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Micro-abrasion transitions of metallic materials
Wear 255 (2003) 14–22
Micro-abrasion transitions of metallic materials M.M. Stack∗ , M. Mathew Department of Mechanical Engineering, University of Strathclyde, 75 Montrose St., Glasgow, G1 1XJ, UK
Abstract Significant progress has been made in the understanding of micro-abrasion transitions of various materials in recent years, where abrasion is caused by particle sizes which are typically less than 10 m. Research has shown reasonable consistency on effects of applied load and sliding distance, for studies carried out in various laboratories. In addition, attempts have been made to construct abrasion “diagrams” showing the transitions between the various regimes as a function of the above parameters. A puzzling aspect of the results to date, however, is the effect of “ridge” development on the micro-abrasion wear pattern. This ridge formation leads to a reduction in wear rate because the size of abrading particles, which account for the three-body effect, is less than the size of the ridge developed as a result of the wear process. The particles thus become lost in the ridge and cause no further abrasion. In this paper, the transition to ridge formation, as a function of sliding distance and load, is described for a range of pure metals of different hardness, using a TE-66, Plint micro-abrasion test rig. Tally-surf and optical microscopy techniques were used to measure the wear rate. Micro-abrasion maps were constructed showing differences in the transition boundaries between the wear regimes, and the implications of such results, for predictive modelling of micro-abrasion, are addressed in this paper. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Micro-abrasion; Pure metals; Wear mechanism maps; Wear regimes
1. Introduction Significant developments have been made in the understanding of micro-abrasion mechanisms of materials in recent years. There is now a considerable body of literature on the performance of a range of materials in such conditions from PVD coatings to paints [1–6]. Mechanisms of micro-abrasion have been identified based on the observed deformation in the various conditions. There is, however, still some confusion on both the possible number of abrasion regimes and the nomenclature used to define such transitions. The three-body abrasion mechanism between particles entrained in the contact may revert to a two-body process if the particles adhere rigidly to one of the surfaces [4]. On the other hand, if the particles disappear between ridges which form during the process, the wear mechanism must involves only the surfaces in contact and this again is a two-body process, albeit entirely different in nature to the latter mechanism [5]. Wear mechanism maps have been produced for such processes [6,7] and again there are some differences in the terminology applied by investigators. Some mapping approaches ∗ Corresponding author. Tel.: +44-141-5483754; fax: +44-141-5525105. E-mail address: [email protected] (M.M. Stack).
have employed the established nomenclature, i.e. three-body, two-body abrasion to describe the process, extending it to distinguish between rolling and grooving wear [4]. Other investigators have differentiated between ridge formation and three-body abrasive wear [5]. It has been pointed out that, for such a complex process, the terminology used to describe the regimes should be carefully chosen in order to describe accurately the surface damage mechanisms [6]. In this paper, a series of tests were carried out for the effect of sliding distance and load on the micro-abrasion of three metallic materials, aluminium, copper and steel. The results are interpreted in terms of different regimes of behaviour and a mechanistic description of the various regimes is proposed. Erosion maps, as a function of sliding distance and load, are presented for each of the materials and the differences between the maps are explained in terms of the tribological properties of the various materials.
2. Experimental details Micro-abrasion tests were performed with a commercially available apparatus, the TE-66, micro-abrasion tester (Fig. 1) (Plint, UK), based on established design (3–4). The details of the experimental rig are as follows. A 25 mm ball is located between two-coaxial shafts each carried in support bearing.
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00204-7
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
Fig. 1. Schematic diagram of experimental apparatus.
One shaft is driven by a variable speed dc geared motor. A batch counter is provided to measure and control the number of shaft revolutions. A peristaltic pump head is connected to the other end of the shaft and this is used for providing slurry feed to the contact. The test sample is clamped onto a platform, which is fitted to the pivoted L-shaped arm. This arm is rotated around its pivot until the sample comes into contact with the ball. The beam is in balance when the samples are just in contact and the load is applied by adding dead weights to a cantilever arm. Each data point represents one experimental result; the error in the results was ±12% based on the average of results from three consecutive tests. There were no measurements taken for the steel balls (although it is acknowledged that wear of the steel balls will also be a major issue in such a test). On average, a new ball was used after eight tests. The slurry is stored in a container that can be agitated on a laboratory magnetic stirrer. The slurry is pumped by the integral peristaltic pump. It is fed to a position just above the contact point and collected in a waste tray underneath. The specification of the apparatus and experimental details are shown in Tables 1 and 2. The arm, which holds the sample, can be moved horizontally in order that several tests on a single sample specimen can be carried out. This configuration has the Table 1 Specification of the micro-abrasion apparatus Name and model Supplier Load range (N) Ball diameter (mm) Ball speed range (rpm) Pump feed rate (ml h−1 ) Software used
Micro-abrasion tester, TE-66 (Plint), UK 0.05–5 25 30–150 Up to 60 (based on 0.5 m bore) COMPEND 2000
Table 2 Experimental details Sample materials Ball materials Speed (rpm) Load (N) Sliding distance (rev) Slurry
Mild steel (2% carbon), aluminium, copper 304 stainless steel (supplied by SKF bearings) 100 1–5 150, 375, 750, 1500, 3000 or (11775, 29438, 58875, 117750, 235500 mm) Slurry composition: silicon carbide (4 m diameter) with distilled water (concentration of 0.025 g cm−1 )
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advantage of accurate control of both normal load (to an accuracy of ±0.01 N) and sliding speed. The sample is then removed from the apparatus and the diameter of the resulting abrasion scars are measured with profile projector and optical microscope. The material of the ball used was 304 stainless steel (of hardness 721 VHN at 5 kg load, 388 kg mm−2 ). Tests were conducted on mild steel (0.2% C), aluminium and copper. The hardness (VHN) of these materials were mild steel, 258; copper, 48 and aluminium, 38 at 3 kg load (138.69, 25.88 and 20.49 kg mm−2 , respectively). The sample surfaces were ground and polished by conventional metallographic methods before testing. Following the test, the worn samples were examined by optical and scanning electron microscopy and by a profilometer. The wear volume was calculated using the standard technique for measuring the wear scar of spherical geometry [3], i.e. the geometry of the wear scar is assumed to reproduce the spherical geometry of the ball, and the wear volume (V) may then be calculated by measurement of either the crater diameter (b) or its depth (h): V ≈
Πb4 , 64R
V ≈ Πh2 R,
for b R for h R
(1) (2)
3. Results 3.1. Effect of sliding distance on applied load on the micro-abrasion rates The results on the variation of wear rate with applied load (Fig. 2a) show that at a sliding distance of 11 775 mm, there was a peak in the wear rate for aluminium; the peak shifted to higher loads for copper (it should be noted that the groove diameter, on which the wear volume measurements were calculated, was measured using profilometry. Optical microscopy was used to confirm these measurements). The micro-abrasion rates of the steel were significantly less than those for the pure metals even at the highest loads. There was very little difference between the wear rates of all three materials at the highest loads of 5 N. As the sliding distance was increased to 29 438 mm (Fig. 2b), the characteristic peak was again observed for the aluminium; for the steel this tended to occur at higher loads. However, at sliding distances of 58 875 mm (Fig. 2c), a transition appeared to have occurred. There appeared to be a continuous increase in wear rate for aluminium; this behaviour now was similar to that of steel. However, for copper, the peak in the wear rate versus applied load occurred at lower loads than that of the other materials. At significantly higher sliding distances, 11 750 mm (Fig. 2d), there was no definite pattern to the results; however, the peak in the wear rate for aluminium appeared to occur at lower loads than at half the sliding distance
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(Fig. 2c). At the highest sliding distance, 235 500 mm, the wear rate of the aluminium was greatest at the lowest load of 1 N; at higher loads there was little differences between the wear rates of the materials. There was some evidence that the peak in the wear rate for the steel had shifted to lower loads compared to that observed at roughly half the sliding distance (Fig. 2c).
3.2. Scanning electron micrographs of the abraded surfaces The micrographs (Fig. 3) show various regimes of micro-abrasion for the various materials. At sliding distances of 58 875 mm (Fig. 3a) for steel, there was significant plastic deformation on the surfaces, some evidence of grooving was
Fig. 2. Variation of wear rate with applied load for the following sliding distances: (a) 11 775 mm; (b) 29 438 mm; (c) 58 875 mm; (d) 11 750 mm; (e) 235 500 mm.
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
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Fig. 2. (Continued ).
Fig. 3. Scanning electron micrographs of abraded surfaces: (a) mild steel, 1 N, 58 875 mm; (b) copper, 2 N, 235 500 mm; (c) copper as in (b) shows ridge formation at high magnification; (d) aluminium, 2 N, 235 500 mm; (e) aluminium as in (d) shows ridge formation at high magnification; (f) aluminium as in (d) shows oxide film at low magnification.
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M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
also observed. At higher sliding distances of 235 500 mm and at higher loads for copper (Fig. 3b) there was evidence of ridge formation. At higher magnifications (Fig. 3c) there was also evidence of ridge formation due to grooving wear. For aluminium, at the highest sliding distance of test regime (Figs. 3d and e) a similar mechanism was observed. However, there also appeared to be some evidence of oxide scale on the surface in these conditions (Fig. 3f). 4. Discussion 4.1. Mechanisms of micro-abrasion of the various materials as a function of applied load and sliding distance The volume loss due to wear, V, is based on Archard’s equation [8] and can be written as follows [9]: V = κSN
(3)
where κ is the constant, S the sliding distance and N the normal load. Clearly, the results (Fig. 2) show that this equation is only valid in a minority of cases for the materials tested because the wear rate is observed to reduce with increased load in many of the tests carried out in this apparatus design [4,5]. However, this pattern, in which there is a transition to ridge formation above a critical load has been seen in previous work in this area [4,5]. That such transitions occur at different loads for a range of pure metals had not been observed. It should be noted that the results are dependent on the approximation to semi-spherical wear scars of the raw data. The scars may on occasion appear more elliptical than semi-spherical (depending on the corresponding wear of the counter face) and this is a major issue for determination of micro-abrasion rates using such an apparatus. The three-body to two-body transition has been attributed to the embedment of particles in the surface which results in a grooving mechanism [4,5]. The resultant decrease in wear rate above this transition load may be due to entrainment of particles in the grooves leading to another kind of two-body abrasion mechanism where the abrading surfaces are in contact. Hence, there are a number of possible mechanisms of micro-abrasion and this is addressed further below. It is interesting that although the micro-abrasion rate largely increased with sliding distance (Fig. 2) such increases were not linear due to the transitions observed above, and again this points to deviation from the Archard’s equation, Eq. (3). It has been pointed out that as sliding distance increases, the area of the contact will change [4] (and generally increase) thus resulting in a possible reduction of the load per abrasive particle with increasing sliding distance. Frictional heating leading to oxide formation will also change the characteristics of the area under contact; thus, there are various reasons why deviation from linear behaviour would be expected.
It is interesting that at low sliding distances (Fig. 2a) the critical load at which the peak in the wear rate occurs is observed, shifts to higher values for copper compared to aluminium. This is likely to be due to the lower hardness of the aluminium compared to the copper, i.e. 21 kg mm−2 compared to 26 kg mm−2 , which would result in ridge formation occurring at lower loads for the softer material. The fact that the micro-abrasion rates approach each other at the high loads is possibly due to a range of factors; differential work hardening rates [9] and greater influence of tribo-chemical mechanisms in such conditions. In the latter case, there may be less difference in the wear rate of such oxides than for the substrate materials. For longer sliding distances (Fig. 2b), the peak in the wear rate for aluminium at low loads indicates that hardness is still playing a role in determining the transition to “ridge-dominated” micro-abrasion behaviour. There is some evidence that the peak for the steel occurs at a significantly higher load, 4 N, and this is attributed to its significantly higher hardness value (139 kg mm−2 ) and the higher work hardening rate [2]. At intermediate sliding distances (Fig. 2c), the peak in the wear rate for the aluminium and steel shifts to significantly higher loads; however, the peak for the copper remains at intermediate loads. For steel, this may be consistent with the fact that the work hardening rate of this material is higher than for copper. However, the reason why the peak should shift to higher values for aluminium is unclear (there is evidence for higher work hardening rates of steel compared to copper in the literature; the picture for aluminium versus copper is less clear [9]). However, for longer time sliding distances, there appears to be very little difference between the results at the various loads and again this may be due to increased influence of tribo-chemical effects as stated above. At low loads and at the highest sliding distance value (Fig. 2e), the wear rate for aluminium is again greater than for the other materials. The fact that this difference in the wear rate compared to the other materials diminishes at higher loads may be due again to a combination of differential work hardening effects [9] and formation of wear resistant oxide scales on the surface. 4.2. Tribo-chemical effects It is clear that tribo-chemical effects may affect the wear process at higher applied loads, particularly when the ridges may become in contact with the counterface. The fact that the wear rate does not increase with further increases in applied load for aluminium (Fig. 2a) at very high loads of 4 and 5 N may be due to a tribo-chemical effect on the surface as suggested above. In this case, there may be some oxide formed on the surface leading to enhanced wear resistance. It should be noted that the classical relationship between wear and applied load [10] shows distinctive transitions with a peak in the wear rate at intermediate loads attributed to the formation of an oxide film on the surface. If this
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
process occurs due to ridge formation in micro-abrasion, then significant reductions in the wear rate may occur with increased frictional heating between the ridges and the counterface materials, due to the formation of protective oxide on the surface. A possible mechanism for this process is outlined below. The major difference between this work and the earlier work of Welsh [10] is the presence of the aqueous environments in this study. This will almost certainly mean that the extent of oxidation is not as high as in the dry environment. Although distilled water is the medium used for the slurry, corrosion of the materials may play a part in the
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micro-abrasion processs. The tribo-chemical film which forms on the surface may be hydrated; however, there is insufficient evidence at this stage to speculate on the properties of such films formed in the contact. Further work will be to investigate the characteristics of such films formed on the various materials at high loads. 4.3. Micro-abrasion regimes and rationale for construction of maps To date, various micro-abrasion mechanisms have been proposed [6,7]:
Fig. 4. Schematic diagram of proposed micro-abrasion mechanisms based on the results.
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(i) 3-body rolling (where particles move in a non-directional fashion. This tends to occur at low loads and for high concentrations of particles) [6]. (ii) 2-body grooving or ridging (where particles adhere to one surface and cause formation of distinctive groove or ridge life features on the surface. This tends to occur at higher loads and low concentrations of particles) [4]. (iii) A mixed regime intermediate between these two mechanisms [4]. (iv) A regime in which ridge formation dominates the wear mechanism and the wear rate tends to reduce with further increases in applied load [5]. The latter wear mechanism where the particles are entrained in the grooves results in a situation where the wear is a two-body process and the wear rate reduces with further increases in applied load (the onset to this regime may be a situation where the load is supported hydrodynamically by the aqueous layer between the two surfaces). With further increases in load, frictional heating may occur on the surface due to the contact of the ridges with the counterface material and in this case the wear process may consist of formation and removal of oxide on the surface. Here, the wear rate is expected to decrease with increased load. In both these cases, the wear is a two-body process, albeit entirely different mechanistically than in regime (ii) above. At progressively higher loads, we might expect to see a transition from regime (iv) to (i) due to wear of the oxide film on the ridges, freeing up the particles in the ridges to cause wear by a rolling mechanism. Hence, we can describe the various regimes as a function of increasing applied load as follows (Figs. 4 and 5) by combining the mechanisms described in various studies carried out to date:
3–2 body 1. 3-body rolling (particles move in a non-directional motion): wear increases with increasing applied load. Particles can roll between the surfaces [4]. 2. Mixed (combination between (i) and (iii)) (some ridge formation occurs): wear increases with further increases in applied load due to the fact that some of the particles can move within the contact and others embed in the surface leading to groove or ridge formation [6]. 3. 2-body grooving (formation of ridges due to particles adhering to surface): wear occurs as a grooving mechanism and increases with further increases of applied load [4]. 2-body-r (i) 2-body wear (wear dominated by ridge formation [5] and interaction between ridges and the counterface material). Wear decreases with further increases in applied load due to (a) particles becoming entrained in the ridges and (b) frictional heating between the ridges and the counterface material at higher loads leading to protective layers of oxide on the surface. We can speculate that at high loads, the ridges would be worn down further resulting in a transition to regime (i) (Fig. 5); however, frictional heating would modify the wear process at this point and thus this regime may only occur in a very narrow range of applied loads compared to that observed at lower loads due to the reduction in substrate hardness which favours the transition to grooving wear. The above observations clearly have implications for predictive modelling of micro-abrasion. Any mathematical model of the process must take into account the complexity of the possible regimes, and the tribo-chemical
Fig. 5. Schematic diagram of transitions between the wear mechanisms as a function of increasing applied load.
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
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Fig. 6. Micro-abrasion mechanism map for steel.
interactions which must inevitably occur on the surface due to frictional heating effects. To date, the latter effects have not been addressed in detail for the micro-abrasion process. The wear maps for the various materials, were constructed by identifying regimes of interaction based on Fig. 5. The 3–2-body-r and 2-body-r regimes are distinguished using various symbols and an approximate outline of the boundary between the two regimes is given on each map. The maps show that ridge formation, not surprisingly for steel (Fig. 6) is significantly less than for aluminium (Fig. 7). This is attributed to the significantly higher hardness of the steel (and generally higher work hardening rate [9] compared with the
aluminium. In fact, the “2-body-r” regime above dominates the wear process for aluminium for the longest sliding distances over the test regime studied and over a large range of applied loads. The wear map for the copper (Fig. 8) shows there is less ridge formation than for aluminium (Fig. 7) and this is thought to be due to the higher hardness of the copper compared to the aluminium. The “2-body-r” regime appeared to preferentially occur at lower loads at higher sliding distances for both aluminium and copper and is consistent with the above results which indicate that the wear volume largely increases with sliding distance. That this regime does not shift to lower loads with increased sliding distance for steel
Fig. 7. Micro-abrasion mechanism map for aluminium.
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M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
Fig. 8. Micro-abrasion mechanism map for copper.
(Fig. 6) is attributed to the higher hardness of this material. Further work will be to investigate the effects of other variables on the transitions between these regimes in addition to examining in more detail the possible tribo-chemical effects involved in such processes.
5. Conclusions (i) The effects of applied load and sliding distance on the micro-abrasion of pure metals and steels have been investigated. (ii) The results have suggested a basis for extension of the range of possible micro-abrasion mechanisms which have been proposed to date. (iii) Micro-abrasion mechanism maps have been constructed as a function of sliding distance and applied load showing significant differences between locations of three-body and two-body abrasion regimes for pure metals and steels.
References [1] V. Thompson, H.E. Hintermann, L. Chollet, The determination of composition depth profiles using spherical erosion and scanning auger electron microscopy, Surf. Technol. 8 (1979) 421. [2] J. Valli, J. Palojarvi, U. Makela, Technical Research Centre of Finland, Report 435, 1985. [3] K.L. Rutherford, I.M. Hutchings, Theory and application of a micro-scale abrasive wear test, J. Test. Eval. 25 (2) (1997) 250. [4] R.I. Trezona, D.N. Allsopp, I.M. Hutchings, Transition between two-body and three-body abrasive wear-influence of test conditions in the micro-scale abrasion test, Wear 205 (1999) 225. [5] F.J. Buchanaan, P.H. Shipway, Micro-abrasion—a simple method to assess surface degradation of UHMPE following sterilization and ageing, Biomaterials 23 (1) (2002) 93. [6] J.C.A. Batista, A. Matthews, C. Godoy, Surf. Coat. Technol. 1137 (2001) 142. [7] E. Rabinowicz, L.A. Dunn, P.G. Russell, A study of abrasive wear under three body conditions, Wear 4 (1961) 345. [8] J.F. Archard, Contact and rubbing of flat surfaces, J. Appl. Phys. 24 (1953) 981. [9] S. Kalpakjian, Manufacturing Processes for Engineering Materials, second ed., Addison-Wesley, Reading, MA, 1991. [10] N.C. Welsh, Philos. Trans. R. Soc. London, Ser. A 257 (1965) 51.
Micro-abrasion transitions of metallic materials M.M. Stack∗ , M. Mathew Department of Mechanical Engineering, University of Strathclyde, 75 Montrose St., Glasgow, G1 1XJ, UK
Abstract Significant progress has been made in the understanding of micro-abrasion transitions of various materials in recent years, where abrasion is caused by particle sizes which are typically less than 10 m. Research has shown reasonable consistency on effects of applied load and sliding distance, for studies carried out in various laboratories. In addition, attempts have been made to construct abrasion “diagrams” showing the transitions between the various regimes as a function of the above parameters. A puzzling aspect of the results to date, however, is the effect of “ridge” development on the micro-abrasion wear pattern. This ridge formation leads to a reduction in wear rate because the size of abrading particles, which account for the three-body effect, is less than the size of the ridge developed as a result of the wear process. The particles thus become lost in the ridge and cause no further abrasion. In this paper, the transition to ridge formation, as a function of sliding distance and load, is described for a range of pure metals of different hardness, using a TE-66, Plint micro-abrasion test rig. Tally-surf and optical microscopy techniques were used to measure the wear rate. Micro-abrasion maps were constructed showing differences in the transition boundaries between the wear regimes, and the implications of such results, for predictive modelling of micro-abrasion, are addressed in this paper. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Micro-abrasion; Pure metals; Wear mechanism maps; Wear regimes
1. Introduction Significant developments have been made in the understanding of micro-abrasion mechanisms of materials in recent years. There is now a considerable body of literature on the performance of a range of materials in such conditions from PVD coatings to paints [1–6]. Mechanisms of micro-abrasion have been identified based on the observed deformation in the various conditions. There is, however, still some confusion on both the possible number of abrasion regimes and the nomenclature used to define such transitions. The three-body abrasion mechanism between particles entrained in the contact may revert to a two-body process if the particles adhere rigidly to one of the surfaces [4]. On the other hand, if the particles disappear between ridges which form during the process, the wear mechanism must involves only the surfaces in contact and this again is a two-body process, albeit entirely different in nature to the latter mechanism [5]. Wear mechanism maps have been produced for such processes [6,7] and again there are some differences in the terminology applied by investigators. Some mapping approaches ∗ Corresponding author. Tel.: +44-141-5483754; fax: +44-141-5525105. E-mail address: [email protected] (M.M. Stack).
have employed the established nomenclature, i.e. three-body, two-body abrasion to describe the process, extending it to distinguish between rolling and grooving wear [4]. Other investigators have differentiated between ridge formation and three-body abrasive wear [5]. It has been pointed out that, for such a complex process, the terminology used to describe the regimes should be carefully chosen in order to describe accurately the surface damage mechanisms [6]. In this paper, a series of tests were carried out for the effect of sliding distance and load on the micro-abrasion of three metallic materials, aluminium, copper and steel. The results are interpreted in terms of different regimes of behaviour and a mechanistic description of the various regimes is proposed. Erosion maps, as a function of sliding distance and load, are presented for each of the materials and the differences between the maps are explained in terms of the tribological properties of the various materials.
2. Experimental details Micro-abrasion tests were performed with a commercially available apparatus, the TE-66, micro-abrasion tester (Fig. 1) (Plint, UK), based on established design (3–4). The details of the experimental rig are as follows. A 25 mm ball is located between two-coaxial shafts each carried in support bearing.
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00204-7
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
Fig. 1. Schematic diagram of experimental apparatus.
One shaft is driven by a variable speed dc geared motor. A batch counter is provided to measure and control the number of shaft revolutions. A peristaltic pump head is connected to the other end of the shaft and this is used for providing slurry feed to the contact. The test sample is clamped onto a platform, which is fitted to the pivoted L-shaped arm. This arm is rotated around its pivot until the sample comes into contact with the ball. The beam is in balance when the samples are just in contact and the load is applied by adding dead weights to a cantilever arm. Each data point represents one experimental result; the error in the results was ±12% based on the average of results from three consecutive tests. There were no measurements taken for the steel balls (although it is acknowledged that wear of the steel balls will also be a major issue in such a test). On average, a new ball was used after eight tests. The slurry is stored in a container that can be agitated on a laboratory magnetic stirrer. The slurry is pumped by the integral peristaltic pump. It is fed to a position just above the contact point and collected in a waste tray underneath. The specification of the apparatus and experimental details are shown in Tables 1 and 2. The arm, which holds the sample, can be moved horizontally in order that several tests on a single sample specimen can be carried out. This configuration has the Table 1 Specification of the micro-abrasion apparatus Name and model Supplier Load range (N) Ball diameter (mm) Ball speed range (rpm) Pump feed rate (ml h−1 ) Software used
Micro-abrasion tester, TE-66 (Plint), UK 0.05–5 25 30–150 Up to 60 (based on 0.5 m bore) COMPEND 2000
Table 2 Experimental details Sample materials Ball materials Speed (rpm) Load (N) Sliding distance (rev) Slurry
Mild steel (2% carbon), aluminium, copper 304 stainless steel (supplied by SKF bearings) 100 1–5 150, 375, 750, 1500, 3000 or (11775, 29438, 58875, 117750, 235500 mm) Slurry composition: silicon carbide (4 m diameter) with distilled water (concentration of 0.025 g cm−1 )
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advantage of accurate control of both normal load (to an accuracy of ±0.01 N) and sliding speed. The sample is then removed from the apparatus and the diameter of the resulting abrasion scars are measured with profile projector and optical microscope. The material of the ball used was 304 stainless steel (of hardness 721 VHN at 5 kg load, 388 kg mm−2 ). Tests were conducted on mild steel (0.2% C), aluminium and copper. The hardness (VHN) of these materials were mild steel, 258; copper, 48 and aluminium, 38 at 3 kg load (138.69, 25.88 and 20.49 kg mm−2 , respectively). The sample surfaces were ground and polished by conventional metallographic methods before testing. Following the test, the worn samples were examined by optical and scanning electron microscopy and by a profilometer. The wear volume was calculated using the standard technique for measuring the wear scar of spherical geometry [3], i.e. the geometry of the wear scar is assumed to reproduce the spherical geometry of the ball, and the wear volume (V) may then be calculated by measurement of either the crater diameter (b) or its depth (h): V ≈
Πb4 , 64R
V ≈ Πh2 R,
for b R for h R
(1) (2)
3. Results 3.1. Effect of sliding distance on applied load on the micro-abrasion rates The results on the variation of wear rate with applied load (Fig. 2a) show that at a sliding distance of 11 775 mm, there was a peak in the wear rate for aluminium; the peak shifted to higher loads for copper (it should be noted that the groove diameter, on which the wear volume measurements were calculated, was measured using profilometry. Optical microscopy was used to confirm these measurements). The micro-abrasion rates of the steel were significantly less than those for the pure metals even at the highest loads. There was very little difference between the wear rates of all three materials at the highest loads of 5 N. As the sliding distance was increased to 29 438 mm (Fig. 2b), the characteristic peak was again observed for the aluminium; for the steel this tended to occur at higher loads. However, at sliding distances of 58 875 mm (Fig. 2c), a transition appeared to have occurred. There appeared to be a continuous increase in wear rate for aluminium; this behaviour now was similar to that of steel. However, for copper, the peak in the wear rate versus applied load occurred at lower loads than that of the other materials. At significantly higher sliding distances, 11 750 mm (Fig. 2d), there was no definite pattern to the results; however, the peak in the wear rate for aluminium appeared to occur at lower loads than at half the sliding distance
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(Fig. 2c). At the highest sliding distance, 235 500 mm, the wear rate of the aluminium was greatest at the lowest load of 1 N; at higher loads there was little differences between the wear rates of the materials. There was some evidence that the peak in the wear rate for the steel had shifted to lower loads compared to that observed at roughly half the sliding distance (Fig. 2c).
3.2. Scanning electron micrographs of the abraded surfaces The micrographs (Fig. 3) show various regimes of micro-abrasion for the various materials. At sliding distances of 58 875 mm (Fig. 3a) for steel, there was significant plastic deformation on the surfaces, some evidence of grooving was
Fig. 2. Variation of wear rate with applied load for the following sliding distances: (a) 11 775 mm; (b) 29 438 mm; (c) 58 875 mm; (d) 11 750 mm; (e) 235 500 mm.
M.M. Stack, M. Mathew / Wear 255 (2003) 14–22
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Fig. 2. (Continued ).
Fig. 3. Scanning electron micrographs of abraded surfaces: (a) mild steel, 1 N, 58 875 mm; (b) copper, 2 N, 235 500 mm; (c) copper as in (b) shows ridge formation at high magnification; (d) aluminium, 2 N, 235 500 mm; (e) aluminium as in (d) shows ridge formation at high magnification; (f) aluminium as in (d) shows oxide film at low magnification.
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also observed. At higher sliding distances of 235 500 mm and at higher loads for copper (Fig. 3b) there was evidence of ridge formation. At higher magnifications (Fig. 3c) there was also evidence of ridge formation due to grooving wear. For aluminium, at the highest sliding distance of test regime (Figs. 3d and e) a similar mechanism was observed. However, there also appeared to be some evidence of oxide scale on the surface in these conditions (Fig. 3f). 4. Discussion 4.1. Mechanisms of micro-abrasion of the various materials as a function of applied load and sliding distance The volume loss due to wear, V, is based on Archard’s equation [8] and can be written as follows [9]: V = κSN
(3)
where κ is the constant, S the sliding distance and N the normal load. Clearly, the results (Fig. 2) show that this equation is only valid in a minority of cases for the materials tested because the wear rate is observed to reduce with increased load in many of the tests carried out in this apparatus design [4,5]. However, this pattern, in which there is a transition to ridge formation above a critical load has been seen in previous work in this area [4,5]. That such transitions occur at different loads for a range of pure metals had not been observed. It should be noted that the results are dependent on the approximation to semi-spherical wear scars of the raw data. The scars may on occasion appear more elliptical than semi-spherical (depending on the corresponding wear of the counter face) and this is a major issue for determination of micro-abrasion rates using such an apparatus. The three-body to two-body transition has been attributed to the embedment of particles in the surface which results in a grooving mechanism [4,5]. The resultant decrease in wear rate above this transition load may be due to entrainment of particles in the grooves leading to another kind of two-body abrasion mechanism where the abrading surfaces are in contact. Hence, there are a number of possible mechanisms of micro-abrasion and this is addressed further below. It is interesting that although the micro-abrasion rate largely increased with sliding distance (Fig. 2) such increases were not linear due to the transitions observed above, and again this points to deviation from the Archard’s equation, Eq. (3). It has been pointed out that as sliding distance increases, the area of the contact will change [4] (and generally increase) thus resulting in a possible reduction of the load per abrasive particle with increasing sliding distance. Frictional heating leading to oxide formation will also change the characteristics of the area under contact; thus, there are various reasons why deviation from linear behaviour would be expected.
It is interesting that at low sliding distances (Fig. 2a) the critical load at which the peak in the wear rate occurs is observed, shifts to higher values for copper compared to aluminium. This is likely to be due to the lower hardness of the aluminium compared to the copper, i.e. 21 kg mm−2 compared to 26 kg mm−2 , which would result in ridge formation occurring at lower loads for the softer material. The fact that the micro-abrasion rates approach each other at the high loads is possibly due to a range of factors; differential work hardening rates [9] and greater influence of tribo-chemical mechanisms in such conditions. In the latter case, there may be less difference in the wear rate of such oxides than for the substrate materials. For longer sliding distances (Fig. 2b), the peak in the wear rate for aluminium at low loads indicates that hardness is still playing a role in determining the transition to “ridge-dominated” micro-abrasion behaviour. There is some evidence that the peak for the steel occurs at a significantly higher load, 4 N, and this is attributed to its significantly higher hardness value (139 kg mm−2 ) and the higher work hardening rate [2]. At intermediate sliding distances (Fig. 2c), the peak in the wear rate for the aluminium and steel shifts to significantly higher loads; however, the peak for the copper remains at intermediate loads. For steel, this may be consistent with the fact that the work hardening rate of this material is higher than for copper. However, the reason why the peak should shift to higher values for aluminium is unclear (there is evidence for higher work hardening rates of steel compared to copper in the literature; the picture for aluminium versus copper is less clear [9]). However, for longer time sliding distances, there appears to be very little difference between the results at the various loads and again this may be due to increased influence of tribo-chemical effects as stated above. At low loads and at the highest sliding distance value (Fig. 2e), the wear rate for aluminium is again greater than for the other materials. The fact that this difference in the wear rate compared to the other materials diminishes at higher loads may be due again to a combination of differential work hardening effects [9] and formation of wear resistant oxide scales on the surface. 4.2. Tribo-chemical effects It is clear that tribo-chemical effects may affect the wear process at higher applied loads, particularly when the ridges may become in contact with the counterface. The fact that the wear rate does not increase with further increases in applied load for aluminium (Fig. 2a) at very high loads of 4 and 5 N may be due to a tribo-chemical effect on the surface as suggested above. In this case, there may be some oxide formed on the surface leading to enhanced wear resistance. It should be noted that the classical relationship between wear and applied load [10] shows distinctive transitions with a peak in the wear rate at intermediate loads attributed to the formation of an oxide film on the surface. If this
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process occurs due to ridge formation in micro-abrasion, then significant reductions in the wear rate may occur with increased frictional heating between the ridges and the counterface materials, due to the formation of protective oxide on the surface. A possible mechanism for this process is outlined below. The major difference between this work and the earlier work of Welsh [10] is the presence of the aqueous environments in this study. This will almost certainly mean that the extent of oxidation is not as high as in the dry environment. Although distilled water is the medium used for the slurry, corrosion of the materials may play a part in the
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micro-abrasion processs. The tribo-chemical film which forms on the surface may be hydrated; however, there is insufficient evidence at this stage to speculate on the properties of such films formed in the contact. Further work will be to investigate the characteristics of such films formed on the various materials at high loads. 4.3. Micro-abrasion regimes and rationale for construction of maps To date, various micro-abrasion mechanisms have been proposed [6,7]:
Fig. 4. Schematic diagram of proposed micro-abrasion mechanisms based on the results.
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(i) 3-body rolling (where particles move in a non-directional fashion. This tends to occur at low loads and for high concentrations of particles) [6]. (ii) 2-body grooving or ridging (where particles adhere to one surface and cause formation of distinctive groove or ridge life features on the surface. This tends to occur at higher loads and low concentrations of particles) [4]. (iii) A mixed regime intermediate between these two mechanisms [4]. (iv) A regime in which ridge formation dominates the wear mechanism and the wear rate tends to reduce with further increases in applied load [5]. The latter wear mechanism where the particles are entrained in the grooves results in a situation where the wear is a two-body process and the wear rate reduces with further increases in applied load (the onset to this regime may be a situation where the load is supported hydrodynamically by the aqueous layer between the two surfaces). With further increases in load, frictional heating may occur on the surface due to the contact of the ridges with the counterface material and in this case the wear process may consist of formation and removal of oxide on the surface. Here, the wear rate is expected to decrease with increased load. In both these cases, the wear is a two-body process, albeit entirely different mechanistically than in regime (ii) above. At progressively higher loads, we might expect to see a transition from regime (iv) to (i) due to wear of the oxide film on the ridges, freeing up the particles in the ridges to cause wear by a rolling mechanism. Hence, we can describe the various regimes as a function of increasing applied load as follows (Figs. 4 and 5) by combining the mechanisms described in various studies carried out to date:
3–2 body 1. 3-body rolling (particles move in a non-directional motion): wear increases with increasing applied load. Particles can roll between the surfaces [4]. 2. Mixed (combination between (i) and (iii)) (some ridge formation occurs): wear increases with further increases in applied load due to the fact that some of the particles can move within the contact and others embed in the surface leading to groove or ridge formation [6]. 3. 2-body grooving (formation of ridges due to particles adhering to surface): wear occurs as a grooving mechanism and increases with further increases of applied load [4]. 2-body-r (i) 2-body wear (wear dominated by ridge formation [5] and interaction between ridges and the counterface material). Wear decreases with further increases in applied load due to (a) particles becoming entrained in the ridges and (b) frictional heating between the ridges and the counterface material at higher loads leading to protective layers of oxide on the surface. We can speculate that at high loads, the ridges would be worn down further resulting in a transition to regime (i) (Fig. 5); however, frictional heating would modify the wear process at this point and thus this regime may only occur in a very narrow range of applied loads compared to that observed at lower loads due to the reduction in substrate hardness which favours the transition to grooving wear. The above observations clearly have implications for predictive modelling of micro-abrasion. Any mathematical model of the process must take into account the complexity of the possible regimes, and the tribo-chemical
Fig. 5. Schematic diagram of transitions between the wear mechanisms as a function of increasing applied load.
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Fig. 6. Micro-abrasion mechanism map for steel.
interactions which must inevitably occur on the surface due to frictional heating effects. To date, the latter effects have not been addressed in detail for the micro-abrasion process. The wear maps for the various materials, were constructed by identifying regimes of interaction based on Fig. 5. The 3–2-body-r and 2-body-r regimes are distinguished using various symbols and an approximate outline of the boundary between the two regimes is given on each map. The maps show that ridge formation, not surprisingly for steel (Fig. 6) is significantly less than for aluminium (Fig. 7). This is attributed to the significantly higher hardness of the steel (and generally higher work hardening rate [9] compared with the
aluminium. In fact, the “2-body-r” regime above dominates the wear process for aluminium for the longest sliding distances over the test regime studied and over a large range of applied loads. The wear map for the copper (Fig. 8) shows there is less ridge formation than for aluminium (Fig. 7) and this is thought to be due to the higher hardness of the copper compared to the aluminium. The “2-body-r” regime appeared to preferentially occur at lower loads at higher sliding distances for both aluminium and copper and is consistent with the above results which indicate that the wear volume largely increases with sliding distance. That this regime does not shift to lower loads with increased sliding distance for steel
Fig. 7. Micro-abrasion mechanism map for aluminium.
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Fig. 8. Micro-abrasion mechanism map for copper.
(Fig. 6) is attributed to the higher hardness of this material. Further work will be to investigate the effects of other variables on the transitions between these regimes in addition to examining in more detail the possible tribo-chemical effects involved in such processes.
5. Conclusions (i) The effects of applied load and sliding distance on the micro-abrasion of pure metals and steels have been investigated. (ii) The results have suggested a basis for extension of the range of possible micro-abrasion mechanisms which have been proposed to date. (iii) Micro-abrasion mechanism maps have been constructed as a function of sliding distance and applied load showing significant differences between locations of three-body and two-body abrasion regimes for pure metals and steels.
References [1] V. Thompson, H.E. Hintermann, L. Chollet, The determination of composition depth profiles using spherical erosion and scanning auger electron microscopy, Surf. Technol. 8 (1979) 421. [2] J. Valli, J. Palojarvi, U. Makela, Technical Research Centre of Finland, Report 435, 1985. [3] K.L. Rutherford, I.M. Hutchings, Theory and application of a micro-scale abrasive wear test, J. Test. Eval. 25 (2) (1997) 250. [4] R.I. Trezona, D.N. Allsopp, I.M. Hutchings, Transition between two-body and three-body abrasive wear-influence of test conditions in the micro-scale abrasion test, Wear 205 (1999) 225. [5] F.J. Buchanaan, P.H. Shipway, Micro-abrasion—a simple method to assess surface degradation of UHMPE following sterilization and ageing, Biomaterials 23 (1) (2002) 93. [6] J.C.A. Batista, A. Matthews, C. Godoy, Surf. Coat. Technol. 1137 (2001) 142. [7] E. Rabinowicz, L.A. Dunn, P.G. Russell, A study of abrasive wear under three body conditions, Wear 4 (1961) 345. [8] J.F. Archard, Contact and rubbing of flat surfaces, J. Appl. Phys. 24 (1953) 981. [9] S. Kalpakjian, Manufacturing Processes for Engineering Materials, second ed., Addison-Wesley, Reading, MA, 1991. [10] N.C. Welsh, Philos. Trans. R. Soc. London, Ser. A 257 (1965) 51.