Wear-mode mapping for the micro-scale abrasion test

Wear 255 (2003) 23–29

Communication

Wear-mode mapping for the micro-scale abrasion test K. Adachi a,∗ , I.M. Hutchings b a

Laboratory of Tribology, School of Mechanical Engineering, Tohoku University, Aramaki-Aza-Aoba 01, Sendai 980-8579, Japan b Department of Engineering, Institute for Manufacturing, University of Cambridge, Cambridge CB2 1RX, UK

Abstract The objective of this study was to develop a theoretical model and associated wear-mode map to identify the regimes in which two-body abrasion (grooving abrasion) and three-body abrasion (rolling abrasion) dominate in the micro-scale abrasive wear test (also known as the ball-cratering abrasion test). The critical condition for the transition between two-body and three-body abrasion was determined from a continuum mechanics model for the penetration of the abrasive particles into the surfaces of the ball and the specimen, coupled with considerations of equilibrium. Micro-scale abrasion tests were performed with different combinations of ball and specimen materials, under different test conditions such as abrasive concentration and load, and a wear-mode map has been produced which defines the regimes of abrasive particle motion. The map is plotted between two dimensionless groups as vertical and horizontal axes: the hardness ratio between the ball and the specimen, and a newly introduced parameter which represents the severity of contact. Experimental data generated in this work and also taken from previous studies show that the map represents behaviour in the micro-scale abrasion test well, for a wide range of ball and specimen counterface materials. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Micro-scale abrasion; Ball cratering; Wear-mode map; Two-body abrasion; Three-body abrasion

1. Introduction The micro-scale abrasion test (also known as the ball-cratering abrasive wear test), in which a ball is rotated against a specimen in the presence of a slurry of fine abrasive particles, is a useful technique for evaluating the wear resistance of surface engineered components [1,2]. The method is currently being assessed as a standard test for the wear resistance of coatings in an EU-funded project [3]. Since wear is a system response and is not a material property, the wear resistance of a material can vary over a wide range if different wear mechanisms are induced by different test conditions [4,5]. For repeatable and reproducible measurements of wear resistance in a standard wear test, the test conditions must be carefully controlled so that the wear mechanism is predictable and reproducible. In the micro-scale abrasion test, which employs free abrasive particles, two distinct wear modes have been identified [6–8]. The dominant process is controlled by the nature of the abrasive particle motion in the contact region between the ball and the specimen. If the particles do not move relative to the ball surface, but act as fixed inden∗ Corresponding author. Tel.: +/fax: +81-22-217-6956. E-mail address: [email protected] (K. Adachi).

ters moving across the specimen, a series of fine parallel grooves is produced on the specimen surface. This leads to so-called grooving wear or two-body abrasion. If on the other hand the abrasive particles roll between the two surfaces, multiple indentations with no evident directionality are produced in a process known as rolling abrasion or three-body abrasion [6–13]. The dominant wear modes in the micro-scale abrasion test have been reported to be influenced by applied load [6], volume fraction of abrasive in the slurry [6], abrasive material [6], materials of both ball and specimen [7], and the surface condition of the ball [8]. In order to use the ball-cratering method as a standard test to evaluate abrasive wear resistance, it is essential to ensure that either two-body or three-body abrasion occurs in the test; there is some evidence that three-body abrasion leads to more reproducible test results [14]. The purpose of this work was therefore to formulate a theoretical model for particle motion in the test, and produce an associated wear-mode map, plotted using dimensionless parameters, which shows the regimes of dominant wear mode observed in experiments carried out for this project and also taken from previous studies. The wear-mode map can then be used to predict the conditions necessary to ensure either two-body or three-body abrasion.

0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00073-5

24

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

Fig. 2. SEM micrograph of F1200 SiC abrasive particles (from Ref. [15]).

Fig. 1. Schematic diagram showing (a) experimental apparatus and (b) details of contact interface and abrasive slurry supply.

2. Experimental methods and materials Fig. 1(a) shows a schematic diagram of the micro-scale abrasive wear tester used for this work (Plint TE-66, Phoenix Tribology, Reading, UK). The ball is clamped between coaxial shafts and driven by an electric motor. A flat specimen is mounted vertically on a pivoted arm and is loaded against the ball by a dead weight hanging from the horizontal shaft. An abrasive slurry containing silicon carbide particles suspended in distilled water is supplied to the contact interface with a syringe as shown in Fig. 1(b) in sufficient quantity to keep the contact surfaces wetted. Three sets of experiments were carried out with two different specimen materials (black polymethylmethacrylate

(PMMA) and tool steel (1.0 wt.% C, 1.2 wt.% Mn, 0.2 wt.% Si, 0.5 wt.% Cr, 0.15 wt.% V, 0.5 wt.% W, balance Fe)) and two different ball materials (polypropylene (PP) and bearing steel AISI 52100), with different abrasive slurry concentrations (ranging from 0.001 to 0.24 volume fraction) and applied loads (0.01–1 N). Additional tests were also carried out with several material combinations for the specimen and ball such as S1C H4 aluminium alloy (0.5 wt.% Si, 0.1 wt.% Cu, 0.7 wt.% Fe, 0.1 wt.% Mn, 0.1 wt.% Zn, balance Al) with bearing steel, phosphor bronze (8 wt.% Sn, balance Cu) and PP balls, PMMA with a phosphor bronze ball, and tool steel with a bearing steel ball to observe the wear modes. The materials used and their properties are listed in Table 1. In all tests the abrasive was silicon carbide (Washington Mills, UK), grade F1200 with a mean particle size of 4.3 ␮m and standard deviation of 1.4 ␮m [1]. The shape of the particles was highly irregular and angular as shown in Fig. 2 [15]. The total sliding distance used in each test was 16 m (200 ball rotations). The relative sliding speed was fixed at the low value of 0.05 m s−1 to avoid hydrodynamic effects. After wear testing the wear scars were observed by optical microscopy to identify the dominant wear mechanism.

Table 1 Properties of materials used for balls and specimens Material

Diameter (mm)

Roughness, Ra (Rmax ) (␮m)

Hardness (GPa)

Young’s modulus (GPa)

Poisson’s ratio

Ball

PP Phosphor bronze Bearing steel AISI 52100

25.0 25.4 25.4

0.334 (3.3) 0.192 (1.92) 0.35 (3.5)

0.062 1.95 9.9

1.5 120 200

0.4 0.38 0.3

Specimen

PMMA S1C H4 aluminium alloy Tool steel

0.216 0.41 7.75

3.5 70 210

0.4 0.3 0.29

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

25

Fig. 3. Optical images showing representative worn surfaces: (a) grooving; (b) multiple indentations. The specimen was PMMA abraded at a normal load of 0.2 N with silicon carbide slurry at volume fraction of: (a) 0.01; (b) 0.2. The counterbody was a hard steel bearing ball.

3. Experimental results Patterns of damage in the wear scars could be clearly classified into two types: either exhibiting parallel grooves as shown in Fig. 3(a), or multiply indented with no evident surface directionality as shown in Fig. 3(b). The former is associated with the micro-cutting or ploughing action of abrasive particles which are dragged through the contact by the ball, remaining essentially fixed to the ball surface during that process, and the latter is formed by the action of many particles rolling through the contact zone. In this work these modes will be referred to as two-body and three-body abrasion, respectively, although it has been pointed out elsewhere that this commonly used terminology can be misleading, since the presence of discrete abrasive particles as a ‘third body’ in a system does not uniquely define the ensuing motion of those particles [6,16]. Most of the wear scars could be identified as caused by either two-body or three-body abrasion, since their whole surfaces showed the features of one mode or the other, as shown in Fig. 4(a) and (b). A small number of wear scars showed two-body abrasion in the centre and three-body abrasion at the edges, as seen in Fig. 4(c); this behaviour was termed ‘mixed-mode’.

Fig. 5 shows the distributions of wear modes observed in three sets of experiments, with different combinations of ball and specimen material, as a function of load and volume fraction of abrasive in the slurry. The wear mode changed from three-body abrasion to two-body abrasion with an increase in load or a decrease in the volume fraction of abrasive. The dependence of wear mode on load and abrasive concentration was different for the different combinations of ball and specimen material. These results show that the motion of the abrasive particles in the contact region between the ball and the specimen, which is reflected by the different wear modes, is influenced not only by load and abrasive concentration, but also by the materials of the ball and specimen.

4. Critical condition for transition from three-body abrasion to two-body abrasion If the transition from three-body abrasion to two-body abrasion is assumed to be caused by a change of particle motion in the contact region from rolling to sliding, a two-dimensional model for particle motion proposed by

Fig. 4. Optical images of three representative worn surfaces showing: (a) two-body (grooving) abrasion; (b) three-body (rolling) abrasion; (c) mixed-mode abrasion. The specimen was black PMMA abraded at a normal load of 0.2 N with silicon carbide slurry at volume fraction of: (a) 0.01; (b) 0.2; (c) 0.1. The counterbody was a hard steel bearing ball.

26

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

Fig. 6. Idealised two-dimensional model proposed in Refs. [17,18].

3. The abrasive particles are sufficiently hard compared with both the ball and the specimen to remain essentially undeformed. 4. All abrasive particles are spherical in shape with diameter D. 5. The concentration of abrasive particles in the contact zone is proportional to the volume fraction of abrasive in the supplied slurry. 6. The surface roughnesses of the ball and the specimen are negligible compared with D. 4.1. Indentation depth of the abrasive particle into the surfaces of the ball and the specimen

Fig. 5. Wear modes observed as a function of load and volume fraction of abrasives on: (a) tool steel specimens with a bearing steel ball counterbody; (b) PMMA with a bearing steel ball; (c) PMMA with a PP ball.

Williams and Hyncica [17,18] suggests that this should be associated with a value of (D/h) lying above a critical value of about 1.74, where D is the particle major axis and h the separation of the two surfaces, as shown in Fig. 6. This critical condition can be analysed with a continuum mechanics model based on the following assumptions: 1. The hydrodynamic effect of the abrasive slurry is negligibly small since the sliding speed and the viscosity of the slurry would lead to a hydrodynamic film which is much thinner than h [1]. 2. The applied load is distributed over and supported by all the abrasive particles instantaneously present in the contact zone.

Fig. 7 shows the schematic diagrams of the contact region between the ball, the plane specimen and the abrasive particles (Fig. 7(a)) and the grooves formed by a single abrasive particle in the ball and the specimen surfaces under load (Fig. 7(b)). The depths of indentation into the ball db and the specimen ds can be estimated by assuming that the load P carried by the particle is supported by plastic flow in the surface at a pressure equal to its indentation hardness. So for grooving wear of the specimen, with hardness Hs , the load is supported over the front part of the spherical particle over a projected area approximately equal to ␲Dds /2 (for ds  D, a reasonable first assumption). The depth of indentation into the specimen surface is therefore given by ds = 2

P 1 πD Hs

(1)

Fig. 7. (a) Model of the contact zone between the ball, the specimen and the abrasive; (b) magnified view showing indentation by a single abrasive particle.

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

27

A similar expression can be derived for the depth of indentation db into the ball with hardness Hb . The separation of the ball and specimen surfaces, h, is equal to (D − ds − db ) and hence
 2P 1 2P 1 h=D− =D− + (2) πD Hb Hs πDH where 1 1 1 = +  H Hb Hs

(3)

This model assumes that the motion of the particle causes grooves in both the specimen and the ball surfaces. If, however, the particle rolls between the surfaces forming multiple indentations, then the normal load will be supported over twice the area of the spherical particles and the separation of the surfaces will be greater, and given by P h=D− (4) πDH From assumption (5) above, the total number of particles N within the interaction area A (the wear scar area) between the ball and the specimen is given by Acv N= (5) πD2 where v is the volume fraction of abrasive in the slurry, and c a constant of proportionality. The load P carried by each abrasive particle is therefore P=

πWD2 Acv

(6)

Fig. 8. Schematic model showing (a) the interaction area and (b) the magnified view.

where W is the applied load between the ball and the specimen. From Eqs. (2) and (6) the separation of the surfaces h is thus
 2W h=D 1− (7) AcvH The critical condition for the transition from grooving to rolling motion, at a certain value of (D/h), is therefore expected to occur at a critical value of (W/AvH ). The dimensionless group (W/AvH ) contains quantities which are all experimentally measurable, and can be termed the ‘severity of contact’, with the symbol S. The condition for the transition can then be written as W S= = S∗ (8) AvH

Fig. 9. Optical image of wear scar on a black PMMA specimen after a very short sliding test of one ball revolution. The contact areas calculated from Hertzian theory and Eq. (9) are marked.

28

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

It is suggested that the wear mode will change from threebody (rolling) abrasion to two-body (grooving) abrasion when S exceeds the threshold value S∗ . 4.2. Estimation of the interaction area A between the abrasive particles and the surfaces of the ball and the specimen Fig. 8 shows the schematic diagram of the interaction area between the abrasive particles and the surfaces of the ball and the specimen (Fig. 8(a)) and a more detailed picture in which the radius of interaction a and the Hertzian contact radius a are shown (Fig. 8(b)). If it is assumed that the macroscopic deformation of the ball and specimen is elastic, and that the interaction area A is the region over which the separation of the ball and the specimen is less than the diameter of the abrasive particles, then A is given by

contact S defined by Eq. (8) (with A computed from Eq. (9)) and of the ratio of the hardnesses of the specimen and the ball (Hs /Hb ). The wear mode dominant in the micro-scale abrasion test is evidently a function of both these quantities, and the condition for rolling abrasion is found empirically to be given by
β W Hs S= ≤α (10)  AvH Hb

where a is the radius of the Hertzian contact area, R the radius of the ball and D the diameter of the abrasive particle. Fig. 9 shows an optical image of the wear scar formed on a specimen after a very short abrasion test with only one ball revolution, on which is marked the contact area predicted from Hertzian theory (diameter 2a) and that derived from Eq. (9) (diameter 2a ), for particles with a mean diameter of 4.3 ␮m. The good agreement with the latter value suggests that Eq. (9) provides a sufficiently accurate estimate of the true interaction area for the purposes of this work.

where α and β are empirical constants related to the critical value of (D/h) and to the constant of proportionality c. For these data, α ≈ 0.0076 and β ≈ −0.49. This empirical relationship, which appears to be valid over a remarkably wide range of experimental conditions, should allow the dominant mode of deformation in the micro-scale abrasion test to be predicted. The threshold value for the transition from three-body abrasion to two-body abrasion evidently depends on the hardness ratio between the ball and the specimen. The reason for this dependence is unclear, but it probably reflects an influence of hardness ratio on the critical value of (D/h) or on the ratio c. The data in Fig. 10 show sharp transitions between the two modes of particle motion, and very few cases where traces of both types of motion can be detected in the wear scar (mixed-mode wear). The change in ball–specimen separation associated with the transition from grooving to rolling motion, as described by Eqs. (2) and (4), may be responsible for this behaviour; once some particles start to roll, they will tend to increase the separation of the surfaces, further decreasing the value of (D/h) and leading to yet more rolling motion.

5. Wear-mode mapping and discussion

6. Conclusions

Fig. 10 summarises wear modes observed in this work and also reported in previous studies carried out predominantly with silicon carbide abrasive but also with diamond and alumina [6–8], plotted as a function of the severity of

A wear-mode map, which defines the regimes for which two-body abrasion (grooving wear) or three-body abrasion (rolling wear) dominate in the micro-scale abrasion test can be plotted using two dimensionless groups as axes: the hardness ratio between the ball and the specimen, and a group S which represents the severity of contact. Data plotted on such a map, obtained from micro-scale abrasion tests performed predominantly with silicon carbide abrasive but also with diamond and alumina over a wide range of conditions, show a very consistent pattern. For the three-body (rolling) mode, the following criterion is proposed:
β W Hs S= ≤ α AvH Hb

A = πa2 = π(a2 + 2RD)

(9)

Fig. 10. Wear-mode map showing wear modes observed in this work and also reported from previous studies [5–7] as a function of S and hardness ratio between the ball and the specimen.

where W is the applied load, A the interaction area (wear scar area), v the volume fraction of abrasive in the slurry and Hs and Hb the hardnesses of the specimen and the ball, respectively. For the data analysed, α and β were found to have the values α = 0.0076 and β = −0.49.

K. Adachi, I.M. Hutchings / Wear 255 (2003) 23–29

Acknowledgements The authors are very grateful to Dr. Y. Kusano at University of Cambridge for helpful discussions. This work has been partially supported by the European Commission through the Framework V Growth programme, contract number GRD1-2000-25020. References [1] K.L. Rutherford, I.M. Hutchings, A micro-abrasive wear test, with particular application to coated systems, Surf. Coat. Technol. 79 (1996) 231–239. [2] K.L. Rutherford, I.M. Hutchings, Theory and application of a micro-scale abrasive wear test, J. Test. Eval. Am. Soc. Test. Mater. (JTEVA) 25 (2) (1997) 250–260. [3] M.G. Gee, A. Gant, I.M. Hutchings, R. Bethke, K. Schiffmann, K. Van Acker, S. Poulat, Y. Gachon, J. von Stebut, Review of ball cratering or micro-abrasion wear testing of coatings, NPL Report MATC(A)62, National Physical Laboratory, Teddington, UK, 2001. [4] K. Adachi, K. Kato, N. Chen, Wear map of ceramics, Wear 203–204 (1997) 291–301. [5] K. Kato, K. Adachi, Wear mechanisms, in: B. Bhushan (Ed.), Modern Tribology Handbook, CRC Press, Boca Raton, FL, 2000, pp. 273–300. [6] R.I. Trezona, D.N. Allsopp, I.M. Hutchings, Transitions between two-body and three-body abrasive wear: influence of test conditions in the micro-scale abrasive wear test, Wear 225–229 (1999) 205–214.

29

[7] R.I. Trezona, I.M. Hutchings, Three-body abrasive wear testing of soft materials, Wear 233–235 (1999) 209–221. [8] D.N. Allsopp, R.I. Trezona, I.M. Hutchings, The effects of ball surface condition in the micro-scale abrasive wear test, Tribol. Lett. 5 (1998) 259–264. [9] N. Axén, S. Jacobson, S. Hogmark, Influence of hardness of the counterbody in three-body abrasive wear—an overlooked hardness effect, Tribol. Int. 27 (1994) 233–241. [10] D.N. Allsopp, I.M. Hutchings, Micro-scale abrasion and scratch response of PVD coatings at elevated temperatures, Wear 251 (2001) 1308–1314. [11] A. Kassman, S. Jacobson, L. Erickson, P. Hedenqvist, M. Olsson, A new test method for the intrinsic abrasion resistance of thin coatings, Surf. Coat. Technol. 50 (1991) 75–84. [12] M.H. Staia, C. Enriquez, E.S. Puchi, Influence of the heat treatment on the abrasive wear resistance of electroless Ni–P, Surf. Coat. Technol. 94–95 (1997) 543–548. [13] M.H. Staia, C. Enriquez, E.S. Puchi, D.B. Lewis, M. Jeandin, Application of ball cratering method to study abrasive wear, Surf. Eng. 14 (1) (1998) 49–54. [14] E. Rabinowicz, L.A. Dunn, P.G. Russell, A study of abrasive wear under three-body conditions, Wear 4 (1961) 345–355. [15] R.I. Trezona, Ph.D. Thesis, University of Cambridge, 2000. [16] J.D. Gates, Two-body and three-body abrasion: a critical discussion, Wear 214 (1998) 139–146. [17] J.A. Williams, A.M. Hyncica, Mechanisms of abrasive wear in lubricated contacts, Wear 152 (1992) 57–74. [18] J.A. Williams, A.M. Hyncica, Abrasive wear in lubricated contacts, J. Phys. D 25 (1992) A81–A90.