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Subsurface characteristics of a Fe–0.4 wt%C martenstic steel abraded with nanoindentation and cross-sectional TEM techniques
Wear 303 (2013) 92–97
Contents lists available at SciVerse ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Subsurface characteristics of a Fe–0.4 wt%C martenstic steel abraded with nanoindentation and cross-sectional TEM techniques Futoshi Katsuki n Advanced Technology Research Laboratories, Nippon Steel & Sumitomo Metal Corporation, 20-1, Shintomi, Futtsu, Chiba 293-8511, Japan
art ic l e i nf o
a b s t r a c t
Article history: Received 27 July 2012 Received in revised form 15 February 2013 Accepted 5 March 2013 Available online 14 March 2013
The present investigation examined the unidirectional abraded surfaces of a martensitic (0.4 wt%C) steel with silicon (1.5 wt%Si), chromium (1.5 wt%Cr) and molybdenum (1.9 wt%Mo) additions in order to elucidate the work hardening and softening near the surface layer caused by abrasion, particularly its relation to wear behavior. The abrasion testing was performed on a pin-abrasion apparatus, and a small pin of the specimen was ground on silica abrasive paper at an applied load of 2.1 N and a sliding speed of 0.66 m/s. The abraded surfaces were examined on a nanometer scale with a nanoindentation apparatus to evaluate the changes in nanohardness as the sliding time progressed. A cross-sectional transmission electron microscope (TEM) technique was also employed to clarify the structural changes in the region close to the abraded surface. The abrasion-induced work hardening with sliding time was observed in the case of chromium and molybdenum addition steels. A fine dispersion of molybdenum carbide (Mo2C) was observed in the surface of the molybdenum steel after abrasion. Mo2C precipitates at approximately 550 1C, indicating that surface and near-surface temperatures exceeded the carbide formation temperature after abrasion-induced frictional heating. In contrast, the silicon addition steel softened, which seems to be caused by abrasion heating that leads to some tempering effects. Work hardening and softening, which are caused by abrasion-induced subsurface deformation and frictional heating, respectively, seem to take place simultaneously and thus counteract each other's effect. Metallurgical reaction, such as precipitation and temper by frictional heating, has been found to play an important role in controlling the wear characteristics of steels. The present article discusses the influence of the alloying element addition on the wear response of the martensitic steel. & 2013 Elsevier B.V. All rights reserved.
Keywords: Abrasive wear Martenstic steel Nanoindentation Molybdenum carbide
1. Introduction Tribological properties (e.g., friction coefficient and wear) are largely determined by the condition and nature of surface layers [1]. The surface films most commonly recognized as wear resistant are those formed by a chemical reaction (such as nitriding) with the atmosphere or a lubricant. However, the topmost surface layers of metal are modified by friction, and Wilson has demonstrated that friction produces a layer of work-hardened metal, which constitutes the so-called Beilby layer [2]. Mechanical friction energy is also transformed into heat through surface and volumetric processes that occur in and around the real area of contact. Welsh found that the high frictional temperatures produced by dry rubbing of simple carbon steel produce gross structural changes and intense hardening of the surface layers, which has been attributed to martensite formation [3]. Modi et al. revealed that frictional heating can soften the region in the very close proximity of the abraded steel surface, which leads to some
* Tel.: þ81 439 80 4586; fax: þ 81 439 80 2910. E-mail address: [email protected] 0043-1648/$ - see front matter & 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.wear.2013.03.006
tempering effects [4]. These results indicate that thermal and mechanical contact phenomena are interrelated and that they can have a dominant influence on the wear response [5]. Attempts to correlate wear response with hardness are very common in the tribological literature [6]. In such studies, the hardness of the surface layer, as opposed to the bulk hardness, is of primary interest, because it reflects the properties of the zone in which wear is occurring. Cross-sectioning the worn surfaces has been conducted numerous times to assess the extent and the characteristics of abrasion-induced microstructure changes [7]. Taper sectioning, described in detail by Blau, has proven to be a more effective method for enhancing the resolution of fine details near the surface layer [8,9]. Blau studied taper sectioning's ability to depict the surfaces of worn copper alloy samples by using a micro-indentation hardness tester [8]. In a recent study of the subsurface characteristics of abraded pearlitic steel (0.4 wt%C), it was found that measuring the surface nanohardness with a depth sensing nanoindentation technique produces results that are in good agreement with the wear resistance, although there was no obvious correlation between these measurements and the bulk hardness measurements produced by Vickers indenters [10,11]. However, there have been no conscious efforts to investigate the
F. Katsuki / Wear 303 (2013) 92–97
changes to microstructures and hardness in steel surfaces, even though these surfaces are subjected to large deformation and high temperatures from abrasion. Steel has been used extensively for different engineering applications such as molds, crank shafts, grinding balls, gears, etc., because of its potential as a wear-resistant material. Medium carbon steel with a martenstic structure is normally considered the best choice for obtaining a good combination of strength, toughness and high wear resistance. The purpose of the present investigation of the unidirectional abraded surface of a martensitic (0.4 wt%C) steel is to examine the hardening, softening and change in the microstructure near the surface layer, especially these changes′ relation to the frictional heating. It also attempts to examine the role of each alloy in controlling wear behavior by testing three samples: one with silicon (1.5 wt%Si), one with chromium (1.5 wt%Cr) and another with molybdenum (1.9 wt%Mo). The nanoindentation technique was applied to very small material volumes to assess localized hardness [12]. Finally, this article presents a TEM investigation of cross-sectional samples that were excised with a focused gallium-ion beam to obtain an electron transparency of the first few micrometers of the contact surface [13].
93
Load 2.1N
Specimen 2 x 3.7mm2
Abrasive paper ( Silica #800) Disc
Sliding speed 0.66m/s (Sliding time 10, 20, 30, 50min) Fig. 1. Schematic illustration of the pin-abrasion apparatus.
dried and weighed prior to and after each wear test. The wear rate was calculated from weight loss measurement. Weight data were converted to volume loss using a steel density of 7760 kg/m3.
2. Experimental 2.1. Steel The martensitic steel utilized in this study was prepared with an induction furnace. Three hot rolled bars 20 mm in diameter were forged above 1100 1C from a bar 90 mm in diameter and were then cooled in an ambient air. Then, the steels were austenitized at 950 1C for 60 min and quenched with ice water held at room temperature. The chemical composition and bulk hardness as measured by a conventional Vickers indenter of steels appear in Table 1. Thermal diffusivity and specific heat were measured via a laser flash method using a thermal constant analyzer (TC-7000, Sinku-Riko). The measurements were made at three temperatures: 20, 400 and 600 1C. Thermal conductivity was calculated from the experimental values of the thermal diffusivity, specific heat and density [14]. 2.2. Abrasive wear tests The two-body abrasive wear test was performed on a small square tip (2 3.7 1.3 mm3) of each specimen. The tests were conducted with a pin-on-disk machine in an ambient atmosphere. A schematic representation of the test configuration is shown in Fig. 1. The testing surface (2 3.7 mm2) of every specimen was mirror-polished by emery paper and buffing and then finished by an electrolytic polishing with the Pt cathode in a solution of 5% percholic acid and 95% acetic acid at 8 1C under the potential of 40–50 V to remove the mechanically damaged surface layer. Afterwards, each specimen was ground on an abrasive paper at an applied load of 2.1 N, a sliding speed of 0.66 m/s and sliding times of 10, 20, 30 and 50 min. When in motion, every specimen ran on the same track under the air blow to remove debris particles. Crushed silica particles (size:15–67 µm) were used as the abrasive medium. The specimens were thoroughly cleaned, Table 1 Specimen's chemical composition (wt%) and bulk hardness according to the conventional Vickers indenter (Hv). No.
C
Si
Cr
Mo
Hv
1 2 3
0.4 0.38 0.35
1.46 o0.01 0.03
o 0.01 1.5 o 0.01
o 0.01 o 0.01 1.90
608 605 583
2.3. Nanoindentation Nanoindentation tests were performed on the steel surfaces, which were all worn under same wear conditions, using a commercially available apparatus (Triboscope, Hysitron Inc.). A Cube-Corner indenter was used for all indentations. The contact area of the indenter tip as a function of contact depth was calibrated by performing a series of indents on a standard fused silica sample. The analyses for the calculation of nanohardness followed the method used by Oliver and Pharr [15]. The testing load values were 200, 500, 1000 and 2000 µN, and the loading rate was a constant of 200 µN/s. A commercial type atomic force microscope (AFM; Nanoscope III, Digital Instruments, Inc.) with a nanoindentation apparatus was used for both the hardness test and surface observation. Probed sites and indent configuration on the specimen surfaces were confirmed before and after the indentation measurements. In order to obtain reliable results, more than 10 indentations were performed for each applied load at randomly selected locations on a worn surface. 2.4. Metallography After the abrasion tests, specimens were sectioned perpendicular to the wear surfaces. Longitudinal sections were examined in the scanning electron microscope (SEM) and transmission electron microscope (TEM) to clarify the structural changes in the region close to the abraded surface. TEM samples were prepared from the cross-section taken from the wear surfaces using the FIB in situ lift-out and H-bar techniques [16]. Thin foil samples for TEM observations of the worn surface were cut parallel to the wear direction and perpendicular to the worn surface inside the chamber of the dual beam FIB system (Hitachi FB-2000A with a Ga ion beam at 30 KV) using the lift-out procedure. The final thickness of the sample was reduced to approximately 100 nm and placed inside the TEM chamber for imaging. Abraded surfaces were also analyzed using the grazing incidence X-ray diffraction (GIXD) method. A commercial rotating anode X-ray diffractometer (Rigaku RINT-2500) combined with a thin-film measurement attached was employed [17].
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F. Katsuki / Wear 303 (2013) 92–97
3. Results 3.1. Abrasion tests The variation of wear depth (wear volume/testing surface area) with sliding time is shown in Fig. 2 for specimen Nos. 1–3. A considerable difference in wear depth for steels with different alloying elements was observed at a sliding time of 50 min. For No. 1 (1.5 wt%Si), the wear depth was found to be larger than that of the chromium (No. 2) and molybdenum (No. 3) addition steels, although the bulk Hv hardness of each specimen was almost equal. 3.2. Surface nanohardness The surface nanohardness of the specimens at a sliding time of 10, 20, 30 and 50 min are plotted as a function of the indentation depth in Fig. 3(a–c). On each specimen, nanoindentation was performed at four different loads—200, 500, 1000 and 2000 µN and the average values at each load are plotted in Fig. 3. Substantial abrasion-induced surface hardening was observed. There was also considerable hardening in the chromium (No. 2) and molybdenum (No. 3) addition steels, as shown in Fig. 3(b and c). On the other hand, as shown in Fig. 3(a), the extent of the silicon addition steel's hardening decreased with sliding time. 3.3. Surface microstructure
direction. The extent of the deformation that occurred very close to the abraded surface was so severe that it would not be possible to resolve the martensitic structure via SEM. TEM bright field (BF) images of the near surface region of specimen Nos. 1, 2 and 3 at a sliding time of 50 min reveal the detailed microstructural features, as shown in Fig. 5(a, b and c), respectively. As the observations approached the surface, increasing plastic deformation was found, which followed the gradient. Under the worn surface, a finegrained layer is visible on each micrograph. As shown in Fig. 5(c), a fine dispersion of precipitates (arrow) appears under the finegrained layer. Fig. 6 shows the multiple plots of GIXD profiles at lower incident angles of 0.5 deg for specimens Nos. 2 and 3. The XRD patterns clearly reveal the existence of ferrite. In the GIXD of Mo addition steel, the presence of a very weak peak at 2θ¼ 46.1 deg agrees with the appearance of the molybdenum carbide phase. The high-resolution image (Fig. 7) shows the details of the precipitation more clearly. The comparison of the bright and dark field image (Fig. 7(a and b)) and the selected diffraction pattern (Fig. 7(c)) reveals the existence of needle-like Mo2C precipitation [18].
4. Discussions The effects of frictional heating on the abrasion resistance of martensitic steel are considered here. Fig. 8 indicates that variations in surface nanohardness correspond to a contact depth of
Metallographic examination of the longitudinal section at a sliding time of 50 min reveals shear deformation under the worn surface (Fig. 4). Below the worn surface, the needle-shaped martensites of specimen No. 2, which were initially perpendicular to the abrasive surface, became displaced towards the sliding
Sliding direction
Wear depth, Wd / μm
60
No. 1 (1.5wt% Si) No. 2 (1.5wt% Cr)
40
20
No. 3 (1.9wt% Mo) 0
0
10
20
30
40
50
Sliding time, t / min
Nanohardness, Hnano / GPa
Fig. 2. Variations of the wear depth with sliding time for steels of different alloying elements.
Fig. 4. Scanning electron microscope image of a longitudinal section of specimen No. 2.
16
16
15
15
14
14
14
13
13
13
12
12
12
11
11
11 10
Softening 0
20
40
60
Indentation depth, d / nm
80 0
16
Hardening
20
40
Hardening
15
60
Indentation depth, d / nm
80 0
20
40
60
80
Indentation depth, d / nm
Fig. 3. Surface nanohardness of Nos. 1–3 at a sliding time of 10, 20, 30 and 50 min: (a) silicon (1.5wt%) (No. 1), (b) chromium (1.5wt%) (No. 2) and (c) molybdenum (1.9wt%) (No. 3) addition steel.
95
-Fe(110)
-Mo2C(110)
Intensity (arb. units)
Sliding direction
-Fe(200)
F. Katsuki / Wear 303 (2013) 92–97
(b) No.3 (1.9 wt% Mo)
(a) No.2 (1.5 wt% Cr)
2 (deg) Fig. 6. GIXD patterns from the abraded surfaces, (a) chromium (No. 2) and (b) molybdenum (No. 3) at 0.5 deg incidence angle.
Mo2C
1μm Fig. 5. Transmission electron micrographs of the wear on the subsurface layers of all three steels: (a) silicon (No. 1), (b) chromium (No. 2) and (c) molybdenum (No. 3).
less than 20 nm with sliding times for steels of differing alloying elements. As shown in Fig. 5(a–c), plastic deformation is observed closer to the surface following the gradient. The fine-grained layers are observed extending approximately 200 nm from the abraded surfaces. The plastic deformation zone of the indenter expands with increasing penetration depth of the indenter. Bhattacharya and Nix suggest that indentations to approximately 20% of the surface layer thickness sample the surface layer alone [19]. The maximum hardness of each specimen appeared to occur at a contact depth of less than 20 nm (Fig. 3). When the contact depth is 20 nm, the corresponding plastically deformed zone beneath the indenter would expand to a diameter of approximately 100 nm on
the hemispherical approximation, which agrees with the depth of the severely deformed layer for a fine grained structure, as shown in Fig. 5. As shown in Fig. 8, the presence of the fine grained layer agreed well with the work hardening and the sliding time of the chromium (No. 2) and molybdenum (No. 3) addition steels. On the other hand, surface softening with sliding time in the case of the silicon addition steel (No. 1) was noted (Fig. 8), even at the extreme subsurface layer. Therefore, silicon addition steel should exhibit lower resistance to abrasive wear than other steel alloys. In a tribological system, the majority of the frictional energy expended in plastic deformation is dissipated as heat, primarily through the top few microns of the contacting bodies [20]. The transformation of frictional energy to heat is responsible for increases in the temperature of the contact surface and could produce temperature gradients within the near-surface layers. As Fig. 3(c) reveals, beneath the fine-grained layer lies a region of a fine dispersion of molybdenum carbide (Mo2C) precipitates (arrow). Mo2C precipitates at approximately 550 1C [21], indicating that surface and near surface temperatures can surpass the carbide formation temperature through abrasion-induced frictional heating. Fig. 9 shows the temperature dependence of the thermal conductivity of the specimens. For the silicon addition steel (No. 1), which had low thermal conductivity, most of the frictional heat was retained in the abrasive paper–specimen interface. The effect of frictional heating is therefore relatively prominent in the first specimen [22]. This causes excessive temperatures in the silicon addition steel and leads to some tempering effects with sliding time in terms of residual stress relieving and surface softening. In contrast, the chromium (No. 2) and molybdenum (No. 3) steels possess high thermal conductivity. A large portion of the frictional heat transferred into the specimen, so the effect of friction heating on the temper of these two alloys was relatively small. Lower temperature causes a high work hardening exponent and resistance to abrasive wear. These results imply that frictional heat generation and thermal conduction are important factors for controlling the abrasive wear characteristics of the martensitic steel. Work hardening caused by abrasion-induced deformation and softening caused by frictional heating seem to be the two processes that take place simultaneously and counteract each other's effects.
5. Conclusions The effects of the abrasion-induced subsurface changes on the wear behavior of martensitic steel have been examined by using nanoindentation and TEM techniques. The findings suggest that the heat produced by abrasion must dissipate. Some heat goes into the abrasive paper, a portion transfers to the steel specimen and
F. Katsuki / Wear 303 (2013) 92–97
Nanohardness, Hnano / GPa
96
16
No. 3 (1.9wt% Mo)
15 14
No. 1
13
(1.5wt% Si) No. 2 (1.5wt% Cr)
12
0
10
20
30
40
50
Sliding time, t / ˚min
100nm
Fig. 8. Variations of surface nanohardness with sliding time for steels of different alloying elements.
60
No. 2 (1.5wt% Cr) 50
No. 3 (1.9wt% Mo)
40
30
No. 1 (1.5wt% Si)
100nm 20
0
200
400
600
800
Temperature, T / ˚C Fig. 9. Thermal conductivity of specimen Nos. 1–3.
11 20
02
suggestions. Thanks also goes to Dr. M. Umino, Mr. K. Yoshino and Mr. H. Tahira for their helpful support over the course of this study; to Ms. H. Horigome for her assistance with the experimental work; to Mr. K. Hanafusa for conducting the TEM observations; to Mr. H. Tamura for GIXD measurements.
00
References
Fig. 7. Transmission electron micrograph of the wear on the subsurface layer of molybdenum addition steel (No. 3): (a) bright field image, (b) dark field image, and (c) selected area diffraction pattern.
the remainder dissipates to the surroundings. The present study found direct microstructural evidence that abrasion increased the temperature of the steel specimens. Depending on the thermal conductivity and frictional energy generated during abrasion, this localized heating can reduce the magnitude of work hardening due to some tempering effects. Thus, frictional heating affected the steel alloy with silicon (No. 1) more than it affected the silicon-free alloys (Nos. 2 and 3) due to the silicon alloy's lower thermal conductivity. These results imply that metallurgical reactions, such as precipitation and temper by frictional heating, play an important role in controlling the wear characteristics of the steel.
Acknowledgments The author would like to thank Dr. Y. Okada and Dr. K. Miyata of Sumitomo Metal Industries Ltd. for their valuable discussions and
[1] D.A. Ringey, J.E. Hammerberg, Unlubricated sliding behavior of metals, MRS Bulletin 23 (6) (1998) 32–36. [2] R.W. Wilson, The contact resistance and mechanical properties of surface films on metals, Proceedings of the Physical Society B68 (1955) 625–641. [3] N.C. Welsh, Frictional heating and Its influence on the wear of steel, Journal of Applied Physics 28 (1957) 960–968. [4] O.P. Modi, B.K. Prasad, S. Das, A.K. Jha, A.H. Yegneswaran, Abrasive wear behavior of an AISI 5132 steel under low stresses, Materials Transactions, JIM 35 (1994) 67–73. [5] Olof Vimgsbo, Sture Hogmark, Fundamental of Friction and Wear of Materials, American Society for Metals, Meals Park, 1980, pp. 373–408; (papers presented at the 1980 ASM Materials Science Seminar). [6] D.A. Rigney, The roles of hardness in the sliding behavior of materials, Wear 175 (1994) 63–69. [7] S.F. Miller, P.J. Blau, J. Shih, Microstructural alteration associated with friction drilling of steel, aluminum, and titanium, Journal of Materials Engineering and Performance 14 (2005) 647–653. [8] P.J. Blau, Investigation of the nature of micro-indentation hardness gradients below slidint cnatacts in five copper alloys worn against 52100 steel, Journal of Materials Science 19 (1984) 1957–1968. [9] P.J. Mutton, J.D. Watson, Some effects of microstructure on the abrasion resistance of metals, Wear 48 (1978) 385–398. [10] F. Katsuki, M. Yonemura, Subsurface charateristics of an abraded Fe–0.4 wt%C pearlitic steel: a nanoindentation study, Wear 263 (2007) 1575–1578. [11] F. Katsuki, K. Watari, H. Tahira, M. Umino, Abrasive wear behavior of a pearlitic (0.4%C) steel microalloyed with vanadium, Wear 264 (2008) 331–336. [12] R. Li, L. Riester, T.R. Watkins, P.J. Blau, A.J. Shih, Metallurgical analysis and nanoindentation characterization of Ti–6Al–4V workpiece and chips in highthroughput drilling, Materials Science and Engineering: A 472 (2008) 115–124.
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[13] W. Österle, I. Dörfel, W. Gesatzle, H. Rooch, I. Urban, A characterization of tribological contacts with FIB and TEM, Praktische Metallographie 41 (2004) 166–179. [14] T. Goto, S. Hayashi, T. Hirai, Thermal properties of amorphous Si3N4–C composites prepared by chemical vapor deposition, Science Reports of the Research Institutes, Tohoku University. Series A, Physics, Chemistry and Metallurgy 29 (1980) 176–183. [15] W.C. Oliver, G.M. Pharr, Improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments, Journal of Materials Research 7 (1992) 1564–1583. [16] X. Meng-Burany, T.A. Perry, A.K. Sachdev, A.T. Alpas, Subsurface sliding wear damage characterization in Al–Si alloys using focused ion beam and crosssectional TEM techniques, Wear 270 (2011) 152–162. [17] T. Takayama, Y. Matsumoto, Effects of refraction and reflection on analysis of thin films by the grazing-incidence X-Ray diffraction method, Advances in XRay Analysis 33 (1989) 109–120.
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[18] S. Hinotani, Y. Ohmori, F. Terasaki, Effects of Fe3C and Mo2C precipitation on hydrogen diffusivity and hydrogen embrittlement in iron alloys, Materials Science and Engineering 76 (1985) 57–69. [19] A.K. Bhattacharya, W.D. Nix, Analysis of elastic and plastic deformation associated with indentation testing of thin films on substrates, International Journal of Solids and Structures 24 (1988) 1287–1298. [20] H. Uets, J. Föhl, Wear as an energy transformation process, Wear 49 (1978) 253–264. [21] R. Uemori, M. Tanino, Study of ultra-fine precipitates in low alloy steels, Journal de Physique C6 (1987) 399–404. [22] F.E. Kennedy Jr., Thermal and thermomechanical effects in dry sliding, Wear 100 (1984) 453–476.
Contents lists available at SciVerse ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Subsurface characteristics of a Fe–0.4 wt%C martenstic steel abraded with nanoindentation and cross-sectional TEM techniques Futoshi Katsuki n Advanced Technology Research Laboratories, Nippon Steel & Sumitomo Metal Corporation, 20-1, Shintomi, Futtsu, Chiba 293-8511, Japan
art ic l e i nf o
a b s t r a c t
Article history: Received 27 July 2012 Received in revised form 15 February 2013 Accepted 5 March 2013 Available online 14 March 2013
The present investigation examined the unidirectional abraded surfaces of a martensitic (0.4 wt%C) steel with silicon (1.5 wt%Si), chromium (1.5 wt%Cr) and molybdenum (1.9 wt%Mo) additions in order to elucidate the work hardening and softening near the surface layer caused by abrasion, particularly its relation to wear behavior. The abrasion testing was performed on a pin-abrasion apparatus, and a small pin of the specimen was ground on silica abrasive paper at an applied load of 2.1 N and a sliding speed of 0.66 m/s. The abraded surfaces were examined on a nanometer scale with a nanoindentation apparatus to evaluate the changes in nanohardness as the sliding time progressed. A cross-sectional transmission electron microscope (TEM) technique was also employed to clarify the structural changes in the region close to the abraded surface. The abrasion-induced work hardening with sliding time was observed in the case of chromium and molybdenum addition steels. A fine dispersion of molybdenum carbide (Mo2C) was observed in the surface of the molybdenum steel after abrasion. Mo2C precipitates at approximately 550 1C, indicating that surface and near-surface temperatures exceeded the carbide formation temperature after abrasion-induced frictional heating. In contrast, the silicon addition steel softened, which seems to be caused by abrasion heating that leads to some tempering effects. Work hardening and softening, which are caused by abrasion-induced subsurface deformation and frictional heating, respectively, seem to take place simultaneously and thus counteract each other's effect. Metallurgical reaction, such as precipitation and temper by frictional heating, has been found to play an important role in controlling the wear characteristics of steels. The present article discusses the influence of the alloying element addition on the wear response of the martensitic steel. & 2013 Elsevier B.V. All rights reserved.
Keywords: Abrasive wear Martenstic steel Nanoindentation Molybdenum carbide
1. Introduction Tribological properties (e.g., friction coefficient and wear) are largely determined by the condition and nature of surface layers [1]. The surface films most commonly recognized as wear resistant are those formed by a chemical reaction (such as nitriding) with the atmosphere or a lubricant. However, the topmost surface layers of metal are modified by friction, and Wilson has demonstrated that friction produces a layer of work-hardened metal, which constitutes the so-called Beilby layer [2]. Mechanical friction energy is also transformed into heat through surface and volumetric processes that occur in and around the real area of contact. Welsh found that the high frictional temperatures produced by dry rubbing of simple carbon steel produce gross structural changes and intense hardening of the surface layers, which has been attributed to martensite formation [3]. Modi et al. revealed that frictional heating can soften the region in the very close proximity of the abraded steel surface, which leads to some
* Tel.: þ81 439 80 4586; fax: þ 81 439 80 2910. E-mail address: [email protected] 0043-1648/$ - see front matter & 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.wear.2013.03.006
tempering effects [4]. These results indicate that thermal and mechanical contact phenomena are interrelated and that they can have a dominant influence on the wear response [5]. Attempts to correlate wear response with hardness are very common in the tribological literature [6]. In such studies, the hardness of the surface layer, as opposed to the bulk hardness, is of primary interest, because it reflects the properties of the zone in which wear is occurring. Cross-sectioning the worn surfaces has been conducted numerous times to assess the extent and the characteristics of abrasion-induced microstructure changes [7]. Taper sectioning, described in detail by Blau, has proven to be a more effective method for enhancing the resolution of fine details near the surface layer [8,9]. Blau studied taper sectioning's ability to depict the surfaces of worn copper alloy samples by using a micro-indentation hardness tester [8]. In a recent study of the subsurface characteristics of abraded pearlitic steel (0.4 wt%C), it was found that measuring the surface nanohardness with a depth sensing nanoindentation technique produces results that are in good agreement with the wear resistance, although there was no obvious correlation between these measurements and the bulk hardness measurements produced by Vickers indenters [10,11]. However, there have been no conscious efforts to investigate the
F. Katsuki / Wear 303 (2013) 92–97
changes to microstructures and hardness in steel surfaces, even though these surfaces are subjected to large deformation and high temperatures from abrasion. Steel has been used extensively for different engineering applications such as molds, crank shafts, grinding balls, gears, etc., because of its potential as a wear-resistant material. Medium carbon steel with a martenstic structure is normally considered the best choice for obtaining a good combination of strength, toughness and high wear resistance. The purpose of the present investigation of the unidirectional abraded surface of a martensitic (0.4 wt%C) steel is to examine the hardening, softening and change in the microstructure near the surface layer, especially these changes′ relation to the frictional heating. It also attempts to examine the role of each alloy in controlling wear behavior by testing three samples: one with silicon (1.5 wt%Si), one with chromium (1.5 wt%Cr) and another with molybdenum (1.9 wt%Mo). The nanoindentation technique was applied to very small material volumes to assess localized hardness [12]. Finally, this article presents a TEM investigation of cross-sectional samples that were excised with a focused gallium-ion beam to obtain an electron transparency of the first few micrometers of the contact surface [13].
93
Load 2.1N
Specimen 2 x 3.7mm2
Abrasive paper ( Silica #800) Disc
Sliding speed 0.66m/s (Sliding time 10, 20, 30, 50min) Fig. 1. Schematic illustration of the pin-abrasion apparatus.
dried and weighed prior to and after each wear test. The wear rate was calculated from weight loss measurement. Weight data were converted to volume loss using a steel density of 7760 kg/m3.
2. Experimental 2.1. Steel The martensitic steel utilized in this study was prepared with an induction furnace. Three hot rolled bars 20 mm in diameter were forged above 1100 1C from a bar 90 mm in diameter and were then cooled in an ambient air. Then, the steels were austenitized at 950 1C for 60 min and quenched with ice water held at room temperature. The chemical composition and bulk hardness as measured by a conventional Vickers indenter of steels appear in Table 1. Thermal diffusivity and specific heat were measured via a laser flash method using a thermal constant analyzer (TC-7000, Sinku-Riko). The measurements were made at three temperatures: 20, 400 and 600 1C. Thermal conductivity was calculated from the experimental values of the thermal diffusivity, specific heat and density [14]. 2.2. Abrasive wear tests The two-body abrasive wear test was performed on a small square tip (2 3.7 1.3 mm3) of each specimen. The tests were conducted with a pin-on-disk machine in an ambient atmosphere. A schematic representation of the test configuration is shown in Fig. 1. The testing surface (2 3.7 mm2) of every specimen was mirror-polished by emery paper and buffing and then finished by an electrolytic polishing with the Pt cathode in a solution of 5% percholic acid and 95% acetic acid at 8 1C under the potential of 40–50 V to remove the mechanically damaged surface layer. Afterwards, each specimen was ground on an abrasive paper at an applied load of 2.1 N, a sliding speed of 0.66 m/s and sliding times of 10, 20, 30 and 50 min. When in motion, every specimen ran on the same track under the air blow to remove debris particles. Crushed silica particles (size:15–67 µm) were used as the abrasive medium. The specimens were thoroughly cleaned, Table 1 Specimen's chemical composition (wt%) and bulk hardness according to the conventional Vickers indenter (Hv). No.
C
Si
Cr
Mo
Hv
1 2 3
0.4 0.38 0.35
1.46 o0.01 0.03
o 0.01 1.5 o 0.01
o 0.01 o 0.01 1.90
608 605 583
2.3. Nanoindentation Nanoindentation tests were performed on the steel surfaces, which were all worn under same wear conditions, using a commercially available apparatus (Triboscope, Hysitron Inc.). A Cube-Corner indenter was used for all indentations. The contact area of the indenter tip as a function of contact depth was calibrated by performing a series of indents on a standard fused silica sample. The analyses for the calculation of nanohardness followed the method used by Oliver and Pharr [15]. The testing load values were 200, 500, 1000 and 2000 µN, and the loading rate was a constant of 200 µN/s. A commercial type atomic force microscope (AFM; Nanoscope III, Digital Instruments, Inc.) with a nanoindentation apparatus was used for both the hardness test and surface observation. Probed sites and indent configuration on the specimen surfaces were confirmed before and after the indentation measurements. In order to obtain reliable results, more than 10 indentations were performed for each applied load at randomly selected locations on a worn surface. 2.4. Metallography After the abrasion tests, specimens were sectioned perpendicular to the wear surfaces. Longitudinal sections were examined in the scanning electron microscope (SEM) and transmission electron microscope (TEM) to clarify the structural changes in the region close to the abraded surface. TEM samples were prepared from the cross-section taken from the wear surfaces using the FIB in situ lift-out and H-bar techniques [16]. Thin foil samples for TEM observations of the worn surface were cut parallel to the wear direction and perpendicular to the worn surface inside the chamber of the dual beam FIB system (Hitachi FB-2000A with a Ga ion beam at 30 KV) using the lift-out procedure. The final thickness of the sample was reduced to approximately 100 nm and placed inside the TEM chamber for imaging. Abraded surfaces were also analyzed using the grazing incidence X-ray diffraction (GIXD) method. A commercial rotating anode X-ray diffractometer (Rigaku RINT-2500) combined with a thin-film measurement attached was employed [17].
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3. Results 3.1. Abrasion tests The variation of wear depth (wear volume/testing surface area) with sliding time is shown in Fig. 2 for specimen Nos. 1–3. A considerable difference in wear depth for steels with different alloying elements was observed at a sliding time of 50 min. For No. 1 (1.5 wt%Si), the wear depth was found to be larger than that of the chromium (No. 2) and molybdenum (No. 3) addition steels, although the bulk Hv hardness of each specimen was almost equal. 3.2. Surface nanohardness The surface nanohardness of the specimens at a sliding time of 10, 20, 30 and 50 min are plotted as a function of the indentation depth in Fig. 3(a–c). On each specimen, nanoindentation was performed at four different loads—200, 500, 1000 and 2000 µN and the average values at each load are plotted in Fig. 3. Substantial abrasion-induced surface hardening was observed. There was also considerable hardening in the chromium (No. 2) and molybdenum (No. 3) addition steels, as shown in Fig. 3(b and c). On the other hand, as shown in Fig. 3(a), the extent of the silicon addition steel's hardening decreased with sliding time. 3.3. Surface microstructure
direction. The extent of the deformation that occurred very close to the abraded surface was so severe that it would not be possible to resolve the martensitic structure via SEM. TEM bright field (BF) images of the near surface region of specimen Nos. 1, 2 and 3 at a sliding time of 50 min reveal the detailed microstructural features, as shown in Fig. 5(a, b and c), respectively. As the observations approached the surface, increasing plastic deformation was found, which followed the gradient. Under the worn surface, a finegrained layer is visible on each micrograph. As shown in Fig. 5(c), a fine dispersion of precipitates (arrow) appears under the finegrained layer. Fig. 6 shows the multiple plots of GIXD profiles at lower incident angles of 0.5 deg for specimens Nos. 2 and 3. The XRD patterns clearly reveal the existence of ferrite. In the GIXD of Mo addition steel, the presence of a very weak peak at 2θ¼ 46.1 deg agrees with the appearance of the molybdenum carbide phase. The high-resolution image (Fig. 7) shows the details of the precipitation more clearly. The comparison of the bright and dark field image (Fig. 7(a and b)) and the selected diffraction pattern (Fig. 7(c)) reveals the existence of needle-like Mo2C precipitation [18].
4. Discussions The effects of frictional heating on the abrasion resistance of martensitic steel are considered here. Fig. 8 indicates that variations in surface nanohardness correspond to a contact depth of
Metallographic examination of the longitudinal section at a sliding time of 50 min reveals shear deformation under the worn surface (Fig. 4). Below the worn surface, the needle-shaped martensites of specimen No. 2, which were initially perpendicular to the abrasive surface, became displaced towards the sliding
Sliding direction
Wear depth, Wd / μm
60
No. 1 (1.5wt% Si) No. 2 (1.5wt% Cr)
40
20
No. 3 (1.9wt% Mo) 0
0
10
20
30
40
50
Sliding time, t / min
Nanohardness, Hnano / GPa
Fig. 2. Variations of the wear depth with sliding time for steels of different alloying elements.
Fig. 4. Scanning electron microscope image of a longitudinal section of specimen No. 2.
16
16
15
15
14
14
14
13
13
13
12
12
12
11
11
11 10
Softening 0
20
40
60
Indentation depth, d / nm
80 0
16
Hardening
20
40
Hardening
15
60
Indentation depth, d / nm
80 0
20
40
60
80
Indentation depth, d / nm
Fig. 3. Surface nanohardness of Nos. 1–3 at a sliding time of 10, 20, 30 and 50 min: (a) silicon (1.5wt%) (No. 1), (b) chromium (1.5wt%) (No. 2) and (c) molybdenum (1.9wt%) (No. 3) addition steel.
95
-Fe(110)
-Mo2C(110)
Intensity (arb. units)
Sliding direction
-Fe(200)
F. Katsuki / Wear 303 (2013) 92–97
(b) No.3 (1.9 wt% Mo)
(a) No.2 (1.5 wt% Cr)
2 (deg) Fig. 6. GIXD patterns from the abraded surfaces, (a) chromium (No. 2) and (b) molybdenum (No. 3) at 0.5 deg incidence angle.
Mo2C
1μm Fig. 5. Transmission electron micrographs of the wear on the subsurface layers of all three steels: (a) silicon (No. 1), (b) chromium (No. 2) and (c) molybdenum (No. 3).
less than 20 nm with sliding times for steels of differing alloying elements. As shown in Fig. 5(a–c), plastic deformation is observed closer to the surface following the gradient. The fine-grained layers are observed extending approximately 200 nm from the abraded surfaces. The plastic deformation zone of the indenter expands with increasing penetration depth of the indenter. Bhattacharya and Nix suggest that indentations to approximately 20% of the surface layer thickness sample the surface layer alone [19]. The maximum hardness of each specimen appeared to occur at a contact depth of less than 20 nm (Fig. 3). When the contact depth is 20 nm, the corresponding plastically deformed zone beneath the indenter would expand to a diameter of approximately 100 nm on
the hemispherical approximation, which agrees with the depth of the severely deformed layer for a fine grained structure, as shown in Fig. 5. As shown in Fig. 8, the presence of the fine grained layer agreed well with the work hardening and the sliding time of the chromium (No. 2) and molybdenum (No. 3) addition steels. On the other hand, surface softening with sliding time in the case of the silicon addition steel (No. 1) was noted (Fig. 8), even at the extreme subsurface layer. Therefore, silicon addition steel should exhibit lower resistance to abrasive wear than other steel alloys. In a tribological system, the majority of the frictional energy expended in plastic deformation is dissipated as heat, primarily through the top few microns of the contacting bodies [20]. The transformation of frictional energy to heat is responsible for increases in the temperature of the contact surface and could produce temperature gradients within the near-surface layers. As Fig. 3(c) reveals, beneath the fine-grained layer lies a region of a fine dispersion of molybdenum carbide (Mo2C) precipitates (arrow). Mo2C precipitates at approximately 550 1C [21], indicating that surface and near surface temperatures can surpass the carbide formation temperature through abrasion-induced frictional heating. Fig. 9 shows the temperature dependence of the thermal conductivity of the specimens. For the silicon addition steel (No. 1), which had low thermal conductivity, most of the frictional heat was retained in the abrasive paper–specimen interface. The effect of frictional heating is therefore relatively prominent in the first specimen [22]. This causes excessive temperatures in the silicon addition steel and leads to some tempering effects with sliding time in terms of residual stress relieving and surface softening. In contrast, the chromium (No. 2) and molybdenum (No. 3) steels possess high thermal conductivity. A large portion of the frictional heat transferred into the specimen, so the effect of friction heating on the temper of these two alloys was relatively small. Lower temperature causes a high work hardening exponent and resistance to abrasive wear. These results imply that frictional heat generation and thermal conduction are important factors for controlling the abrasive wear characteristics of the martensitic steel. Work hardening caused by abrasion-induced deformation and softening caused by frictional heating seem to be the two processes that take place simultaneously and counteract each other's effects.
5. Conclusions The effects of the abrasion-induced subsurface changes on the wear behavior of martensitic steel have been examined by using nanoindentation and TEM techniques. The findings suggest that the heat produced by abrasion must dissipate. Some heat goes into the abrasive paper, a portion transfers to the steel specimen and
F. Katsuki / Wear 303 (2013) 92–97
Nanohardness, Hnano / GPa
96
16
No. 3 (1.9wt% Mo)
15 14
No. 1
13
(1.5wt% Si) No. 2 (1.5wt% Cr)
12
0
10
20
30
40
50
Sliding time, t / ˚min
100nm
Fig. 8. Variations of surface nanohardness with sliding time for steels of different alloying elements.
60
No. 2 (1.5wt% Cr) 50
No. 3 (1.9wt% Mo)
40
30
No. 1 (1.5wt% Si)
100nm 20
0
200
400
600
800
Temperature, T / ˚C Fig. 9. Thermal conductivity of specimen Nos. 1–3.
11 20
02
suggestions. Thanks also goes to Dr. M. Umino, Mr. K. Yoshino and Mr. H. Tahira for their helpful support over the course of this study; to Ms. H. Horigome for her assistance with the experimental work; to Mr. K. Hanafusa for conducting the TEM observations; to Mr. H. Tamura for GIXD measurements.
00
References
Fig. 7. Transmission electron micrograph of the wear on the subsurface layer of molybdenum addition steel (No. 3): (a) bright field image, (b) dark field image, and (c) selected area diffraction pattern.
the remainder dissipates to the surroundings. The present study found direct microstructural evidence that abrasion increased the temperature of the steel specimens. Depending on the thermal conductivity and frictional energy generated during abrasion, this localized heating can reduce the magnitude of work hardening due to some tempering effects. Thus, frictional heating affected the steel alloy with silicon (No. 1) more than it affected the silicon-free alloys (Nos. 2 and 3) due to the silicon alloy's lower thermal conductivity. These results imply that metallurgical reactions, such as precipitation and temper by frictional heating, play an important role in controlling the wear characteristics of the steel.
Acknowledgments The author would like to thank Dr. Y. Okada and Dr. K. Miyata of Sumitomo Metal Industries Ltd. for their valuable discussions and
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