**1. Introduction**

Machine parts are subject to the following wear types: Abrasive, adhesive, fatigue and corrosive. In abrasive wear, chipping of harder material a micrometer scale occurs as result of rubbing the soft member. The wear is formed as result of cutting, hitting, and scratching. Abrasion takes places at the solid-solid, particle-solid, solid-liquid interface [1].

If one of the surfaces which are in touch is rough and hard, it chips the other surface due to relative motion or touching forces. The wear is called two-body abrasive wear. If there are free abrasive particles between the two bodies, the wear is called three-body abrasive wear. The free abrasive particles may be external material dust or the remains of chipping. Usually, the wear starts as a two-body abrasive or adhesive wear and then becomes a threebody wear as dust form between the two surfaces due to external particles, chipping remains, or oxide particles. In three-body abrasive wear, wear rate increases as diameter of abrasive particles increases. Gouging, high stress abrasion and low stress abrasion are types of three-body abrasive wear [1-3].

In gouging, surface wear is formed using large abrasive particles. A gouging mechanism is common in ground leveling machines and excavation and digging machines. In these machines, wear occurs on moving, digging, and excavating members. High-stress abrasion occurs when the sharp edged small abrasive particles, which are formed by the crushed particles under excessive loads, scratch the surface. Ball-bearing grinders are mostly subject to high-stress abrasion [2]. These grinders are predominately used to crush the metallic ores and minerals. High-stress abrasion contributes to the significant portion of the wear in grinders. Low-stress abrasion takes place when there is no crushing or grinding in the abrasive particles and one of the surfaces is subject to wear [3].

Abrasive wear experiments have been made with substances containing one or more abrasive. Abrasive statements, which are obtained through single abrasive end patterns (i.e.

© 2013 Sevim, licensee InTech. This is an open access chapter distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited. © 2013 Sevim, licensee InTech. This is a paper distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

sphere, pyramid, and cone) are adapted to abrasive wear cases with abrasive particle more than one based on some assumptions. The abrasive particle is generally modeled as a cone shape [4]. Rabinowicz [5] has aimed at a simple expression for the volume of material removed during two-body abrasion by a conical abrasive particle;

$$\frac{V}{L} = \left(\frac{2\tan\alpha}{\pi}\right)\left(\frac{F\_N}{H}\right) \tag{1}$$

Effect of Abrasive Particle Size on Abrasive Wear Resistance in Otomotive Steels 31

(3)

*m* abrasive

Generally, (2) does not agree with the experimental results. The main reasons for this incompatibility are the changes of wear coefficient *k* depending on abrasive grit size [5, 6]. In literature, there are many investigations about the effect of the abrasive grit size on abrasive wear rate in zone *I*. Avient et al [7] have examined the abrasive behavior of many materials and realized that the clogging of the interstices between the finer abrasive grains by wear debris is responsible for the grit size effect. This decreases the number of abrasive grains, which contact the surface and remove material, thus decreasing the abrasive wear rate. Mulhearn and Samuel [8] studied samples of silicon carbide (SiC) abrasive papers. They believe that the mechanical properties of coarse and of fine abrasive grains are different, and that the fine grains have a needle-like shape and contain many cracks, thus braking up more readily. In this way, abrasive wear rate becomes zero, because small grains are flattened. Rabinowich and Mutis [9], have aimed an account of the size effect using adhesive wear particles. Using a surface energy criterion, they theoretically show that the critical abrasive particle size is a function of the adhesive particle size of the material being worn away. Sin et.al [5] have used the critical depth of penetration to explain the effect of grit size on abrasive wear loss and have found that there was not a critical abrasive particle size for a

specific material. They also showed that the constant wear rate starts at 80

particle size for all metals used in the experiments. The elastic contact hypothesis was first suggested by Larsen-Badse [10] who measured the size and number of grooves formed on polished copper specimens abraded by SiC abrasive papers and estimated the real contact area. He postulated that many fine grit have elastic interaction with the surface. It was also suggested that the fraction of the load carried by particle in elastic contact increased with decreasing grit size since it is unlikely that the abrasive grits gradually become more angular with increased size. Moore and Douthwaite [11], have tried to explain the size effect by plastic deformation concept below worn surfaces. They estimated the equivalent plastic strain and the flow stress as a function of depth below worn surface and calculated the work done in deforming the material below the groove and energy absorbed in plowing the surface. They concluded that the energy expended in plastic deformation of material to form the grooves and deform the surface account for almost all the external work done for all grit sizes in abrasion and that wear volume is dependent on the grit size probably because the deterioration and pick up of abrasive particles become more intense at small grit sizes. Hutchings [12], has stated that the size effect is due to the variation of shape changing rate dependent to abrasive particle size. However, Misra and Finnie [13], have found that the shape-changing rate has only changed the wear resistance, and has no effect on the dependency of abrasive particle size. Many researchers have examined the abrasive particle size effect in the zone *II* [14, 15]. Rabinowich [14] determined empirically the following abrasive wear rate expression for the *zone II* using only one type of abrasive particle size

> 1 2 3 3 *<sup>O</sup>*

where *H* is the hardness of alloy, *HO* is the hardness of the alloy in the fully soft condition

*H H*

*<sup>P</sup> W k*

and *P* is the pressure applied to the surface.

where *V* is the volume loss due to wear, *L* the sliding distance, *FN* the normal load on the conical particle and *H* hardness of wearing surface and α the attack angle of the abrasive particle.

The equation (1), for linear wear density can be written as follows [1];

$$\mathcal{W} = k \frac{P}{H} \tag{2}$$

Where; *W*: linear wear density, *k*: wear coefficient, *P*: pressure applied on surface, *H*: hardness of abraded material.

For pure metals and annealed steels, the wear resistance versus hardness is a line passing through the origin. The linear zone is called zone *I* throughout the paper. The abrasive wear resistance versus hardness graph of the heat-treated steels is a line not passing through the origin [3]. This behavior cannot be derived from (2). The zone corresponding this is called zone *II* . The zones *II* and *I* are shown in Figure 1 [1, 4]. (2) is similar to the Archard's adhesive wear expression.

**Figure 1.** Relationship between wear resistance and hardness [1, 4].

Generally, (2) does not agree with the experimental results. The main reasons for this incompatibility are the changes of wear coefficient *k* depending on abrasive grit size [5, 6]. In literature, there are many investigations about the effect of the abrasive grit size on abrasive wear rate in zone *I*. Avient et al [7] have examined the abrasive behavior of many materials and realized that the clogging of the interstices between the finer abrasive grains by wear debris is responsible for the grit size effect. This decreases the number of abrasive grains, which contact the surface and remove material, thus decreasing the abrasive wear rate. Mulhearn and Samuel [8] studied samples of silicon carbide (SiC) abrasive papers. They believe that the mechanical properties of coarse and of fine abrasive grains are different, and that the fine grains have a needle-like shape and contain many cracks, thus braking up more readily. In this way, abrasive wear rate becomes zero, because small grains are flattened. Rabinowich and Mutis [9], have aimed an account of the size effect using adhesive wear particles. Using a surface energy criterion, they theoretically show that the critical abrasive particle size is a function of the adhesive particle size of the material being worn away. Sin et.al [5] have used the critical depth of penetration to explain the effect of grit size on abrasive wear loss and have found that there was not a critical abrasive particle size for a specific material. They also showed that the constant wear rate starts at 80 *m* abrasive particle size for all metals used in the experiments. The elastic contact hypothesis was first suggested by Larsen-Badse [10] who measured the size and number of grooves formed on polished copper specimens abraded by SiC abrasive papers and estimated the real contact area. He postulated that many fine grit have elastic interaction with the surface. It was also suggested that the fraction of the load carried by particle in elastic contact increased with decreasing grit size since it is unlikely that the abrasive grits gradually become more angular with increased size. Moore and Douthwaite [11], have tried to explain the size effect by plastic deformation concept below worn surfaces. They estimated the equivalent plastic strain and the flow stress as a function of depth below worn surface and calculated the work done in deforming the material below the groove and energy absorbed in plowing the surface. They concluded that the energy expended in plastic deformation of material to form the grooves and deform the surface account for almost all the external work done for all grit sizes in abrasion and that wear volume is dependent on the grit size probably because the deterioration and pick up of abrasive particles become more intense at small grit sizes. Hutchings [12], has stated that the size effect is due to the variation of shape changing rate dependent to abrasive particle size. However, Misra and Finnie [13], have found that the shape-changing rate has only changed the wear resistance, and has no effect on the dependency of abrasive particle size. Many researchers have examined the abrasive particle size effect in the zone *II* [14, 15]. Rabinowich [14] determined empirically the following abrasive wear rate expression for the *zone II* using only one type of abrasive particle size

30 Tribology in Engineering

particle.

hardness of abraded material.

adhesive wear expression.

Wear Resistance

*W-1, (GPa)*

sphere, pyramid, and cone) are adapted to abrasive wear cases with abrasive particle more than one based on some assumptions. The abrasive particle is generally modeled as a cone shape [4]. Rabinowicz [5] has aimed at a simple expression for the volume of material

> 2tan *<sup>N</sup> V F L H*

where *V* is the volume loss due to wear, *L* the sliding distance, *FN* the normal load on the conical particle and *H* hardness of wearing surface and α the attack angle of the abrasive

*<sup>P</sup> W k*

Where; *W*: linear wear density, *k*: wear coefficient, *P*: pressure applied on surface, *H*:

For pure metals and annealed steels, the wear resistance versus hardness is a line passing through the origin. The linear zone is called zone *I* throughout the paper. The abrasive wear resistance versus hardness graph of the heat-treated steels is a line not passing through the origin [3]. This behavior cannot be derived from (2). The zone corresponding this is called zone *II* . The zones *II* and *I* are shown in Figure 1 [1, 4]. (2) is similar to the Archard's

Vickers Hardness *HV10, (GPa)*

Zone *I*

(1)

*<sup>H</sup>* (2)

Heat Treated Steels

Zone *II*

Pure Metals and Annealed Steels

removed during two-body abrasion by a conical abrasive particle;

The equation (1), for linear wear density can be written as follows [1];

**Figure 1.** Relationship between wear resistance and hardness [1, 4].

$$\mathcal{W} = k \frac{P}{\frac{1}{3}H + \frac{2}{3}H\_O} \tag{3}$$

where *H* is the hardness of alloy, *HO* is the hardness of the alloy in the fully soft condition and *P* is the pressure applied to the surface.

Khruschov [15] has studied experimentally the zone *I* in a stationary abrasive particle size using the non-heat treated steels and he found the relative wear resistance –hardness relationship for metals as follows;

$$
\mathfrak{E} = \mathfrak{b}H \tag{4}
$$

Effect of Abrasive Particle Size on Abrasive Wear Resistance in Otomotive Steels 33

Ni

Al

Cu

Ti

V

(%)

(%)

Vickers hardness *HV*10 *(MPa)*

(%)

(%)

(%)

**Alloys** <sup>C</sup>

(%)

Si

Mn

P

S

Cr

Mo

(%)

(%)

**1010** 0.107 0.11 0.413 0.019 0.025 \_ 0.003 \_ 0.032 0.031 0.002 \_

**1030** 0.328 0.069 0.673 0.015 0.019 \_ 0.001 \_ \_ 0.037 0.002 0.005

**1040** 0.402 0.247 0.82 0.012 0.028 0.025 0.001 0.003 0.014 0.032 0.001 0.003

**1050** 0.506 0.252 0.654 0.014 0.006 0.251 0.002 \_ 0.006 0.017 0.002 0.006

**50CrV4** 0.523 0.394 0.915 0.021 0.027 0.917 0.025 0.034 \_ 0.183 \_ 0.095

AISI1010 - 1648±10 AISI1030 - 1716±20 AISI1040 - 1961±29 AISI1050 - 2175±34 50CrV4 - 2549±49 AISI1010 Water quenched from 900-925 oC 2255±54 AISI1030 Water quenched from 830-850 oC 5609±20 AISI1040 Water quenched from 820-850 oC 6276±15 AISI1050 Water quenched from 810-840 oC 6570±0 50CrV4 Water quenched from 830-850 oC 8895±0 AISI1010 Water quenched from 900-925 oC + 2 hours refrigerated at –25 oC 2256±10 AISI1030 Water quenched from 830-850 oC + 2 hours refrigerated at –25 oC 6767±25 AISI1040 Water quenched from 820-850 oC + 2 hours refrigerated at –25 oC 7100±39 AISI1050 Water quenched from 810-840 oC + 2 hours refrigerated at –25 oC 7875±20 50CrV4 Water quenched from 830-850 oC + 2 hours refrigerated at –25 oC 8895±0 AISI1010 Water quenched from 900-925 oC + tempered at 250 oC 1873±25 AISI1030 Water quenched from 830-850 oC + tempered at 250 oC 5551±34 AISI1040 Water quenched from 820-850 oC + tempered at 250 oC 5943±17

(%)

(%)

**Table 1.** The chemical compositions of experiment sample (wt. %) [19, 20]

Materials Heat Treatment

(%)

(%)

where is the relative wear resistance, *b* a constant coefficient and *H* the initial hardness.

Furthermore the following relationship has been determined to be in zone *II*, between the relative wear resistances of heat-treated steels and hardness;

$$
\varepsilon = \left(\varepsilon\_0 - \mathbb{C}\_0 H\_0\right) + \mathbb{C}\_1 H \tag{5}
$$

where 0 and *H*0 are the relative wear resistance and hardness of annealed steel, and *C*0 and *C*1 are constants.

There are numerous explanations in the literature to explain the abrasive grit size effect. However, most of them have been insufficient since they have not been able to explain the grit size effect encountered in all abrasive wear mechanisms (for example erosive wear) [15-18].

The focus of this study is to investigate the effect of abrasive particle size on abrasive wear resistance in zone *I*, *II* and to develop the equations of empirical abrasive wear resistance connected to abrasive particle size. Moreover, to search for the effects of relative wear resistance in zone *I*, *II* and to develop the equations of empirical relative wear resistance connected to abrasive particle size.
