- Email: [email protected]
Ball cratering test on ductile materials
Wear 271 (2011) 770–774
Contents lists available at ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Ball cratering test on ductile materials Rafael Vieira Camerini a,b,c , Rodrigo Buchfink de Souza a,b,c , Felipe de Carli a , Altair Soria Pereira a,d , Naira Maria Balzaretti a,∗ a
Instituto de Fisica, UFRGS, Av. Bento Goncalves, 9500, Porto Alegre, RS, CP 15051, 91501-970, Brazil Jomon Indústria de Artefatos Cerâmicos Ltda, Porto Alegre, RS, Brazil c PGCIMAT, UFRGS, Av. Bento Goncalves, 9500, Porto Alegre, RS, 91501-970, Brazil d Escola de Engenharia, UFRGS, Porto Alegre, RS, Brazil b
a r t i c l e
i n f o
Article history: Received 26 June 2010 Received in revised form 23 February 2011 Accepted 25 March 2011 Available online 2 April 2011 Keywords: Micro-scale abrasion Hardness Steel Non-ferrous metals Wear testing
a b s t r a c t Micro-abrasive wear test, also called ball cratering test, is normally used to investigate the wear behavior of hard bulk materials or wear-resistant coatings. In this work, this test was used to investigate the wear behavior of metallic samples. The results showed that the embedment of abrasive particles and the formation of grooves in the crater’s surface may induce to artificially small wear coefficient when the test is applied to ductile materials. Surface grooving was observed when the normal force applied by the sphere was higher than a critical value. This critical value depends on the hardness and mechanical strength of the material being tested. © 2011 Elsevier B.V. All rights reserved.
1. Introduction The abrasive wear resistance is an important property for a wide range of applications in engineering. Ball cratering test – also called micro-abrasion test – allows the quantitative evaluation of this property even for small volumes of material [1]. The technique consists in the rotation of a sphere of radius R in contact with a polished surface under the action of a normal force (N). During the test, an abrasive suspension is dropped between the ball and the sample. As the geometry of the wear crater produced on the sample surface is related to the spherical shape of the counter-body, the crater volume (V) can be obtained from the ratio between the crater diameter and the sphere diameter. For homogeneous materials, the wear coefficient can be calculated from a simple model that is equivalent to the Archard equation [1]: V = kL N
(1)
In this equation, L is the apparent sliding distance traveled by a fixed point at the surface of the sphere and k is the wear coefficient. A detailed description of this equation can be found in Ref. [1].
The wear mode depends on the abrasive particle motion at the interface between sample and counter-body. A three-body abrasion mechanism occurs when the abrasive particles stay free and roll along the interface between the sphere and the sample surface [1]. In this case, the crater surface shows a homogeneous scar pattern and directional grooves are not observed. On the other hand, a two-body abrasion mode is present when the abrasive particles are attached to the ball or to the specimen surface, producing grooves and scratch marks. In this case, the volume of the removed material in the crater cannot be calculated directly from the starting geometry of the sphere [1]. Therefore, the choice of the testing parameters should minimize the two-body abrasion mechanism [1–7]. The allowed range for these parameters will be strongly dependent of the sample’s mechanical properties. There are several studies in the literature dealing with grooving abrasion associated to the attachment of abrasive particles to the ball [3,6]. However, the consequences of this attachment on the determination of the wear coefficient have not been investigated in detail so far. In this paper, we present results for the ball cratering test on ductile soft materials where the attachment of abrasive particles to the crater surface is expected to occur. The main goal of this study is to evaluate the effect of the abrasive particles embedment on the determination of the wear coefficient. 2. Material and methods
∗ Corresponding author. Av. Bento Goncalves, 9500, Porto Alegre, RS, CP 15051, 91501-970, Brazil. Tel.: +55 51 3308 6489; fax: +55 51 3308 7286. E-mail address: [email protected] (N.M. Balzaretti). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2011.03.013
The wear behavior of aluminum, copper, AISI 1020 steel, AISI 1045 steel and quenched AISI 1045 steel (heat treated at 900 ◦ C
R.V. Camerini et al. / Wear 271 (2011) 770–774
771
Table 1 Vickers hardness and wear coefficient of the samples calculated using Eq. (1). The hardness of the ball used as counter-body in the micro-abrasive wear tests is also included. Sample
Vickers hardness (GPa)
Wear coefficient (m2 /N)
L<1m Aluminum Copper AISI 1020 steel AISI 1045 steel AISI 1045 quenched steel Mullite Stainless steel sphere
L>1m −12
1.72 2.00 2.86 2.94 8.83
10 × 10 5 × 10−12 5 × 10−12 5 × 10−12 2 × 10−12
9 × 10−12 4 × 10−12 3 × 10−12 2 × 10−12 2 × 10−12
8.58 9.02
5 × 10−12
5 × 10−12
during 30 min, followed by cooling in mineral oil) was investigated by the ball cratering test. For comparison, the same experimental procedure was applied to a ceramic specimen of mullite. The Vickers hardness of the sphere and the samples were measured using a micro-hardness device (Shimadzu A-108) with a load of 300 g (2.94 N) (Table 1). The ball cratering wear tests were performed using the free ball/single shaft geometry (CSEM Calowear – Fig. 1), where the ball
Fig. 2. SEM image showing the morphology of the SiC abrasive powder.
is driven by friction with a shaft. The normal force is measured by a load cell positioned below the sample table, and its value is adjusted by the slope of the table, which modifies the component of the sphere’s weight perpendicular to the sample’s surface. The maximum force used in this setting usually does not exceed 0.7 N. A hard steel ball with 25.40 mm in diameter was used to perform the tests. For every sample investigated, the first crater was discarded to ensure the ball would be in the same condition for every consecutive crater. Before the abrasive tests, all samples were polished using an aqueous suspension of alumina powder (0.4 m grain size) and ultrasonic bath cleaned first in distilled water and then in ethanol for 20 min. At the surface of each sample, 11 craters were made, corresponding to sliding distances from 0.4 m to 4.0 m (5–50 rotations of the sphere with a tangential velocity of 0.05 m/s). An abrasive suspension of silicon carbide (SiC) powder in distilled water (0.75 g/cm3 ) was dropped in a rate of 1 drop per sphere’s rotation. The morphology of the abrasive particles was characterized by SEM microscopy (Fig. 2). The particle size distribution was measured by laser diffraction (Cilas 1180LD device) and the results are shown in Table 2. The normal force was kept constant during the tests and all craters in the same sample were produced under the same load. The value of normal force was between 0.265 N and 0.285 N for all tests. The exception occurred for aluminum, the softest studied material (Table 1). In this case, to reduce the scratch inside the craters (groove formation), it was used 0.220 N. Also, additional tests using 0.132 N and 0.315 N were made on AISI 1045 steel to investigate the effect of the normal force on the density of abrasive particles trapped in the craters. Images of the wear craters were obtained using a metallographic microscope with a low-magnification objective. The diameters of the craters were evaluated by software for image analysis (Optimas). A scanning electron microscope with an energy-dispersive spectrometer (SEM/EDS) was used for chemical analysis of the samples.
Table 2 Grain size distribution of the SiC abrasive powder.
Fig. 1. Ball cratering apparatus. (a) Picture of CSEM Calowear; (b) schematic representation of the experimental geometry.
Diameter at 10% Diameter at 50% Diameter at 90% Mean diameter
0.74 m 3.13 m 5.11 m 3.04 m
772
R.V. Camerini et al. / Wear 271 (2011) 770–774
Fig. 3. Volume of the crater divided by the normal force as a function of the sliding distance of a fixed point at the surface of the sphere. (a) Aluminum, (b) copper, (c) AISI 1020 steel, (d) AISI 1045 steel, (e) quenched steel (AISI 1045) and (f) mullite.
3. Results and discussion The results obtained for all the samples are shown in Fig. 3, where the ratio between the volume of the crater and the normal load during the test (V/N) is presented as a function of the sliding distance (L) traveled by the sphere. The dashed lines correspond to linear fittings. For mullite and tempered steel, the hardest investigated materials, the results can be well described by a single straight line. For all other samples, a single linear fitting was not enough to give a good representation of the wear behavior for the whole range of sliding distances. In these cases, two straight lines, one up to L ≈ 1 m, and other for larger sliding distances, can be used to describe quite well the experimental results. Systematically, the slope of the line for larger distances was smaller, corresponding to a smaller wear coefficient (k). Table 1 shows the estimated values of k obtained for both straight lines for each sample. As can be seen, the wear coefficients of the metallic samples are of the same order of magnitude than the ceramic sample and the tempered steel. On the other hand, the hardness values of the ceramic and tempered steel sample are three to four times higher than of the softer samples. A possible explanation for the anomalous low wear coefficient values obtained for the softer materials is related to the embedment of the abrasive particles at the surface of the craters, inhibiting the abrasion of the metal surface. The attachment of abrasive particles in relatively soft samples has been reported by Bello and Wood [8], Acselrad et al. [9] and Bose
and Wood [10]. None of these studies, however, investigated the effects of the attachment on the wear coefficient of these materials. The nonlinear behavior depicted in Fig. 3 for soft metallic materials is normally related to a transition from 3 body to 2 body abrasion wear modes [6,11]. Braga et al. [11] investigated the behavior of steels under micro abrasion test using a ball cratering equipment with fixed-ball. Their study corroborates the idea that this nonlinear behavior should be a consequence of two phenomena: (a) the transition between wear modes and (b) the low resolution of the edge of the crater, required during the evaluation of the volume of the crater. The results shown in the present work pointed out to a third reason which contributes to the nonlinear behavior of the abrasion wear. The attachment of abrasive particles inside the craters may explain why the soft and ductile materials exhibit an apparently anomalous low abrasive wear behavior. The concentration of Si at the surface of the crater can be associated to the embedment of SiC particles. Fig. 4 shows the increase of Si concentration as a function of the sliding distance traveled by the ball for AISI 1020, AISI 1045 quenched steel and copper measured by SEM/EDS analysis. As can be seen, the concentration practically saturates for sliding distances greater than ∼1 m, the sliding distance corresponding to the transition between the two straight lines in Fig. 3 for the soft materials. Fig. 4 also shows that the amount of embedded particles is different for each sample, ranging from ∼6 at% for quenched AISI 1045 steel to ∼20 at% for Copper and AISI 1020 steel.
R.V. Camerini et al. / Wear 271 (2011) 770–774
773
Fig. 5. Abrasive wear of AISI 1045 for different values of normal force: (䊉) 0.132 N, () 0.220 N, (×) 0.315 N.
Fig. 4. (a) Atomic concentration of Si in the craters as a function of the sliding distance traveled by the sphere. () AISI 1020 steel; () AISI 1045 quenched steel, () copper. (b) SEM image of a crater in AISI 1045 quenched steel. (c) EDS analysis of region (1), inside the crater. (d) EDS analysis of region (2), outside the crater.
The density of SiC particles attached on the scars of the softer steel (AISI 1020) is higher than in the harder steel. This is probably a consequence of the degree of anchoring of the abrasive particles in this material, without the complete burial of the grains. On the other hand, for the harder steel (AISI 1045), its resistance to plastic deformation leads to the formation of shallow indentations that are
not always able to promote the adhesion of the abrasive particle. For the softer metallic samples, on the contrary, the low resistance to plastic deformation allows the complete burial of abrasive particles in these materials. Thereby, despite the graph in Fig. 4a shows a high concentration of silicon carbide inside the craters on copper specimen, the presence of the grooves at the surface with an abrasive particle stuck to the end were not easily found probably because the particles were completely buried at the surface. Fig. 4b shows an image of the crater on AISI 1045 quenched steel. Fig. 4c and d shows EDS analyses of regions inside (1) and outside (2) the crater, indicating there was no Si outside the crater region. Fig. 5 compares the wear behavior of AISI 1045 steel for three different values of normal force: 0.132 N, 0.220 N and 0.315 N. For short sliding distances (when L < 1 m) the volume of craters normalized by the normal force was almost the same. However, for longer sliding distances, the slope and, consequently, the wear coefficient, decreased as the normal force increased. This phenomenon was already reported by Adachi and Hutching [12] for tool steel samples. The SEM images shown in Fig. 6 indicate that the increase in the normal force induces a gradual change in the wear mode on AISI 1045 steel from rolling abrasion to a mixed wear mode. As can be seen, there is no grooving for N = 0.132 N (Fig. 6a), although Fig. 5 shows the decrease on the slope even for this load. On the other hand, in the craters shown in Fig. 6b and c, corresponding to higher normal forces, N = 0.220 N and N = 0.315 N, respectively, the grooves are clearly seen. The nonlinear behavior shown in Fig. 5 occurs independently of the existence of grooves, so the decrease of the coefficient cannot be explained simply by the changing on the wear mode. The critical normal force (NC ) is defined as the minimum value of the normal force required for starting the formation of grooves inside the crater. NC depends on the hardness of the material and it is an important parameter for ductile materials, especially for the softer, deserving further studies. The results observed in this study indicates a strong influence of the incorporation of silicon carbide, also affected by the normal force, on the ball cratering tests carried out on ductile materials. The results shown in Fig. 6 revealed that, for AISI 1045 steel, the critical normal force is above 0.132 N. The formation of grooves for normal forces above NC probably enhances the incorporation of abrasive particles at the surface of the sample, which may affect the measurement of the wear coefficient. Abrasive particles attached inside the wear scars modify the behavior of the surface of the material during the test, leading to an artificial decrease in wear coefficient, as shown in Fig. 3. Thus, the values of k given by the linear fittings for L > 1 m in Table 1 are underestimated when compared to the real k of the metallic sample. However, the k values given by the fitting for small sliding distances, L < 1 m, are strongly influenced by specific features related to the initial stages of the test, before the steady state is reached. These initial stages determines, for example, that in many
774
R.V. Camerini et al. / Wear 271 (2011) 770–774
rials, caused by the formation of a region of scuffing that occurs around the edge of the crater [4]. As demonstrated by Trezona and Hutchings [5], the extension of this poorly defined edge region is inversely proportional to the diameter of the crater, which can lead to the overestimation of the craters diameter at low sliding distances. This is the most likely reason for the overestimation of the measured volume of the craters for short sliding distances. Moreover, at the initial stages of the test the wear coefficient may be affected by the surface roughness, wettability of the abrasive suspension on the ball and the sample and by the cold working induced by the metallographic sample preparation procedure, among other factors. 4. Conclusions The behavior of ball cratering test applied to different metallic samples was investigated. For ductile materials, it was possible to observe that the embedment of the abrasive particles at the surface of the crater and the formation of grooves induced to an artificially smaller wear coefficient for sliding distances greater than 1 m. For larger distances, the sphere is sliding over a silicon carbide coated surface and the value of the wear coefficient measured is related to this composition. There is a critical value for the normal force (NC ), that depends on the hardness and mechanical strength of the material, above which the formation of grooves is clearly observed. When the ball cratering test is carried out on ductile materials, the normal force should be smaller than NC to minimize the influence of the attachment of the abrasive particles at the surface of the sample on the measurement of the wear coefficient. Acknowledgements The authors thank to CAPES, CNPQ and FINEP for financial support and to the Electron Microscopy Center of UFRGS. References
Fig. 6. SEM images (backscattering mode) showing the effect of the normal force on the wear mode of AISI 1045 steel. The arrow indicates the direction of ball rolling. All parameters of the test were the same, including the sliding distance (L = 4 m), except the normal force: (a) N = 0.132 N, (b) N = 0.220 N, and (c) N = 0.315 N.
materials – the soft metals are clearly among them – the straight line that represents the best fit does not pass through the origin of the graph. This deviation is due, among other factors, to the low resolution of the edge of the crater, especially in soft mate-
[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] D.N. Allsopp, R.I. Trezona, I.M. Hutchings, The effects of ball surface condition in the micro-scale abrasive, Tribol. Lett. 5 (1998) 259–264. [3] M.G. Gee, A. Gant, I. Hutchings, R. Bethke, K. Schiffman, K. Van Acker, S. Poulat, Y. Gachon, J. von Stebut, Progress towards standardization of ball cratering, Wear 255 (2003) 1–13. [4] A. Ramalho, A geometrical model to predict the wear evolution of coated surfaces, Wear 264 (2008) 775–780. [5] R.I. Trezona, I.M. Hutchings, Three-body abrasive wear testing of soft materials, Wear 233–235 (1999) 209–221. [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 microscale abrasive wear test, Wear 225–229 (1999) 205–214. [7] L. Fang, J. Xing, W. Liu, Q. Xue, G. Wu, X. Zhang, Computer simulation of twobody abrasion process, Wear 251 (2001) 1356–1360. [8] J.O. Bello, R.J.K. Wood, Grooving micro-abrasion of polyamide 11 coated carbon steel tubular for down hole application, Wear 255 (2003) 1157–1167. [9] O. Acselrad, A.R. de Souza, I.S. Kalashniko, S.S. Camargo Jr., A first evaluation of the abrasive wear of an austenitic FeMnAlC steel, Wear 257 (2004) 999–1005. [10] K. Bose, R.J.K. Wood, Optimum tests conditions for attaining uniform rolling abrasion in ball cratering tests on hard coatings, Wear 258 (2005) 322–332. [11] D. Braga, A. Ramalho, P.N. Silva, A. Cavaleiro, Study of abrasion resistance of steels by micro-scale tests, Mater. Sci. Forum 514–516 (2006) 544–548. [12] K. Adachi, I.M. Hutchings, Sensitivity of wear rates in the micro-scale abrasion test to test conditions and material hardness, Wear 258 (2005) 318– 321.
Contents lists available at ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Ball cratering test on ductile materials Rafael Vieira Camerini a,b,c , Rodrigo Buchfink de Souza a,b,c , Felipe de Carli a , Altair Soria Pereira a,d , Naira Maria Balzaretti a,∗ a
Instituto de Fisica, UFRGS, Av. Bento Goncalves, 9500, Porto Alegre, RS, CP 15051, 91501-970, Brazil Jomon Indústria de Artefatos Cerâmicos Ltda, Porto Alegre, RS, Brazil c PGCIMAT, UFRGS, Av. Bento Goncalves, 9500, Porto Alegre, RS, 91501-970, Brazil d Escola de Engenharia, UFRGS, Porto Alegre, RS, Brazil b
a r t i c l e
i n f o
Article history: Received 26 June 2010 Received in revised form 23 February 2011 Accepted 25 March 2011 Available online 2 April 2011 Keywords: Micro-scale abrasion Hardness Steel Non-ferrous metals Wear testing
a b s t r a c t Micro-abrasive wear test, also called ball cratering test, is normally used to investigate the wear behavior of hard bulk materials or wear-resistant coatings. In this work, this test was used to investigate the wear behavior of metallic samples. The results showed that the embedment of abrasive particles and the formation of grooves in the crater’s surface may induce to artificially small wear coefficient when the test is applied to ductile materials. Surface grooving was observed when the normal force applied by the sphere was higher than a critical value. This critical value depends on the hardness and mechanical strength of the material being tested. © 2011 Elsevier B.V. All rights reserved.
1. Introduction The abrasive wear resistance is an important property for a wide range of applications in engineering. Ball cratering test – also called micro-abrasion test – allows the quantitative evaluation of this property even for small volumes of material [1]. The technique consists in the rotation of a sphere of radius R in contact with a polished surface under the action of a normal force (N). During the test, an abrasive suspension is dropped between the ball and the sample. As the geometry of the wear crater produced on the sample surface is related to the spherical shape of the counter-body, the crater volume (V) can be obtained from the ratio between the crater diameter and the sphere diameter. For homogeneous materials, the wear coefficient can be calculated from a simple model that is equivalent to the Archard equation [1]: V = kL N
(1)
In this equation, L is the apparent sliding distance traveled by a fixed point at the surface of the sphere and k is the wear coefficient. A detailed description of this equation can be found in Ref. [1].
The wear mode depends on the abrasive particle motion at the interface between sample and counter-body. A three-body abrasion mechanism occurs when the abrasive particles stay free and roll along the interface between the sphere and the sample surface [1]. In this case, the crater surface shows a homogeneous scar pattern and directional grooves are not observed. On the other hand, a two-body abrasion mode is present when the abrasive particles are attached to the ball or to the specimen surface, producing grooves and scratch marks. In this case, the volume of the removed material in the crater cannot be calculated directly from the starting geometry of the sphere [1]. Therefore, the choice of the testing parameters should minimize the two-body abrasion mechanism [1–7]. The allowed range for these parameters will be strongly dependent of the sample’s mechanical properties. There are several studies in the literature dealing with grooving abrasion associated to the attachment of abrasive particles to the ball [3,6]. However, the consequences of this attachment on the determination of the wear coefficient have not been investigated in detail so far. In this paper, we present results for the ball cratering test on ductile soft materials where the attachment of abrasive particles to the crater surface is expected to occur. The main goal of this study is to evaluate the effect of the abrasive particles embedment on the determination of the wear coefficient. 2. Material and methods
∗ Corresponding author. Av. Bento Goncalves, 9500, Porto Alegre, RS, CP 15051, 91501-970, Brazil. Tel.: +55 51 3308 6489; fax: +55 51 3308 7286. E-mail address: [email protected] (N.M. Balzaretti). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2011.03.013
The wear behavior of aluminum, copper, AISI 1020 steel, AISI 1045 steel and quenched AISI 1045 steel (heat treated at 900 ◦ C
R.V. Camerini et al. / Wear 271 (2011) 770–774
771
Table 1 Vickers hardness and wear coefficient of the samples calculated using Eq. (1). The hardness of the ball used as counter-body in the micro-abrasive wear tests is also included. Sample
Vickers hardness (GPa)
Wear coefficient (m2 /N)
L<1m Aluminum Copper AISI 1020 steel AISI 1045 steel AISI 1045 quenched steel Mullite Stainless steel sphere
L>1m −12
1.72 2.00 2.86 2.94 8.83
10 × 10 5 × 10−12 5 × 10−12 5 × 10−12 2 × 10−12
9 × 10−12 4 × 10−12 3 × 10−12 2 × 10−12 2 × 10−12
8.58 9.02
5 × 10−12
5 × 10−12
during 30 min, followed by cooling in mineral oil) was investigated by the ball cratering test. For comparison, the same experimental procedure was applied to a ceramic specimen of mullite. The Vickers hardness of the sphere and the samples were measured using a micro-hardness device (Shimadzu A-108) with a load of 300 g (2.94 N) (Table 1). The ball cratering wear tests were performed using the free ball/single shaft geometry (CSEM Calowear – Fig. 1), where the ball
Fig. 2. SEM image showing the morphology of the SiC abrasive powder.
is driven by friction with a shaft. The normal force is measured by a load cell positioned below the sample table, and its value is adjusted by the slope of the table, which modifies the component of the sphere’s weight perpendicular to the sample’s surface. The maximum force used in this setting usually does not exceed 0.7 N. A hard steel ball with 25.40 mm in diameter was used to perform the tests. For every sample investigated, the first crater was discarded to ensure the ball would be in the same condition for every consecutive crater. Before the abrasive tests, all samples were polished using an aqueous suspension of alumina powder (0.4 m grain size) and ultrasonic bath cleaned first in distilled water and then in ethanol for 20 min. At the surface of each sample, 11 craters were made, corresponding to sliding distances from 0.4 m to 4.0 m (5–50 rotations of the sphere with a tangential velocity of 0.05 m/s). An abrasive suspension of silicon carbide (SiC) powder in distilled water (0.75 g/cm3 ) was dropped in a rate of 1 drop per sphere’s rotation. The morphology of the abrasive particles was characterized by SEM microscopy (Fig. 2). The particle size distribution was measured by laser diffraction (Cilas 1180LD device) and the results are shown in Table 2. The normal force was kept constant during the tests and all craters in the same sample were produced under the same load. The value of normal force was between 0.265 N and 0.285 N for all tests. The exception occurred for aluminum, the softest studied material (Table 1). In this case, to reduce the scratch inside the craters (groove formation), it was used 0.220 N. Also, additional tests using 0.132 N and 0.315 N were made on AISI 1045 steel to investigate the effect of the normal force on the density of abrasive particles trapped in the craters. Images of the wear craters were obtained using a metallographic microscope with a low-magnification objective. The diameters of the craters were evaluated by software for image analysis (Optimas). A scanning electron microscope with an energy-dispersive spectrometer (SEM/EDS) was used for chemical analysis of the samples.
Table 2 Grain size distribution of the SiC abrasive powder.
Fig. 1. Ball cratering apparatus. (a) Picture of CSEM Calowear; (b) schematic representation of the experimental geometry.
Diameter at 10% Diameter at 50% Diameter at 90% Mean diameter
0.74 m 3.13 m 5.11 m 3.04 m
772
R.V. Camerini et al. / Wear 271 (2011) 770–774
Fig. 3. Volume of the crater divided by the normal force as a function of the sliding distance of a fixed point at the surface of the sphere. (a) Aluminum, (b) copper, (c) AISI 1020 steel, (d) AISI 1045 steel, (e) quenched steel (AISI 1045) and (f) mullite.
3. Results and discussion The results obtained for all the samples are shown in Fig. 3, where the ratio between the volume of the crater and the normal load during the test (V/N) is presented as a function of the sliding distance (L) traveled by the sphere. The dashed lines correspond to linear fittings. For mullite and tempered steel, the hardest investigated materials, the results can be well described by a single straight line. For all other samples, a single linear fitting was not enough to give a good representation of the wear behavior for the whole range of sliding distances. In these cases, two straight lines, one up to L ≈ 1 m, and other for larger sliding distances, can be used to describe quite well the experimental results. Systematically, the slope of the line for larger distances was smaller, corresponding to a smaller wear coefficient (k). Table 1 shows the estimated values of k obtained for both straight lines for each sample. As can be seen, the wear coefficients of the metallic samples are of the same order of magnitude than the ceramic sample and the tempered steel. On the other hand, the hardness values of the ceramic and tempered steel sample are three to four times higher than of the softer samples. A possible explanation for the anomalous low wear coefficient values obtained for the softer materials is related to the embedment of the abrasive particles at the surface of the craters, inhibiting the abrasion of the metal surface. The attachment of abrasive particles in relatively soft samples has been reported by Bello and Wood [8], Acselrad et al. [9] and Bose
and Wood [10]. None of these studies, however, investigated the effects of the attachment on the wear coefficient of these materials. The nonlinear behavior depicted in Fig. 3 for soft metallic materials is normally related to a transition from 3 body to 2 body abrasion wear modes [6,11]. Braga et al. [11] investigated the behavior of steels under micro abrasion test using a ball cratering equipment with fixed-ball. Their study corroborates the idea that this nonlinear behavior should be a consequence of two phenomena: (a) the transition between wear modes and (b) the low resolution of the edge of the crater, required during the evaluation of the volume of the crater. The results shown in the present work pointed out to a third reason which contributes to the nonlinear behavior of the abrasion wear. The attachment of abrasive particles inside the craters may explain why the soft and ductile materials exhibit an apparently anomalous low abrasive wear behavior. The concentration of Si at the surface of the crater can be associated to the embedment of SiC particles. Fig. 4 shows the increase of Si concentration as a function of the sliding distance traveled by the ball for AISI 1020, AISI 1045 quenched steel and copper measured by SEM/EDS analysis. As can be seen, the concentration practically saturates for sliding distances greater than ∼1 m, the sliding distance corresponding to the transition between the two straight lines in Fig. 3 for the soft materials. Fig. 4 also shows that the amount of embedded particles is different for each sample, ranging from ∼6 at% for quenched AISI 1045 steel to ∼20 at% for Copper and AISI 1020 steel.
R.V. Camerini et al. / Wear 271 (2011) 770–774
773
Fig. 5. Abrasive wear of AISI 1045 for different values of normal force: (䊉) 0.132 N, () 0.220 N, (×) 0.315 N.
Fig. 4. (a) Atomic concentration of Si in the craters as a function of the sliding distance traveled by the sphere. () AISI 1020 steel; () AISI 1045 quenched steel, () copper. (b) SEM image of a crater in AISI 1045 quenched steel. (c) EDS analysis of region (1), inside the crater. (d) EDS analysis of region (2), outside the crater.
The density of SiC particles attached on the scars of the softer steel (AISI 1020) is higher than in the harder steel. This is probably a consequence of the degree of anchoring of the abrasive particles in this material, without the complete burial of the grains. On the other hand, for the harder steel (AISI 1045), its resistance to plastic deformation leads to the formation of shallow indentations that are
not always able to promote the adhesion of the abrasive particle. For the softer metallic samples, on the contrary, the low resistance to plastic deformation allows the complete burial of abrasive particles in these materials. Thereby, despite the graph in Fig. 4a shows a high concentration of silicon carbide inside the craters on copper specimen, the presence of the grooves at the surface with an abrasive particle stuck to the end were not easily found probably because the particles were completely buried at the surface. Fig. 4b shows an image of the crater on AISI 1045 quenched steel. Fig. 4c and d shows EDS analyses of regions inside (1) and outside (2) the crater, indicating there was no Si outside the crater region. Fig. 5 compares the wear behavior of AISI 1045 steel for three different values of normal force: 0.132 N, 0.220 N and 0.315 N. For short sliding distances (when L < 1 m) the volume of craters normalized by the normal force was almost the same. However, for longer sliding distances, the slope and, consequently, the wear coefficient, decreased as the normal force increased. This phenomenon was already reported by Adachi and Hutching [12] for tool steel samples. The SEM images shown in Fig. 6 indicate that the increase in the normal force induces a gradual change in the wear mode on AISI 1045 steel from rolling abrasion to a mixed wear mode. As can be seen, there is no grooving for N = 0.132 N (Fig. 6a), although Fig. 5 shows the decrease on the slope even for this load. On the other hand, in the craters shown in Fig. 6b and c, corresponding to higher normal forces, N = 0.220 N and N = 0.315 N, respectively, the grooves are clearly seen. The nonlinear behavior shown in Fig. 5 occurs independently of the existence of grooves, so the decrease of the coefficient cannot be explained simply by the changing on the wear mode. The critical normal force (NC ) is defined as the minimum value of the normal force required for starting the formation of grooves inside the crater. NC depends on the hardness of the material and it is an important parameter for ductile materials, especially for the softer, deserving further studies. The results observed in this study indicates a strong influence of the incorporation of silicon carbide, also affected by the normal force, on the ball cratering tests carried out on ductile materials. The results shown in Fig. 6 revealed that, for AISI 1045 steel, the critical normal force is above 0.132 N. The formation of grooves for normal forces above NC probably enhances the incorporation of abrasive particles at the surface of the sample, which may affect the measurement of the wear coefficient. Abrasive particles attached inside the wear scars modify the behavior of the surface of the material during the test, leading to an artificial decrease in wear coefficient, as shown in Fig. 3. Thus, the values of k given by the linear fittings for L > 1 m in Table 1 are underestimated when compared to the real k of the metallic sample. However, the k values given by the fitting for small sliding distances, L < 1 m, are strongly influenced by specific features related to the initial stages of the test, before the steady state is reached. These initial stages determines, for example, that in many
774
R.V. Camerini et al. / Wear 271 (2011) 770–774
rials, caused by the formation of a region of scuffing that occurs around the edge of the crater [4]. As demonstrated by Trezona and Hutchings [5], the extension of this poorly defined edge region is inversely proportional to the diameter of the crater, which can lead to the overestimation of the craters diameter at low sliding distances. This is the most likely reason for the overestimation of the measured volume of the craters for short sliding distances. Moreover, at the initial stages of the test the wear coefficient may be affected by the surface roughness, wettability of the abrasive suspension on the ball and the sample and by the cold working induced by the metallographic sample preparation procedure, among other factors. 4. Conclusions The behavior of ball cratering test applied to different metallic samples was investigated. For ductile materials, it was possible to observe that the embedment of the abrasive particles at the surface of the crater and the formation of grooves induced to an artificially smaller wear coefficient for sliding distances greater than 1 m. For larger distances, the sphere is sliding over a silicon carbide coated surface and the value of the wear coefficient measured is related to this composition. There is a critical value for the normal force (NC ), that depends on the hardness and mechanical strength of the material, above which the formation of grooves is clearly observed. When the ball cratering test is carried out on ductile materials, the normal force should be smaller than NC to minimize the influence of the attachment of the abrasive particles at the surface of the sample on the measurement of the wear coefficient. Acknowledgements The authors thank to CAPES, CNPQ and FINEP for financial support and to the Electron Microscopy Center of UFRGS. References
Fig. 6. SEM images (backscattering mode) showing the effect of the normal force on the wear mode of AISI 1045 steel. The arrow indicates the direction of ball rolling. All parameters of the test were the same, including the sliding distance (L = 4 m), except the normal force: (a) N = 0.132 N, (b) N = 0.220 N, and (c) N = 0.315 N.
materials – the soft metals are clearly among them – the straight line that represents the best fit does not pass through the origin of the graph. This deviation is due, among other factors, to the low resolution of the edge of the crater, especially in soft mate-
[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] D.N. Allsopp, R.I. Trezona, I.M. Hutchings, The effects of ball surface condition in the micro-scale abrasive, Tribol. Lett. 5 (1998) 259–264. [3] M.G. Gee, A. Gant, I. Hutchings, R. Bethke, K. Schiffman, K. Van Acker, S. Poulat, Y. Gachon, J. von Stebut, Progress towards standardization of ball cratering, Wear 255 (2003) 1–13. [4] A. Ramalho, A geometrical model to predict the wear evolution of coated surfaces, Wear 264 (2008) 775–780. [5] R.I. Trezona, I.M. Hutchings, Three-body abrasive wear testing of soft materials, Wear 233–235 (1999) 209–221. [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 microscale abrasive wear test, Wear 225–229 (1999) 205–214. [7] L. Fang, J. Xing, W. Liu, Q. Xue, G. Wu, X. Zhang, Computer simulation of twobody abrasion process, Wear 251 (2001) 1356–1360. [8] J.O. Bello, R.J.K. Wood, Grooving micro-abrasion of polyamide 11 coated carbon steel tubular for down hole application, Wear 255 (2003) 1157–1167. [9] O. Acselrad, A.R. de Souza, I.S. Kalashniko, S.S. Camargo Jr., A first evaluation of the abrasive wear of an austenitic FeMnAlC steel, Wear 257 (2004) 999–1005. [10] K. Bose, R.J.K. Wood, Optimum tests conditions for attaining uniform rolling abrasion in ball cratering tests on hard coatings, Wear 258 (2005) 322–332. [11] D. Braga, A. Ramalho, P.N. Silva, A. Cavaleiro, Study of abrasion resistance of steels by micro-scale tests, Mater. Sci. Forum 514–516 (2006) 544–548. [12] K. Adachi, I.M. Hutchings, Sensitivity of wear rates in the micro-scale abrasion test to test conditions and material hardness, Wear 258 (2005) 318– 321.