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Subsurface characteristics of an abraded low carbon steel
WEAR ELSEVIER
Wear 210 ¢1997) 204-210
Subsurface characteristics of an abraded low carbon steel
Hilseyin ~imeno~lu lstanbul Technical Universi~*. Departmem of Metallurgical En&ineering. 80626 Maslak. istanbul, Turkey
Received 25 November 1996; accepted 4 March 1997
Abstraet Ductile metals commonly exhibit plastic deformation at and near the worn surface and their flow behaviour at large strains has a cleareffect on wear resistance. In this study, the characteristics of the near-surface region ofa ferritic-peaditic steel (0.2% C, 1.2% Mn), subjected to abrasive wear tests, were examined. Wear tests were performed under different loads by robbing the specimens on sliding 60 mesh A!203 abrasive band. The metallographic technique used to determine the magnitude of plastic deformation was based on measurement of the displacements of pearlite bands. The hardne~ of the plastic deformation zone was determined by performing ultramicrohardness tests along fertile bands with a Vickers indenter. Microscopic examinations of the near-surface regions revealed the wear mechanism to be ploughing and the deformation mechanism to be cross-slip. It was observed that plastic strain (more than 6) occurred on the abraded surface, and increased the hardness to about !.5 times the original value. The strain and hardness gradient extended to a larger depth into the bulk with increasing wear test load. It is concluded that the wear resistance of the investigated steel increases by work hardening of the near-surface region which is required to consume high energy for abrasion, during sliding. Ultramicrohardness measurements performed on worn specimens revealed high har4ness, as the indent size decreased. The indentation size-hardness relation was explained by adislocation model incorporating geometrically necessary dislocations due to the presence of strain gradients in the deformation region around the indent. © 1997 Elsevier Science S.A. Ke~n~,ords: Abrasivewear;.Indentation size effect: Low carbon steel: Uhram]crohardness'Work hardening
1, Introduction Wear is a surface phenomenon that causes alteration of the dimensions of materials by microchip formation due to the mechanical action of contacting materials. When a material surface is rubbed against another material under stress, the contact surface of the metal is worn and grooves are formed on it, in the direction parallel to sliding. In abrasive wear, material is removed or displaced from the surface by hard particles as they embed in the surface and move relative to it [ i-31. Sliding contact is often accompanied by plastic deformation and/or microfracture which influence the wear behaviour and may be associated with changes in the microstructure [ 4 - 6 ] . In ductile materials, the strains at the worn surfaces are much larger than conventional deformations, and the wear resistance correlates well with the surface hardness rather than the bulk hardness [2,7,8]. Previously the effect of bulk hardness on the abrasive wear performance of carbon steels has been studied [9,10]. Recently it has been reported ,hat. the bulk hardness can be used as a predictor only for annealed steels with the same type of microstructure [ 11 ]. 0043-1648/97/$17.00 © 1997ElsevierScience S.A. All rights reserved PIi S0043-1648 ( 97 )00051-3
Xu and Kennon [ 11] developed an abrasion grooving model for annealed steels, based on the groove size and thickness of the plastic deformation zone under the worn surface, by assuming the decrease in these parameters with increasing carbon content. However, they did not make any attempt to measure the thickness of the plastic deformation zone. The variatiop of the thickness of the plastic deformation zone with .,~iding distance, load and abrasive grid size has been studied in some non-ferrous metals [ 12-15]. The present study aims to perform metallographic examinations and hardness surveys on the longitudinal cross-sections of an annealed low carbon steel worn by sliding on abrasive particles, to determine the strain and hardness distribution under ,*he ,';am surface with respect to load. Special attention is given to the load-indentation depth-time data obtained from ultramicrohardness tests. 2, Experimental procedure The low carbon steel (0.21% C, 0.23% Si, i.18% Mn, 0.05% P, 0.05% S) utilized in this study was a hot rolled bar ( I0 mm diameter). Wear specimens with 3 X 3 mm 2 square tips w e ~ machined from the steel rod and annealed at 950
H. ~imeJ:J.,: '.. / Wear 210 (1997) 204-210
Lo~t~
(::) Fig. 1. Schematic view of the abrasive wear tester utilized in this inveztigation. °C for 1 h. The surfaces of the annealed specimens were etched with an HI= solution after grinding to reveal the micmstructure. Wear tests were carded out using the wear tester shown in Fig. 1, under four different loads ( 16.4 N to 65.6 N). At each load level two specimens were tested. During wear tests, specimens that were rubbing on 60 mesh AI203 abrasive band were moving perpendicular to the sliding direction so that they always passed over fresh abrasive. The sliding distance and the speed of the abrasive band were 4.6 m and 0.12 m s - t, respectively. The wear volume of the specimens was calculated from weight loss measurements, taken to the nearest 0.1 mg. Longitudinal sections of the worn specimens (on a plane normal to the wear surface and parallel to the sliding direction) were examined by optical and scanning electron microscopes. Hardness surveys of the sectioned specimens were made with a Fisher H 100 XYPROG ultramicrohardness tester [ 16]. During the hardness tests, load-indentation depth-time data were recorded as a Vickers indenter was driven into the sample (loading segment) and then withdrawn from it (unloading segment). A maximum test load of I00 mN was applied in 50 steps and the hold time at each step was 0.1 s. The loading/unloading rate of the indenter was 3.6 mN s - '.
3. Results and discussion
20
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Load (N) Fig. 2. The effect of abrasiveweartest loadon rig wearrateof the investigatedsteel.
3.2. Subsurface microstructure Metaliographic examinations of longitudinal sections revealed shear deformation under the worn surface. As shown in Fig. 3, the microstmcture of the investigated steel consisted of ferrite and pearlite bands elongated in the rolling direction. Below the worn surface, pearlite bands, which were initially perpendicular to the sliding abrasive, were displaced towards the slip direction. Ferfite grains were also bent and reoriented in the slip direction. Fine fenite grains were observed just beneath the worn surface (Fig. 4). It is reviewed [2,4,6] that the formation of the fine grained substructure under the worn surface is due to the heavy shear deformation introduced by wear, and such a sliding contact creates a near-surface layer having different structure and properties than the base matefiat. Fig. 5, which we,,,; obtained from a specimen etched before wear testing, shows the effects of abrasive wear on the subsurface. The rough surface and slip bands in ferrite grains indicate heavy plastic deformation concentrated in ferrite grains. A high magnification scanning electron microscopy image of the near-surface region (Fig. 6) reveals wavy slip
3.1. Wearrate Fig. 2 shows the effect of applied load on the wear rate, which is described as the volume loss per unit sliding distance ( r a m 3 m - t ). The wear rate increased from 0.6 mm 3 m - ~to 1.8 mm 3 m - t linearly when the load increased from 16.4 N tO 65.6 N. The slope of Fig. 2 (25 × 10 -3 mm 3 N - i m-~) is expressed as the dimensional wear coefficient k in the Archa~l equation [ 1-3]. The monotonic increase in wear rate with respect to load without showing any transient behavlout, indicates that the same wear mechanism is effective through the entire load range. Therefore, in the following sections, only the results of metallographic and hardness examinations of worn specimens under the test loads of 16.4 N and 65.6 N are given and discussed.
Fig. 3. The banded microm'actam o f t l ~ iavesdgmed low carbo~ steel subjcctcd to abrasive wear testing under a load of 16.4 N.
206
1t. ~'imeno~lu / Wear 210 (1997) 204-210
Fig, 7 shows a groove and the flow pattern beneath it, caused by penetration of abrasive to the worn surface. Thus, penetration of the abrasive particle displaced material from the groove to the ridges at the sides of the groove (by ploughing) and this displaced material was then removed as microchips. In a previous investigation [ 11], the abrasive wear mechanism of annealed low carbon steels was reported as microcutting and microploughing, which is in agreement with this observation.
3.3. Subsurface plastic strain
1,L Fig. 4, Optical micrograph oflhe subsurface showing fine fertile grainsjust beneath tbe worn surface (test load .65.6 N).
Rigney el al. [ 7 ] noticed that some microstructural features such as grain boundaries, twin boundaries and lamellae can be utilized to provide information about the deformation characteristics of layers below the worn surfaces and that optical microscopy was sufficient for this purpose. In the present study, the strain dis~bution under the worn surface was obtained by measuring the displacements of several pearlite bands in the sliding direction. Measurements were taken on printouts of the optical images that were transferred to a PC, equipped with a videographic card, and magnified. The displacements of pearlite bands at each depth were utilized to determine the shear deformation. The effective plastic strain at any given depth was calculated from the shear angle of pearlite bands ~ [ 12-14] ; e = (~/3/3) tan ~
Fig. 5. Opticalmicrographof the longitudinalsectionof the.wornspecimen (testload 65.6 N) etcbed before wear testing.
( I)
The variation of average effective strain with distance from the worn surface is plotted in Fig. 8. It should be mentioned that the standard deviation of the mean values was about I. 1 at depths le,ss than 20 Ixm, and 0.2 at greater depths, Fig. 8 shows that at higher loads, a strain gradient exists up to larger depths. AccordirJg to the results of abrasive wear tests conducted on copper-silver solder material, the strain as a function of depth is proportional to the square root of load [ 12]. Al-SiC composites also exhibited an increase in the thickness of the plastic deformation zone with load [ 13,15]. The strain profile of Fig. 8 can be described by an exponential [ 12,141;
Fig, 6. Scanning electron microgmph of the near-surface layer of the worn specimen (test load 65.6 N). Ferrit¢ and pearlite grains are ::hown by the symbols F and P respectively.
lines in ferrite grains which are evidence of cross-slip. In the literature [ 17,18], it is reported that, at large plastic s~.rains, the deformation mechanism of metals is controlled by crossslip which is associated with cell formation in the parabolic (stage Ill) hardening region. Transmission electron microscopy examinations revealed dislocation cell and subgrain formation under the worn surfaces of copper based metals, subjected to sliding wear and erosion tests [ 7,191.
Fig. 7. Optical micrograph of the longitudinal section showing a groove and flow lines beneath it (test load 64.6 N),
H. (?imeno~,ht/ Wear 210 ( ! 997) 204-216 m
.,,,..
$1Rnncd ~ •
[
IQ4 N O~IBN
5
°i
1
! 20
40
60
80
100
1~o
Distance from worn surface (pm) Fig. 8. The variation of effective strain with distance from the worn surface.
(2)
E=A exp(--BZ)
where Z is the surface depth, A is the extrapolated strain at the surface and B is a constant related to the strain rate at the surface [ 7 ]. The numerical values of A and B were calculated from Fig. 9 as A = 9, I i~m Ixm- i and B - 0,09 i,tm- ' for 16.4 N and A = 6.7 p.m i.tm-n and B---0.05 p.m- ~ for 65.6 N. Alpas and Zhang [20] also noted that the values of coefficients A and B depend on load. According to Fig. 9, abrasion under the loads of 16.4 and 65.6 N caused plastic strains of about 7 and 9 at the worn st'Maces of the investigated steel respectively. These strain levels reached at the worn surfaces are extremely high compared with those reached under conventional deformation processes.
3.4. Subsurface hardness In addition to the measurements of strain gradients below the abraded surface, the hardness distribution in this region
207
was determined. Since heavy deformation was ob~rved in ferrite grains (Figs. 5 and 6), ultramicrohardness measurements were carried out along ferrite bands with 25 Ixm interval. On each specimen, hardness tests were performed at four different locations and average values are plotted in Fig. 10. The standard deviations of the mean values were about 10% of the mean. Fig. l0 shows that the layers that were strained owing to abrasion (Fig. 9) were hardened, and the hardness gradient extends to a I~rger depth into the bulk with increasing load. The increase in hardness of fertte (about 50% with respect to the original hardness) at the worn surface can be attributed to work hardening. Some non-ferrous f.c.c, metals subjected to abrasive wear exhibited a hardness !ncrement two or three times that of the bulk at their worn surfaces [2,12,14]. It is well known that work hardening occurs as a result of dislocation interaction and, therefore, the degree of hardening is a function of dislocation density [ 17,18 ]. According to Moore et al. [ 21 ], hardening of the worn metal surface due to dislocation interactions is small compared with interstitial-dislocation interactions in b.c.c, solid solutions even at high dislocation densities, and b.c.c, metals show lower degrees of work hardening than f.c.c, metals at their worn surfaces.
3.5. hldentation si=e effect Ultramicrohardness measurements performed on the longitudinal sections of the worn specimens exhibited indentation depth dependence of hardness. Fig. I I compares the variation of hardness with indentation depth for work hardened and unhardened layers of the worn specimen tested at 65.6 N. In both layers the hardness increases as the indent .~:.ze decreases. This phenomenon is called the indentation size effect. It should be mentioned that the hardness values plotted in Fig. 10 corresponded to the maximum indentation depths under 100 raN. As shown in Fig. i i, maximum depths of
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60
80
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100
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Distance from worn s u r f a c e ( g m ) Fig. 9. Plot of/_~ effective strain against distance from the worn surface, to determine the values of coefficients A and B of Eq. (21.
"=7~ 0
=
l
1
!
I
20
40
60
80
100
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Distance from wom surface (gin) Fig. 10. The variation of hardness with distance from Ihe worn suff~ce.
208
14. ~'imeno~lu/ Wear 210 (1997) 204-210 m
- -
.,
°t 0
"
I
I
I
I
I 15
Indentaeo. depth (pro) Fig. 11. Hardness as a function of indentation depth for work hardened (dislince from worn surface 10 tun) and m f l l d e n e d (distance from worn surface I i0 pan) layers of the specimen wont under 65.6 N.
indent of the work hardened and unhardened layers are 1.07 g m and 1.39 Ixm, respectively. This indicates that, under similar wear test conditions (load, grid size, etc.), work hardening increases the resistance of the contact surface to penetration of abrasive particles and decreases the groove size, compared with bulk material. Therefore, better correlation is often found in ductile metals between the wear resistance and the hardness of the worn surface rather than the hardness of the bulk material [2,81. Recently, Poole et al. [22l developed a dislocation model to analyse the indentation size effect on materials that can deform plastically by glide of dislocations. They tested their model experimentally on fully annealed and fully work hardened copper. This dislocation model is based on the creation of geometrically necessary dislocations around the indemand assumes the dislocation density is the sum of the background dislocation density O, and geometrically necessary dislocation density Or Since hardness is proportional to flow stress which is, in turn, pmpoRional to the square root of dislocation density, the indentation size dependence of hardness was expressed by the equation [ 22]; /42 ffi (3ap.bM) 2(p, + ta,t2[i/2bD)
because of the occurrence of a high strain gradient in the deformation region around the indent, and they observed experimentally a steeper slope for worked copper than for annealed copper. Fig. 12 shows that an approximately linear relationship exists for both work hardened and unhardened layers of the worn specimen. The slopes of the work hardened and unhardened layers are i.46 GPa 2 t i m - ' and 1.03 GPa 2 I~m-' respectively. From Eq. (3) one can calculate the slope d(HZ)/(D - ' ) theoretically as !.03 GPa z t t m - ' (by taking /~ = 75 GPa and b ffi 2.5 × 1 0 - ' ttm for a low carbon ferritic steel) which corresponds m the slope of the unhardened layer. The higher slope of the work hardened layer than the unhardened layer implies a localized plastic deformation region around the indent in the work hardened layer. However, in the unhardened layer, which has a high work hardening rate, a rather large plastic deformation region is generated under the indenter [ 17]. This implies that work hardening of the abraded surface limits the spread of the material under the groove over a wide volume, by plastic flow.
3.6. Subsurface plastic energy The ultramicrohardness tester utilized in this study was also capable of generating plots of indentation depth vs. load. These data are given in Fig. 13. for work hardened and unhardened layers discussed in Section 3.5. The area between loading and unloading segnmnts of the load-indentation depth plot gives the plastic energy expended during indentation, which was obtained for the work hardened and unhardened layers as 47.66 nJ and 34.92 nJ respectively. S i ~ e these layers have different plasticity, plastic energies expended during indentation create different indent volumes ( V = 8.17LP) on the surfaces being tested. Therefore, the energy absorbed by plastically deformed metal under the indenter should be expressed as the plastic energy expended 12--
.,,,
_
,,
,,
(3)
which is essentially the same equation given by Stelmansbeko et al. [23]. in Eq. (3), the constant a is normally taken as !/3 [23],/z is the shear modulus, b is the Burgers vector, M is the Taylor factor which was found to be 3.0 for both the f.c.c, and b.c.c, structures [ 18],/] is the inclination angle between the face of the indenter and tbe surface of the material being indented (22' for the Vickers indenter) and D is the indentation depth. According to Eq. (3), if the two components of the dislocation density (O, and Ps) add in a linear manner, there should be a linear proportionality between a//2 and D - ~plot. Poole et at. [22] suggested that the slope of this plot will be steep if the plasticity of the material being tested is low,
$ I= "R
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x: v 4
-
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I
I
I
I
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I
0.8
1.0
12
1.4
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2.
(Indentation depth)' (gmj' Fig. 12. The data of Fig. i I plotted as the square of hardness vs. reciprocal of indentation depth according to Eq. (3).
H. ~'imeno~lu / Wear 210 (1997) 204-210 120
--
"
:
,=
Work hardened layer
Unhardened tlyer
100
A
z E
1. 20
0
t O0
05
1o
oo
05
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Indentation depth (pm) Fig. 13. Load-indentation depth plots of work hardened and unhardened layers shown in Fig. il. The dashed areas represent the plastic energy expended during hardness measurements.
per unit volume of indent. From this point of view, Berriche I241 proposed a new method for calculating the Vickers hardness by taking into account the ratio of the plastic energy dissipated to the volume of indentation produced. The plastic energies expended for the formation of a unit volume of indent were calculated as 3.5 J mm -3 for the work hardened layer and 2.2 J mm-3 for the unhardened layer. This result indicates that, to obtain the same indent volumes in both layers, about 60% more work should be done in the wo,k hardened layer than in the unhardened layer. Moore and Douthwaite [12] reported that the external work done in abrasion of ductile materials is related to the energy expended in plastic deformation of material from grooves and below the surface. Therefore, high plastic energy absorption by a unit volume of indent during hardness measurements can be correlated with the requirement of high external work for abrasive wear. In this case, the abrasion of the work hardened layer would be expected to occur by consuming high energy, during sliding. Thus, tf the surface of the material does not work harden, it will wear easily by consuming low energy. Wang and Lei [25] also suggested that a material which has a high work hardening exponent will consume greater energy during sliding and exhibit higher wear resistance.
4. Conclusions I. Microscopic studies show that sliding on abrasive particles not only causes wear of low carbon ferritic-pearlitic steels by ploughing, but also introduces plastic strains of more than 6 into the abraded surface. At this strain level, which is much higher than that of conventional deformations, ferrite grains deform by cross-slip and work harden about 50% of the original value. Plastic strain and hardness gradients extend to larger depths into the bulk with
209
increasing wear test load. The investigated 0.2%Ci.2%Mn steel exhibited a wear coefficient value of 25 X 10 - 3 mm 3 N -I m -I in abrasive wear tests performed under a load range between 16.4 N and 65.6 N. . Work hardening of the subsurface tends to restrict the plastic flow of the material under the groove and decrease the groove size. In this case, the external work necessary to form a unit volume of groove on the work hardened layer increases. Therefore, greater energy consumption is required for abrasion, during sliding. The higher the energy consumption, the better is the wear resistance. The indentation depth dependence of hardness observed during hardness measurements of worn specimens is in agreement with the dislocation model developed by Poole et al. 122]. This model is proposed for materials which can deform plastically by dislocation glide and explains the indentation size effect by the increase in dislocation density as the indent size decreases.
Acknowledgements "fhe author gratefully acknowledges the support of NATO TU PVD Coatings Project for the equipment supplied to the Metallurgical Engineering Department of ITU which was utilized in this study.
References I 1 [ ASM Metals Handbook, Properties and Selection: Iron and Steels, Vol, I. ASM. Metals Park, OH, 1978, 9th edn. 121 M. Hutchings. Tdbology: Friction and WearofEngineedngMaterials, Edward Arnold, London, 1992. [ 3 ]I.M. Hutchings, Wear resistant t~tefials: into the next century, Mater. Sci. Eng. A!84 (1994) 185-195. [ 4 ] D.A. Rigney. Viewpoint set on materials aspects of wear-introdection, Scri. Metall. 24 (1990) 799-803. [ 5 ] K. Kato. Wear mode transition, Scr. Metall. 24 (1990) 815-820. [ 61 L.H. Chen. Deformation, transfer and debris formation during sliding wear of metals, Scr. Metall. 24 (1990) 827-832. [7] D.A. Rigney. M.G.S. Naylor, H. Divakar, Low energy dislocation stn tures caused by sliding and particle impact. Mater. Sci. Eng. 81 ( 1986 ) 409-425. [8] J.L. Bas~, Role of microstrecture and mechanical properties in abrasion. Scr. Metall. 24 (1990) 821--826. 191 M.A. Moore. The relationship between the abrasive wear resistance. hardness and miemstructure of ferritie materials, Wear 28 (1974) 5968. ilOI M.M. Khru~hov, Principles of abrasive wear, Wear 28 (1974) 6988. Iill L. Xu. F. Kennon, A study of abrasive wear of carbon steels, Wear 148 ( 1991 ) 101-112. [121 M.A. Moore, R.M. Douthwaite, Plastic deformation below the worn surface, Metail. Trans. A 7 (1976) 1833-1839. [131 J. Zhang, A.T. Alpas, Delamination wear in ductile materials containing second phase particles, Mater. Set. Eng. AI60 (1993) 2535. 114l A.T. Alpas, H. Hu, J. Zhang, Plastic deformation and damage accumulation below ~be worn surfaces, Wear 162 (1993) 188-195.
210
1t. ~'imenol!iu ! Wear 210 (1997) 204-210
[151 B. Venkataraman, G. Sundarerajan, The sliding wear behavior of AISiC pa~culate composites !I, characterization of subsurface deformation and correlation with wear behavior, Acta Metall. 44 (1996) 461-473. [i6l W. Weiler, Hardness testing---a new method for economical and phy,~ically meaningful microhardness testing, Br. J. Non-Destruct. Test. 31 (1989) 253-258. [17] M.A. Meyers, K.K. Chawla, Mechanical Metallurgy, Prentice.Hall, Englewood Cliffs, NJ, 1984. [lSl T.H. Coertney, Mechanical Behavior of Materials, McGraw-Hill, Singapore, 1990. [Jg] P. Heilmann. W.A.T. Clark. D.A. Rigney. Orientation determination of subsurface cells generated by sliding. Acta Metall. 3 ! (1983) 12931305.
[201 A.T. AIpas, J. Zhang, Effect of microstructure (particulate size and volume fraction) and counterface material on the sliding wear resistance of paniculate reinforeed aluminum matrix composite, Metall. Mater. Trans. A 25 (1993) 969--983. [211 M.A. Moore, R.C.D. Richardson, D.G. Attwood, The limiting strength of worn metal surfaces. Metall. Trans. 3 (1972) 2485-2491. [221 WJ. Poole. M.F. Ashby, N.A. Flenk. Micrehardness of annealed and work ha~lened copper polyclystals, Scr. Mctall. 34 (1996) 559-564. [23] N.A. Stelmansheko, M.G. Walls, L.M. Brown. V. Milman, Microindentations on W and Mo oriented single crystals---an STM study. Acta Mater. Metall. 41 (1993) 2855-2685. [241 R. Berriche, Vickers hardness from plastic energy, Scr. Met,all. 32 (1995) 617-620. 1251 Y. Wang, T. Lei, Wear behavior of steel 1080 with different microstructures during dry sliding, Wear 194 (1996) 44-53.
Wear 210 ¢1997) 204-210
Subsurface characteristics of an abraded low carbon steel
Hilseyin ~imeno~lu lstanbul Technical Universi~*. Departmem of Metallurgical En&ineering. 80626 Maslak. istanbul, Turkey
Received 25 November 1996; accepted 4 March 1997
Abstraet Ductile metals commonly exhibit plastic deformation at and near the worn surface and their flow behaviour at large strains has a cleareffect on wear resistance. In this study, the characteristics of the near-surface region ofa ferritic-peaditic steel (0.2% C, 1.2% Mn), subjected to abrasive wear tests, were examined. Wear tests were performed under different loads by robbing the specimens on sliding 60 mesh A!203 abrasive band. The metallographic technique used to determine the magnitude of plastic deformation was based on measurement of the displacements of pearlite bands. The hardne~ of the plastic deformation zone was determined by performing ultramicrohardness tests along fertile bands with a Vickers indenter. Microscopic examinations of the near-surface regions revealed the wear mechanism to be ploughing and the deformation mechanism to be cross-slip. It was observed that plastic strain (more than 6) occurred on the abraded surface, and increased the hardness to about !.5 times the original value. The strain and hardness gradient extended to a larger depth into the bulk with increasing wear test load. It is concluded that the wear resistance of the investigated steel increases by work hardening of the near-surface region which is required to consume high energy for abrasion, during sliding. Ultramicrohardness measurements performed on worn specimens revealed high har4ness, as the indent size decreased. The indentation size-hardness relation was explained by adislocation model incorporating geometrically necessary dislocations due to the presence of strain gradients in the deformation region around the indent. © 1997 Elsevier Science S.A. Ke~n~,ords: Abrasivewear;.Indentation size effect: Low carbon steel: Uhram]crohardness'Work hardening
1, Introduction Wear is a surface phenomenon that causes alteration of the dimensions of materials by microchip formation due to the mechanical action of contacting materials. When a material surface is rubbed against another material under stress, the contact surface of the metal is worn and grooves are formed on it, in the direction parallel to sliding. In abrasive wear, material is removed or displaced from the surface by hard particles as they embed in the surface and move relative to it [ i-31. Sliding contact is often accompanied by plastic deformation and/or microfracture which influence the wear behaviour and may be associated with changes in the microstructure [ 4 - 6 ] . In ductile materials, the strains at the worn surfaces are much larger than conventional deformations, and the wear resistance correlates well with the surface hardness rather than the bulk hardness [2,7,8]. Previously the effect of bulk hardness on the abrasive wear performance of carbon steels has been studied [9,10]. Recently it has been reported ,hat. the bulk hardness can be used as a predictor only for annealed steels with the same type of microstructure [ 11 ]. 0043-1648/97/$17.00 © 1997ElsevierScience S.A. All rights reserved PIi S0043-1648 ( 97 )00051-3
Xu and Kennon [ 11] developed an abrasion grooving model for annealed steels, based on the groove size and thickness of the plastic deformation zone under the worn surface, by assuming the decrease in these parameters with increasing carbon content. However, they did not make any attempt to measure the thickness of the plastic deformation zone. The variatiop of the thickness of the plastic deformation zone with .,~iding distance, load and abrasive grid size has been studied in some non-ferrous metals [ 12-15]. The present study aims to perform metallographic examinations and hardness surveys on the longitudinal cross-sections of an annealed low carbon steel worn by sliding on abrasive particles, to determine the strain and hardness distribution under ,*he ,';am surface with respect to load. Special attention is given to the load-indentation depth-time data obtained from ultramicrohardness tests. 2, Experimental procedure The low carbon steel (0.21% C, 0.23% Si, i.18% Mn, 0.05% P, 0.05% S) utilized in this study was a hot rolled bar ( I0 mm diameter). Wear specimens with 3 X 3 mm 2 square tips w e ~ machined from the steel rod and annealed at 950
H. ~imeJ:J.,: '.. / Wear 210 (1997) 204-210
Lo~t~
(::) Fig. 1. Schematic view of the abrasive wear tester utilized in this inveztigation. °C for 1 h. The surfaces of the annealed specimens were etched with an HI= solution after grinding to reveal the micmstructure. Wear tests were carded out using the wear tester shown in Fig. 1, under four different loads ( 16.4 N to 65.6 N). At each load level two specimens were tested. During wear tests, specimens that were rubbing on 60 mesh AI203 abrasive band were moving perpendicular to the sliding direction so that they always passed over fresh abrasive. The sliding distance and the speed of the abrasive band were 4.6 m and 0.12 m s - t, respectively. The wear volume of the specimens was calculated from weight loss measurements, taken to the nearest 0.1 mg. Longitudinal sections of the worn specimens (on a plane normal to the wear surface and parallel to the sliding direction) were examined by optical and scanning electron microscopes. Hardness surveys of the sectioned specimens were made with a Fisher H 100 XYPROG ultramicrohardness tester [ 16]. During the hardness tests, load-indentation depth-time data were recorded as a Vickers indenter was driven into the sample (loading segment) and then withdrawn from it (unloading segment). A maximum test load of I00 mN was applied in 50 steps and the hold time at each step was 0.1 s. The loading/unloading rate of the indenter was 3.6 mN s - '.
3. Results and discussion
20
205 :
:
/
A
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10
0
5 0
~ 2O
4O
~0
8
Load (N) Fig. 2. The effect of abrasiveweartest loadon rig wearrateof the investigatedsteel.
3.2. Subsurface microstructure Metaliographic examinations of longitudinal sections revealed shear deformation under the worn surface. As shown in Fig. 3, the microstmcture of the investigated steel consisted of ferrite and pearlite bands elongated in the rolling direction. Below the worn surface, pearlite bands, which were initially perpendicular to the sliding abrasive, were displaced towards the slip direction. Ferfite grains were also bent and reoriented in the slip direction. Fine fenite grains were observed just beneath the worn surface (Fig. 4). It is reviewed [2,4,6] that the formation of the fine grained substructure under the worn surface is due to the heavy shear deformation introduced by wear, and such a sliding contact creates a near-surface layer having different structure and properties than the base matefiat. Fig. 5, which we,,,; obtained from a specimen etched before wear testing, shows the effects of abrasive wear on the subsurface. The rough surface and slip bands in ferrite grains indicate heavy plastic deformation concentrated in ferrite grains. A high magnification scanning electron microscopy image of the near-surface region (Fig. 6) reveals wavy slip
3.1. Wearrate Fig. 2 shows the effect of applied load on the wear rate, which is described as the volume loss per unit sliding distance ( r a m 3 m - t ). The wear rate increased from 0.6 mm 3 m - ~to 1.8 mm 3 m - t linearly when the load increased from 16.4 N tO 65.6 N. The slope of Fig. 2 (25 × 10 -3 mm 3 N - i m-~) is expressed as the dimensional wear coefficient k in the Archa~l equation [ 1-3]. The monotonic increase in wear rate with respect to load without showing any transient behavlout, indicates that the same wear mechanism is effective through the entire load range. Therefore, in the following sections, only the results of metallographic and hardness examinations of worn specimens under the test loads of 16.4 N and 65.6 N are given and discussed.
Fig. 3. The banded microm'actam o f t l ~ iavesdgmed low carbo~ steel subjcctcd to abrasive wear testing under a load of 16.4 N.
206
1t. ~'imeno~lu / Wear 210 (1997) 204-210
Fig, 7 shows a groove and the flow pattern beneath it, caused by penetration of abrasive to the worn surface. Thus, penetration of the abrasive particle displaced material from the groove to the ridges at the sides of the groove (by ploughing) and this displaced material was then removed as microchips. In a previous investigation [ 11], the abrasive wear mechanism of annealed low carbon steels was reported as microcutting and microploughing, which is in agreement with this observation.
3.3. Subsurface plastic strain
1,L Fig. 4, Optical micrograph oflhe subsurface showing fine fertile grainsjust beneath tbe worn surface (test load .65.6 N).
Rigney el al. [ 7 ] noticed that some microstructural features such as grain boundaries, twin boundaries and lamellae can be utilized to provide information about the deformation characteristics of layers below the worn surfaces and that optical microscopy was sufficient for this purpose. In the present study, the strain dis~bution under the worn surface was obtained by measuring the displacements of several pearlite bands in the sliding direction. Measurements were taken on printouts of the optical images that were transferred to a PC, equipped with a videographic card, and magnified. The displacements of pearlite bands at each depth were utilized to determine the shear deformation. The effective plastic strain at any given depth was calculated from the shear angle of pearlite bands ~ [ 12-14] ; e = (~/3/3) tan ~
Fig. 5. Opticalmicrographof the longitudinalsectionof the.wornspecimen (testload 65.6 N) etcbed before wear testing.
( I)
The variation of average effective strain with distance from the worn surface is plotted in Fig. 8. It should be mentioned that the standard deviation of the mean values was about I. 1 at depths le,ss than 20 Ixm, and 0.2 at greater depths, Fig. 8 shows that at higher loads, a strain gradient exists up to larger depths. AccordirJg to the results of abrasive wear tests conducted on copper-silver solder material, the strain as a function of depth is proportional to the square root of load [ 12]. Al-SiC composites also exhibited an increase in the thickness of the plastic deformation zone with load [ 13,15]. The strain profile of Fig. 8 can be described by an exponential [ 12,141;
Fig, 6. Scanning electron microgmph of the near-surface layer of the worn specimen (test load 65.6 N). Ferrit¢ and pearlite grains are ::hown by the symbols F and P respectively.
lines in ferrite grains which are evidence of cross-slip. In the literature [ 17,18], it is reported that, at large plastic s~.rains, the deformation mechanism of metals is controlled by crossslip which is associated with cell formation in the parabolic (stage Ill) hardening region. Transmission electron microscopy examinations revealed dislocation cell and subgrain formation under the worn surfaces of copper based metals, subjected to sliding wear and erosion tests [ 7,191.
Fig. 7. Optical micrograph of the longitudinal section showing a groove and flow lines beneath it (test load 64.6 N),
H. (?imeno~,ht/ Wear 210 ( ! 997) 204-216 m
.,,,..
$1Rnncd ~ •
[
IQ4 N O~IBN
5
°i
1
! 20
40
60
80
100
1~o
Distance from worn surface (pm) Fig. 8. The variation of effective strain with distance from the worn surface.
(2)
E=A exp(--BZ)
where Z is the surface depth, A is the extrapolated strain at the surface and B is a constant related to the strain rate at the surface [ 7 ]. The numerical values of A and B were calculated from Fig. 9 as A = 9, I i~m Ixm- i and B - 0,09 i,tm- ' for 16.4 N and A = 6.7 p.m i.tm-n and B---0.05 p.m- ~ for 65.6 N. Alpas and Zhang [20] also noted that the values of coefficients A and B depend on load. According to Fig. 9, abrasion under the loads of 16.4 and 65.6 N caused plastic strains of about 7 and 9 at the worn st'Maces of the investigated steel respectively. These strain levels reached at the worn surfaces are extremely high compared with those reached under conventional deformation processes.
3.4. Subsurface hardness In addition to the measurements of strain gradients below the abraded surface, the hardness distribution in this region
207
was determined. Since heavy deformation was ob~rved in ferrite grains (Figs. 5 and 6), ultramicrohardness measurements were carried out along ferrite bands with 25 Ixm interval. On each specimen, hardness tests were performed at four different locations and average values are plotted in Fig. 10. The standard deviations of the mean values were about 10% of the mean. Fig. l0 shows that the layers that were strained owing to abrasion (Fig. 9) were hardened, and the hardness gradient extends to a I~rger depth into the bulk with increasing load. The increase in hardness of fertte (about 50% with respect to the original hardness) at the worn surface can be attributed to work hardening. Some non-ferrous f.c.c, metals subjected to abrasive wear exhibited a hardness !ncrement two or three times that of the bulk at their worn surfaces [2,12,14]. It is well known that work hardening occurs as a result of dislocation interaction and, therefore, the degree of hardening is a function of dislocation density [ 17,18 ]. According to Moore et al. [ 21 ], hardening of the worn metal surface due to dislocation interactions is small compared with interstitial-dislocation interactions in b.c.c, solid solutions even at high dislocation densities, and b.c.c, metals show lower degrees of work hardening than f.c.c, metals at their worn surfaces.
3.5. hldentation si=e effect Ultramicrohardness measurements performed on the longitudinal sections of the worn specimens exhibited indentation depth dependence of hardness. Fig. I I compares the variation of hardness with indentation depth for work hardened and unhardened layers of the worn specimen tested at 65.6 N. In both layers the hardness increases as the indent .~:.ze decreases. This phenomenon is called the indentation size effect. It should be mentioned that the hardness values plotted in Fig. 10 corresponded to the maximum indentation depths under 100 raN. As shown in Fig. i i, maximum depths of
2O
164N 656N
10
Symbol
B
Load
.c_ tO
i'1
27.~
v
W IU e-
~
-r
•, 3 0
I
I
0
20
:
I
40
" =
I
I
60
80
"
l
100
12
Distance from worn s u r f a c e ( g m ) Fig. 9. Plot of/_~ effective strain against distance from the worn surface, to determine the values of coefficients A and B of Eq. (21.
"=7~ 0
=
l
1
!
I
20
40
60
80
100
12
Distance from wom surface (gin) Fig. 10. The variation of hardness with distance from Ihe worn suff~ce.
208
14. ~'imeno~lu/ Wear 210 (1997) 204-210 m
- -
.,
°t 0
"
I
I
I
I
I 15
Indentaeo. depth (pro) Fig. 11. Hardness as a function of indentation depth for work hardened (dislince from worn surface 10 tun) and m f l l d e n e d (distance from worn surface I i0 pan) layers of the specimen wont under 65.6 N.
indent of the work hardened and unhardened layers are 1.07 g m and 1.39 Ixm, respectively. This indicates that, under similar wear test conditions (load, grid size, etc.), work hardening increases the resistance of the contact surface to penetration of abrasive particles and decreases the groove size, compared with bulk material. Therefore, better correlation is often found in ductile metals between the wear resistance and the hardness of the worn surface rather than the hardness of the bulk material [2,81. Recently, Poole et al. [22l developed a dislocation model to analyse the indentation size effect on materials that can deform plastically by glide of dislocations. They tested their model experimentally on fully annealed and fully work hardened copper. This dislocation model is based on the creation of geometrically necessary dislocations around the indemand assumes the dislocation density is the sum of the background dislocation density O, and geometrically necessary dislocation density Or Since hardness is proportional to flow stress which is, in turn, pmpoRional to the square root of dislocation density, the indentation size dependence of hardness was expressed by the equation [ 22]; /42 ffi (3ap.bM) 2(p, + ta,t2[i/2bD)
because of the occurrence of a high strain gradient in the deformation region around the indent, and they observed experimentally a steeper slope for worked copper than for annealed copper. Fig. 12 shows that an approximately linear relationship exists for both work hardened and unhardened layers of the worn specimen. The slopes of the work hardened and unhardened layers are i.46 GPa 2 t i m - ' and 1.03 GPa 2 I~m-' respectively. From Eq. (3) one can calculate the slope d(HZ)/(D - ' ) theoretically as !.03 GPa z t t m - ' (by taking /~ = 75 GPa and b ffi 2.5 × 1 0 - ' ttm for a low carbon ferritic steel) which corresponds m the slope of the unhardened layer. The higher slope of the work hardened layer than the unhardened layer implies a localized plastic deformation region around the indent in the work hardened layer. However, in the unhardened layer, which has a high work hardening rate, a rather large plastic deformation region is generated under the indenter [ 17]. This implies that work hardening of the abraded surface limits the spread of the material under the groove over a wide volume, by plastic flow.
3.6. Subsurface plastic energy The ultramicrohardness tester utilized in this study was also capable of generating plots of indentation depth vs. load. These data are given in Fig. 13. for work hardened and unhardened layers discussed in Section 3.5. The area between loading and unloading segnmnts of the load-indentation depth plot gives the plastic energy expended during indentation, which was obtained for the work hardened and unhardened layers as 47.66 nJ and 34.92 nJ respectively. S i ~ e these layers have different plasticity, plastic energies expended during indentation create different indent volumes ( V = 8.17LP) on the surfaces being tested. Therefore, the energy absorbed by plastically deformed metal under the indenter should be expressed as the plastic energy expended 12--
.,,,
_
,,
,,
(3)
which is essentially the same equation given by Stelmansbeko et al. [23]. in Eq. (3), the constant a is normally taken as !/3 [23],/z is the shear modulus, b is the Burgers vector, M is the Taylor factor which was found to be 3.0 for both the f.c.c, and b.c.c, structures [ 18],/] is the inclination angle between the face of the indenter and tbe surface of the material being indented (22' for the Vickers indenter) and D is the indentation depth. According to Eq. (3), if the two components of the dislocation density (O, and Ps) add in a linear manner, there should be a linear proportionality between a//2 and D - ~plot. Poole et at. [22] suggested that the slope of this plot will be steep if the plasticity of the material being tested is low,
$ I= "R
6--
x: v 4
-
06
I
I
I
I
•
I
0.8
1.0
12
1.4
I~6
18
2.
(Indentation depth)' (gmj' Fig. 12. The data of Fig. i I plotted as the square of hardness vs. reciprocal of indentation depth according to Eq. (3).
H. ~'imeno~lu / Wear 210 (1997) 204-210 120
--
"
:
,=
Work hardened layer
Unhardened tlyer
100
A
z E
1. 20
0
t O0
05
1o
oo
05
10
Indentation depth (pm) Fig. 13. Load-indentation depth plots of work hardened and unhardened layers shown in Fig. il. The dashed areas represent the plastic energy expended during hardness measurements.
per unit volume of indent. From this point of view, Berriche I241 proposed a new method for calculating the Vickers hardness by taking into account the ratio of the plastic energy dissipated to the volume of indentation produced. The plastic energies expended for the formation of a unit volume of indent were calculated as 3.5 J mm -3 for the work hardened layer and 2.2 J mm-3 for the unhardened layer. This result indicates that, to obtain the same indent volumes in both layers, about 60% more work should be done in the wo,k hardened layer than in the unhardened layer. Moore and Douthwaite [12] reported that the external work done in abrasion of ductile materials is related to the energy expended in plastic deformation of material from grooves and below the surface. Therefore, high plastic energy absorption by a unit volume of indent during hardness measurements can be correlated with the requirement of high external work for abrasive wear. In this case, the abrasion of the work hardened layer would be expected to occur by consuming high energy, during sliding. Thus, tf the surface of the material does not work harden, it will wear easily by consuming low energy. Wang and Lei [25] also suggested that a material which has a high work hardening exponent will consume greater energy during sliding and exhibit higher wear resistance.
4. Conclusions I. Microscopic studies show that sliding on abrasive particles not only causes wear of low carbon ferritic-pearlitic steels by ploughing, but also introduces plastic strains of more than 6 into the abraded surface. At this strain level, which is much higher than that of conventional deformations, ferrite grains deform by cross-slip and work harden about 50% of the original value. Plastic strain and hardness gradients extend to larger depths into the bulk with
209
increasing wear test load. The investigated 0.2%Ci.2%Mn steel exhibited a wear coefficient value of 25 X 10 - 3 mm 3 N -I m -I in abrasive wear tests performed under a load range between 16.4 N and 65.6 N. . Work hardening of the subsurface tends to restrict the plastic flow of the material under the groove and decrease the groove size. In this case, the external work necessary to form a unit volume of groove on the work hardened layer increases. Therefore, greater energy consumption is required for abrasion, during sliding. The higher the energy consumption, the better is the wear resistance. The indentation depth dependence of hardness observed during hardness measurements of worn specimens is in agreement with the dislocation model developed by Poole et al. 122]. This model is proposed for materials which can deform plastically by dislocation glide and explains the indentation size effect by the increase in dislocation density as the indent size decreases.
Acknowledgements "fhe author gratefully acknowledges the support of NATO TU PVD Coatings Project for the equipment supplied to the Metallurgical Engineering Department of ITU which was utilized in this study.
References I 1 [ ASM Metals Handbook, Properties and Selection: Iron and Steels, Vol, I. ASM. Metals Park, OH, 1978, 9th edn. 121 M. Hutchings. Tdbology: Friction and WearofEngineedngMaterials, Edward Arnold, London, 1992. [ 3 ]I.M. Hutchings, Wear resistant t~tefials: into the next century, Mater. Sci. Eng. A!84 (1994) 185-195. [ 4 ] D.A. Rigney. Viewpoint set on materials aspects of wear-introdection, Scri. Metall. 24 (1990) 799-803. [ 5 ] K. Kato. Wear mode transition, Scr. Metall. 24 (1990) 815-820. [ 61 L.H. Chen. Deformation, transfer and debris formation during sliding wear of metals, Scr. Metall. 24 (1990) 827-832. [7] D.A. Rigney. M.G.S. Naylor, H. Divakar, Low energy dislocation stn tures caused by sliding and particle impact. Mater. Sci. Eng. 81 ( 1986 ) 409-425. [8] J.L. Bas~, Role of microstrecture and mechanical properties in abrasion. Scr. Metall. 24 (1990) 821--826. 191 M.A. Moore. The relationship between the abrasive wear resistance. hardness and miemstructure of ferritie materials, Wear 28 (1974) 5968. ilOI M.M. Khru~hov, Principles of abrasive wear, Wear 28 (1974) 6988. Iill L. Xu. F. Kennon, A study of abrasive wear of carbon steels, Wear 148 ( 1991 ) 101-112. [121 M.A. Moore, R.M. Douthwaite, Plastic deformation below the worn surface, Metail. Trans. A 7 (1976) 1833-1839. [131 J. Zhang, A.T. Alpas, Delamination wear in ductile materials containing second phase particles, Mater. Set. Eng. AI60 (1993) 2535. 114l A.T. Alpas, H. Hu, J. Zhang, Plastic deformation and damage accumulation below ~be worn surfaces, Wear 162 (1993) 188-195.
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[151 B. Venkataraman, G. Sundarerajan, The sliding wear behavior of AISiC pa~culate composites !I, characterization of subsurface deformation and correlation with wear behavior, Acta Metall. 44 (1996) 461-473. [i6l W. Weiler, Hardness testing---a new method for economical and phy,~ically meaningful microhardness testing, Br. J. Non-Destruct. Test. 31 (1989) 253-258. [17] M.A. Meyers, K.K. Chawla, Mechanical Metallurgy, Prentice.Hall, Englewood Cliffs, NJ, 1984. [lSl T.H. Coertney, Mechanical Behavior of Materials, McGraw-Hill, Singapore, 1990. [Jg] P. Heilmann. W.A.T. Clark. D.A. Rigney. Orientation determination of subsurface cells generated by sliding. Acta Metall. 3 ! (1983) 12931305.
[201 A.T. AIpas, J. Zhang, Effect of microstructure (particulate size and volume fraction) and counterface material on the sliding wear resistance of paniculate reinforeed aluminum matrix composite, Metall. Mater. Trans. A 25 (1993) 969--983. [211 M.A. Moore, R.C.D. Richardson, D.G. Attwood, The limiting strength of worn metal surfaces. Metall. Trans. 3 (1972) 2485-2491. [221 WJ. Poole. M.F. Ashby, N.A. Flenk. Micrehardness of annealed and work ha~lened copper polyclystals, Scr. Mctall. 34 (1996) 559-564. [23] N.A. Stelmansheko, M.G. Walls, L.M. Brown. V. Milman, Microindentations on W and Mo oriented single crystals---an STM study. Acta Mater. Metall. 41 (1993) 2855-2685. [241 R. Berriche, Vickers hardness from plastic energy, Scr. Met,all. 32 (1995) 617-620. 1251 Y. Wang, T. Lei, Wear behavior of steel 1080 with different microstructures during dry sliding, Wear 194 (1996) 44-53.