechanical Engineering, University of Strathclyde, James Weir Building, 75 Montrose St., Glasgow G1 1XJ, UK
Abstract In studies of sliding wear in dry environments, there have been many attempts at mapping the processes. For steels, the classical “wear map” has been developed and various other studies have extended the approach to composites, ceramics and coated materials. Despite this work, there have been few attempts to extend the methodology to wet conditions, where the wear process interacts with solutions that are defined by pH and electrochemical potential. The object of this work was to study the sliding wear–corrosion behaviour of steels in a pin-on-disc apparatus in aqueous conditions. The effects of applied load and velocity were evaluated at various electrochemical potentials in carbonate/bicarbonate solution pH 9.8). The results were analysed using weight loss and scanning electron microscopy techniques. Wear mechanisms were identified in the various environments and a method of identifying the wear–corrosion transitions, in aqueous conditions, was proposed in the work. These regimes were superimposed on wear–corrosion maps, where the change in wear–corrosion regime was identified as a function of velocity and electrochemical potential. Possible reasons for the differences in the boundaries of the map at various applied loads are discussed in this paper. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Sliding wear corrosion; Maps; Wear mechanisms; Aqueous environments
1. Introduction Significant advances in recent years have been made in the study of tribo-corrosion in aqueous conditions. In particular, there has been much work carried out in the field of erosion–corrosion by solid particles [1–5]. Regimes of interaction have been identified, ranging from “erosion-dominated” where the material loss is predominately from the substrate to “corrosion-dominated” where the losses are as a result of dissolution or formation of corrosion product [6–8]. Various methodologies have now been developed for characterising the tribo-corrosion interaction in erosion–corrosion environments, and in particular, in dry conditions [6]. There many “schools of thought” on the best method of describing such interactions and these have been the subject of a recent review [9]. However, such mechanistic descriptions are important because they enable the construction of a useful engineering tool for application to such environments, i.e. the engineering map.
∗ Corresponding author. Tel.: +44-141-5483754; fax: +44-141-5525105. E-mail address: [email protected] (M.M. Stack).
Erosion [10] and wear maps [11–14] are now part of mainstream tribology but it is worth noting that there are still many “territories” in this area that remain uncharted by such mapping approaches. In particular, although maps have been constructed for aqueous-corrosion (e.g. Pourbaix diagrams [15]) and erosion in such conditions [7,8], there has been no similar approach to characterising sliding wear in aqueous conditions. This is despite the fact that a map for dry sliding wear of steels (based on experimental results and constructed using analytical tools) [11] has been developed for over a decade. Some recent work by Jiang et al. [16] has provided a basis for construction of a simple map, based on mathematical models of wear–aqueous–corrosion in sliding contacts. However, there has been no laboratory work carried out to develop the map from raw data. In this work, the sliding wear behaviour in steels in an aqueous solution of buffered pH (carbonate/bicarbonate) was investigated at a range of velocities and applied loads. The results were used to establish regimes of wear–corrosion. Wear–corrosion maps were constructed as a function of sliding velocity and applied potential and the effects of increasing applied load on the wear–corrosion boundaries are discussed in terms of the competition between the frictional heating effects in the aqueous solutions versus the enhancement of mechanical wear.
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00203-5
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
2. Experimental description A pin-on-disc apparatus was used in this study (Fig. 1). The asymmetric test rig employed a pin pressed against a rotating ring both making conformal contact when the track rotates. Both the pin and ring were made of carbon steel of composition: 0.21% C, 0.38% Mn, 0.46% Si, 0.005% P and <0.004% S. The pre-weighed ring was fixed into a perspex trough acting as the container for the electrolyte (0.5 M NaHCO3 + 0.5 M Na2 CO3 solution prepared using de-ionised water). Importantly, the rear face and edges of the ring, and the cylindrical circumference of the pin, were coated with a lacquer (Lacomit) so as to limit the electrochemically exposed area solely to the abraded track (and also, necessarily, the pin face). The pin holder made it possible to apply variable force using a total weight of either 500 or 1000 g and also to connect the pin as the working electrode. A graphite auxiliary electrode ring was concentric to, but insulated from the pin. The rig was connected to a reference electrode (standard calomel electrode, SCE) by a salt bridge with a capillary placed close to, but not in contact with, the ring. The whole system was software controlled using a GillAC electrochemical interface from ACM Instruments (UK). Following an initial stabilisation period at −1000 mV (SCE) for 120 s, polarisation curves during simultaneous sliding wear were measured from −1000 to +500 mV (SCE) at a sweep rate of 1 mV s−1 . Necessarily, in this apparatus, the solution was exposed to air. Thus, the measured
457
current densities were corrected, at all potentials, for the presence of residual cathodic current by subtraction of the oxygen reduction current at −920 mV (SCE). The polarisation curves shown in this paper are presented after this correction. Prior to this process, the currents at −920 mV were plotted as a function of the square root of the rotation speed; the expected linear relationship confirmed that diffusion controlled kinetics dominated at this potential and that the correction was justified. Although a resistive (iR) drop may also have been present, it was assumed, due to the relatively high conductivity of the 0.5 M electrolyte, that this was minimal and no further corrections to the polarisation data were carried out. Wear–corrosion tests were carried out rotation speeds equivalent to linear velocities at the pin of 0.157, 0.314, 0.471, 0.628 m s−1 . The corrosion rates (Kc ) during sliding wear–corrosion were obtained from the current densities at potentials of −600, −400, −200, 0, 200 mV (SCE). Thus, the potential range covered active corrosion (−600 and −400 mV), pre-passivation (−400 and −200 mV) and passivation (−200 to 200 mV), depending on the sliding wear conditions. The total mass losses due to sliding wear and corrosion (Kwc ) at the above potentials after 1 h were obtained from the weight change of the ring; this was measured using a Mettler College 150 electronic balance, with an error of ±0.1 mg (it should be noted that for these experiments, the corrosion and wear rates were not separated; the corrosion rate refers to the corrosion of the pin and ring. The total area of the wear track on the ring was 31.42 cm2 ; the area of the pin was 0.5 cm2 . The wear and wear–corrosion
Fig. 1. Schematic diagram of experimental apparatus used for the wear–corrosion tests.
458
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
rates refer to the ring only; this was because tests on the pin showed that the wear and wear–corrosion losses were very low). Thus, the mass loss due to sliding wear during wear–corrosion (Kw ) can be obtained from subtraction of Kwc and Kc . The background corrosion rate in the absence of wear (Kco ) was obtained under static conditions. Finally, the background mass losses due to sliding wear in the absence of corrosion (Kwo ) were obtained during cathodic protection at −800 mV (SCE).
3. Results 3.1. Polarisation curves The polarisation curves for wear at 500 and 1000 g loads demonstrate the change in behaviour at various velocities (Fig. 2(a) and (b)). For loads of 500 g (Fig. 2(a)), the steel shows the expected active-to-passive transition between −700 and −500 mV
Fig. 2. Polarisation curves for steel at various sliding speeds and applied loads: (a) 500 g; (b) 1000 g.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
SCE. As the sliding wear rate increases, the maximum current density in the active loop tends to increase while the passive current density also increases. Although the corrosion (zero current) potential values are obscured by the high noise in the data, there is evidence that it tends to become more negative with increasing sliding wear. For 1000 g loads (Fig. 2(b)), the trends are similar to those at 500 g. Thus, the passive current densities increase, and the zero current density potentials become more negative, with increasing rotational velocity. Significantly, the currents were generally higher at all potentials with the higher load, this being particularly evident in the passive regime. Also, the potential at which the active-to-passive transition occurred was significantly delayed to more positive values. These effects are presumably due to the considerably more effective removal of the passive film under the higher load conditions. 3.2. Weight change data If
Kwc = Kw + Kc
459
Table 2 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.314 m s−1 and at 500 g load Potential Kwc
Kco
Kc
Kc
Kwo
Kw
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.963 0.203 0.244 0.243 0.246
0.830 0.151 0.200 0.201 0.203
1.100 1.100 1.100 1.100 1.100
0.737 −0.363 1.305 1.297 0.197 0.156 0.856 −0.244 0.286 0.657 −0.443 0.370 1.354 0.254 0.181
1.7 1.5 1.1 0.9 1.6
Kw
Kc /Kw
Table 3 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.471 m s−1 and at 500 g load Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
1.3 3 1.1 0.9 1.1
0.133 0.052 0.045 0.042 0.042
0.620 0.224 0.270 0.263 0.263
0.487 0.172 0.226 0.221 0.220
0.400 0.400 0.400 0.400 0.400
0.680 2.776 0.830 0.637 0.837
0.280 2.376 0.430 0.237 0.437
0.912 0.081 0.326 0.413 0.313
(i)
where, Kwc is the total wear–corrosion rate (measured as described above); Kw the total wear rate; Kc is the total corrosion rate. Kw can be written as Kwo + Kw
(ii)
where Kwo is the wear rate in the absence of corrosion; Kw is the “synergistic” effect of corrosion on the wear rate. Kc can be written as Kco + Kc where Kco is the corrosion rate in the absence of wear; Kc is the enhancement of corrosion due to wear, the so-called “additive” effect. The above analysis is based on an earlier approach used to characterise the various components of the wear–corrosion process. The results for the various contributions to the weight change are given in Tables 1–8 and Figs. 3–4. The corrosion rate data, Kc and Kco , were derived using Faraday’s law, e.g. Q Kc = (iii) ZF Mw it Kc = (iv) ZF
Table 1 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.157 m s−1 and at 500 g load
Table 4 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.628 m s−1 and at 500 g load Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.9 1.7 0.9 2.2 2.1
0.137 0.052 0.045 0.043 0.042
0.588 0.259 0.323 0.314 0.312
0.455 0.207 0.279 0.272 0.269
2.3 2.3 2.3 2.3 2.3
0.312 1.44 0.577 1.886 1.789
−1.988 −0.858 −1.723 −0.414 −0.512
1.89 0.180 0.560 0.167 0.174
Table 5 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.157 m s−1 and at 1000 g load Potential Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.551 0.181 0.214 0.212 0.206
0.419 0.129 0.169 0.170 0.164
1.200 1.200 1.200 1.200 1.200
0.449 0.919 0.486 0.588 1.094
−0.751 −0.281 −0.714 −0.612 −0.106
1.229 0.197 0.439 0.361 0.189
1.0 1.1 0.7 0.8 1.3
Table 6 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.314 m s−1 and at 1000 g load
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
1.1 0.7 0.9 1.2 2.0
0.133 0.052 0.045 0.043 0.042
0.587 0.159 0.201 0.198 0.199
0.454 0.108 0.156 0.156 0.156
1 1 1 1 1
0.514 0.541 0.699 1.001 1.801
−0.487 −0.469 −0.301 0.002 0.801
1.142 0.299 0.288 0.198 0.110
−0.6 −0.4 −0.2 0 0.2
1.7 1.5 1.2 2.0 1.5
0.133 0.052 0.045 0.042 0.042
0.506 0.197 0.275 0.256 0.258
0.374 0.146 0.230 0.214 0.216
0.700 0.700 0.700 0.700 0.700
1.194 1.303 0.925 1.744 1.242
0.494 0.603 0.225 1.044 0.542
0.424 0.152 0.297 0.147 0.208
460
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Table 7 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.471 m s−1 and at 1000 g load
Table 8 Mass loss data for corrosion, wear and wear–corrosion contributions at 0.628 m s−1 and at 1000 g load
Potential
Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
Potential Kwc
Kco
Kc
Kc
Kwo
Kw
Kw
Kc /Kw
−0.6 −0.4 −0.2 0 0.2
0.9 2.2 1.7 0.8 0.9
0.900 2.200 1.700 0.800 0.900
0.133 0.052 0.045 0.042 0.042
0.738 0.266 0.283 0.277 0.280
0.606 0.215 0.238 0.235 0.238
0.500 0.500 0.500 0.500 0.500
0.162 1.934 1.417 0.523 0.620
−0.338 1.434 0.917 0.023 0.120
−0.6 −0.4 −0.2 0 0.2
0.133 0.052 0.045 0.042 0.042
0.738 0.266 0.283 0.277 0.280
0.606 0.215 0.238 0.235 0.238
2.500 2.500 2.500 2.500 2.500
0.962 0.734 1.617 0.423 0.920
−1.538 −1.766 −0.883 −2.077 −1.580
0.768 0.363 0.175 0.655 0.305
1.7 1.0 1.9 0.7 1.2
Fig. 3. Variation of weight change loss due to corrosion, Kc , due to wear, Kw , and due to wear–corrosion, Kwc , with velocity at applied loads of 500 g: (a) −0.6 V; (b) −0.4 V; (c) −0.2 V; (d) 0 V; (e) 0.2 V.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
461
Fig. 3. (Continued ).
where Q is the charge passed; F the Faraday’s constant (96494); n the number of electrons involved in corrosion process; i the current density; t the exposure time; Mw is the molecular weight of material The weight loss due to wear in the absence of corrosion, Kwo , was estimated by measuring the weight change in cathodic conditions. The results at the lower load of 500 g (Fig. 3) shows that at −0.6 V (Fig. 3(a)), the corrosion rate, Kc , exhibited a peak at intermediate velocities. There also appeared to be a peak in the wear rate, Kw and Kwc , at intermediate velocities. At higher potentials (Fig. 3(b)) −0.4 V, the corrosion rate was significantly lower with very little difference between the values of Kw and Kwc , with the peak in these values occurring at intermediate speeds. At higher potentials Fig. 3(c) −0.2 V, the values of Kc were higher than at −0.4 V (Fig. 3(b)). At potentials of −0.2 V (Fig. 3(d)), there was also evidence of a peak in the weight loss as a function of increasing velocity. At higher potentials of 0 and 0.2 V (Fig. 3(e) and (f)), there was an initial reduction in the Kw and Kwc values with increasing velocity, with these values increasing again at the highest rotational speeds, where the corrosion rates were at a maximum. The results at the higher load 1000 g (Fig. 4(a)), show a largely similar pattern to that observed at lower loads, at
potentials of −0.6 V, with a peak in the wear rate recorded as a function of increasing velocity; however in this case the values of Kw and Kwc attained minimum values at intermediate speeds and subsequently increased at higher values. The Kc value was high; it approached the value of Kw at the maximum rotational velocity. At higher potential values, −0.4 V, the characteristic peaks in the values of Kw and Kwc were recorded. At higher values, however (Fig. 4(c)), the peak disappeared and there was a continuous increase in these values with increasing rotational speed. The results at potentials of 0 and 0.2 V (Fig. 4(d) and (e)) showed that peaks in the values of Kw and Kwc as a function of velocity were also recorded; however, at the highest potential the wear rate appeared to attain a minimum value at intermediate velocities; further increases in rotational velocity caused an increase in the recorded weight changes. 3.3. Interactions of corrosion and wear on the steel surface The results show that there was a positive enhancement of corrosion rate due to wear, Kc , at the exposure potentials (Tables 1–8), for each of the applied loads. This was consistent with the polarisation data (Fig. 2), which showed that increasing the sliding speed shifted the passive current
462
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Fig. 4. Variation of weight change due to corrosion, Kc , due to wear, Kw , and due to wear–corrosion, Kwc with velocity at applied loads of 1000 g: (a) −0.6 V; (b) −0.4 V; (c) −0.2 V; (d) 0 V; (e) 0.2 V.
densities to higher values. The corrosion enhancement was highest at the active-to-passive transition, −0.6 V. Above this value, where the passive film formed, the values of Kc were lower. The Kc is the so-called “additive” effect of wear on corrosion and can be measured in situ from the Faradaic conversion of current density to weight change as described above. The values of Kw , the so-called “true” “synergistic” effect, varied as a function of potential, for the conditions tested (Tables 1–8). This indicated that for tests carried out
at 500 g loads, enhanced corrosion during the wear process, impeded wear at intermediate potentials (Tables 1–3), and at low applied loads and velocities, i.e. less than 0.471 m s−1 . At this velocity, corrosion appeared to enhance wear at the higher potentials, thus indicating that, at least for these conditions, corrosion was not effective in reducing wear. At the higher applied load of 1000 g, the values of Kw were negative at high velocities and at high potentials, i.e. 0.628 m s−1 . This was attributed to the formation of a protective tribo-chemical film in these conditions.
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
463
Fig. 4. (Continued ).
3.4. Scanning electron microscopy of worn surfaces The micrographs (Fig. 5), show typical examples of the worn morphologies at various applied potentials. The results (at loads of 500 g) (Fig. 5(a)) showed that at low potentials of −0.8 V, at 0.314 m s−1 , there was extensive plastic deformation on the surface. As the potential was increased to −0.4 V, there was some evidence of oxide flakes on the surface consistent with the increase in corrosion at this potentials (Fig. 5(b)). At higher potentials of 0.2 V, the surface was heavily deformed (Fig. 5(c)). The corrosion on the surface appeared to be confined to the deformed surface proud of the wear track; the fact that there appeared to be no increase in corrosion rate compared to that at lower potentials (Fig. 5(b)) was consistent with the polarisation data (Fig. 2), and mass loss results (Table 3). 4. Discussion 4.1. Trends in wear–corrosion data as a function of increasing velocity and applied potential The results showed that the trends in wear rate data as a function of increasing velocity were similar to that in dry conditions, in which the classical peak in the wear rate has
been observed for steels [17]. This peak may be due to the effect of frictional heating, leading to formation of an oxide which provides some protection against wear (Figs. 3 and 4). That this peak changes at higher potentials at the lower load (Fig. 3(c)-(e)) may be indicative of the deleterious effects of aqueous-corrosion on the surface of this normally “protective” tribo-chemical film. The results at the higher load of 1000 g (Fig. 4(a) and (b)) at the lower potential values are largely similar to those recorded at 500 g (Fig. 3(a) and (b)) with the classical peak in the wear rate recorded with increasing rotational speed of the pin. The patterns at the higher loads (Fig. 4(c)-(e)) differ to those observed at the lower loads (Fig. 3(c)-(e)). The continuous increase in wear rate with increasing velocity, at −0.2 V (Fig. 4(c)) is not observed at high potentials (Fig. 4(d) and (e)) where the peak in the wear–corrosion rates is again observed. In these conditions (Fig. 4(d)), when the corrosion rates at high velocities, Kc , approach the wear rates, Kw , a reduction in the wear–corrosion rate, Kwc , is recorded. At 0.2 V (Fig. 4(e)), there appears to be range of rotational velocities in which the wear–corrosion rate is at a minimum indicating the ability of the corrosion film to provide protection against wear; at higher velocities, this film is removed again due mechanical wear. This indicates that the overall wear–corrosion rate Kwc can be reduced at higher velocities due to the protective nature of the tribo-chemical
464
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
Fig. 5. Scanning electron micrographs of worn surfaces at 500 g loads and at 0.314 m s−1 : (a) −0.8 V; (b) −0.4 V; (c) 0.2 V.
film formed; wear occurring at higher loads tends to promote the development of this layer. Hence, the results indicate that a significant difference between the wear process in these passivating carbonate/bicarbonate conditions and wear in dry conditions, is the competition between frictional heating and corrosion of the iron oxide film which provides protection at high velocities. In these conditions, the iron oxide film appears to loose protection in aqueous conditions at low loads and high velocities and in the highest potential range studied. At high loads and velocities, the film which may be considerably thicker as indicated by the higher corrosion rates, tends to provide better protection during sliding, albeit in a narrow range of conditions. 4.2. Wear–corrosion regimes and wear maps in aqueous conditions The concept of wear–corrosion regimes can be applied to these conditions to distinguish between wear- and corrosion-dominated conditions.
The regimes can be defined as follows: Kc ≤ 0.1 Kw 0.1 < 1<
Kc ≤1 Kw
(wear) (wear − corrosion)
(v) (vi)
Kc ≤ 10 Kw
(corrosion − wear)
(vii)
Kc > 10 Kw
(corrosion)
(viii)
The corrosion process is further sub-divided to indicate “dissolution” (active corrosion) or “passivation” (film formation). The active-to-passive transition at various velocities is estimated from polarisation data (Fig. 2) (however, the corrosion process on the surface may involve processes of dissolution and film formation, even at constant potential due to the competition between the local oxidation rate due to frictional heating at the sliding contact, and the steady state electrochemical reaction defined by potential and pH). Clearly, the results at 500 g loads (Fig. 6(a)) shows that the
M.M. Stack, K. Chi / Wear 255 (2003) 456–465
465
5. Conclusions (i) A study has been carried out of the sliding wear of steels in a carbonate/bicarbonate solution. (ii) While no direct interfacial temperature measurements were made in these experiments, the results indicate that frictional heating may have played a part in the wear–corrosion mechanism. (iii) Wear–corrosion maps have been constructed based on the results and show the differences in tribo-corrosion regimes as a function of electrochemical potential, applied load and sliding speed. References
Fig. 6. Velocity-potential wear–corrosion maps for steel at various applied loads: (a) 500 g; (b) 1000 g.
dissolution-wear-dominated regime occurs at low potentials. As the potential is increased, “wear-passivation” commences to dominate, due to the transition from active-to-passive corrosion behaviour (Fig. 2). The fact that there is a transition between “wear-dissolution” to “dissolution-wear”, and subsequently to “wear-dissolution” behaviour as a function of increasing velocity, at low potentials, is due to the distinctive effect of velocity in sliding wear. The predominance of the “wear-dissolution” regime at intermediate velocities is probably due to enhancement of mechanical damage on the surface as a function of increasing velocity. That the “dissolution-wear” regime extends to higher velocities is due to tribo-chemical effects as a result of increased frictional heating at such speeds leading to an enhanced role of the corrosion process. At 1000 g (Fig. 6(b)), the shift of the “dissolution-wear” regime on the map to higher velocities compared with that observed at 500 g (Fig. 6(a)), at low potentials, is probably due to the enhancement of mechanical wear. It is interesting that the “dissolution-wear” regime dominates the map to higher potential values, compared with the results at 500 g loads (Fig. 5(a)). This is probably due to the higher rates of frictional heating at the higher loads (such flash temperatures may increase significantly as a result of increasing load and this is well established in the literature [18]). This and the above observations indicate that in aqueous environments, the extent of tribo-corrosion on the surface is significant. Further work will be to investigate the wear–corrosion behaviour at other loads and pH values in order to understand more fully the characteristics of such sliding wear–corrosion maps in aqueous environments.
[1] Z. Yue, P. Zhou, J. Shi, Some factors influencing erosion–corrosion performance of materials, in: K.C. Ludema (Ed.), Proceedings of the Conference on Wear of Materials, Pub. ASME, 1987, p. 763. [2] M. Matsumura, Y. Oka, Hiura, M. Yano, The role of the passivating film in preventing slurry erosion–corrosion of austenitic stainless steel, ISIJ Int. 31(2) (1991) 168. [3] M. Bjordal, E. Bardal, T. Rogne, Erosion and corrosion properties of WC coatings and duplex stainless steel in sand containing synthetic sea water, Wear 186 (2) (1995) 508. [4] Y. Li, G.T. Burstein, I.M. Hutchings, Influence of environmental composition and electrochemical potential on the slurry erosion of aluminium, Wear 186–187 (1995) 515. [5] I. Garcia, D. Drees, J.P. Celis, Corrosion–wear of passivating materials in sliding contacts based on a concept of active wear track area, Wear 249 (2001) 452. [6] M.M. Stack, S. Zhou, R.C. Newman, Effects of particle velocity and applied potential on erosion of mild steel in carbonate/bicarbonate slurry, Mater. Sci. Tech. 12 (1996) 26. [7] M.M. Stack, N. Corlett, S. Zhou, A methodology for the construction of the erosion–corrosion map in aqueous environments, Wear 203–204 (1997) 474. [8] M.M. Stack, S. Zhou, R.C. Newman, Identification of transitions in erosion–corrosion regimes in aqueous environments, Wear 186–187 (1995) 523. [9] M.M. Stack, S. Lekatos, F.H. Stott, Erosion–corrosion regimes: number, nomenclature and justification, Tribol. Int. 28 (7) (1995) 445. [10] I.M. Hutchings, Ductile-brittle transitions and wear maps for the erosion and abrasion of brittle materials, J. Phys. D: App. Phys. 25 (1992) A121–A221. [11] S.C. Lim, M.F. Ashby, Wear mechanism maps, Acta. Metall. 35 (1) (1987) 1. [12] S.M. Hsu, Y.S. Yang, R.G. Munro, Quantitative wear maps as a visualization of wear mechanism transitions in ceramic materials, Wear 134 (1989) 1. [13] T.H.C. Childs, The sliding wear mechanisms of metals, mainly steels, Tribol. Int. 13 (1980) 285. [14] K. Adachi, K. Kato, N. Chen, Wear map of ceramics, Wear 203 (1987) 291. [15] M. Pourbaix, Atlas of Electrochemical Equilibria in Aqueous Solutions, Section 2, NACE, Houston, 1974. [16] J. Jiang, M.M. Stack, A. Neville, Modelling the tribo-corrosion interaction in aqueous sliding conditions, Tribol. Int. (2002). [17] W. Hirst, J.K. Lancaster, The influence of speed on metallic wear, Proc. R. Soc. London A259 (1960) 228. [18] J.F. Archard, Contact and rubbing of flat surfaces, J. Appl. Phys. 24 (1953) 98.