Effect of grain size and hardness on fretting wear behavior of Inconel 600 alloys

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Effect of grain size and hardness on fretting wear behavior of Inconel 600 alloys J. Li, Y.H. Lu, H.Y. Zhang, L. Xin

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S0301-679X(14)00300-4 http://dx.doi.org/10.1016/j.triboint.2014.08.005 JTRI3414

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Received date: 10 April 2014 Revised date: 13 July 2014 Accepted date: 7 August 2014 Cite this article as: J. Li, Y.H. Lu, H.Y. Zhang, L. Xin, Effect of grain size and hardness on fretting wear behavior of Inconel 600 alloys, Tribology International, http://dx.doi.org/10.1016/j.triboint.2014.08.005 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

  Effect of grain size and hardness on fretting  wear behavior of Inconel 600 alloys   

J. Li, Y. H. Lu*, H.Y. Zhang, L. Xin National Center for Materials Service Safety, University of Science and Technology Beijing, Beijing 100083, P. R. China

Abstract:

The effects of grain size and bulk hardness on fretting wear behaviors were investigated by solution annealing and subsequently fretting wear test in Inconel 600 alloys. The results indicated that, with increase of solution temperature, the grain size increased while the hardness decreased. The average friction coefficients are the almost same, independent of grain size and hardness; while the wear volume decreased with increase of grain size, but the hardness played little role. The smaller grain was conductive to formation of tribological transformed structure (TTS) layer, and produced shorter delamiantion cracks in the TTS layer than larger one. Keywords: Fretting wear; Non-ferrous metal; Grain size; Crack

                                                                *   Author to whom correspondence should be addressed; E‐mail: [email protected]    1   

1. Introduction Fretting usually refers to small amplitude oscillatory movement occurring at tribosurfaces. They are encountered in many mechanical systems, such as bolted and riveted joints, turbine blade, leaf springs, shrink-fitted couplings, etc [1-4]. The degradation of surface and subsurface caused by fretting is a complex process involving friction, wear, plastic deformation, oxidation and crack [5]. Although the loss of material due to fretting may not be significant, it often leads to premature nucleation and subsequent growth of crack in the subsurface of the contact zone, which further degrades the functionality of mechanical system. Previous studies show that fretting sliding conditions can be either partial slip or gross slip, which is dependent of the contact condition[6]. Regarding the material’s response to fretting sliding conditions, adhesive wear and cracking are mainly encountered in partial slip regime, whereas abrasive and delamination wears associated with oxidation are observed in gross slip regime [7]. Lately, the concept of plastic ratcheting mechanism in sliding wear to predict delamination wear for metallic materials has also been applied to the computational analysis of fretting behavior [8-9]. Steam generator (SG) is one of the most important components in a nuclear power plant and its main function is to transfer the heat from the primary reactor coolant to make steam in the secondary side for driving turbine generator. However, damages and even failures occur during the heat exchange due to flow-induced vibration, which are attributed to the combination of high flow rate and small clearance between these tubes and their supports [10]. Because the tube vibration is 2   

inevitable, fretting wear damage has been considered as an important factor for degradation mechanisms in nuclear power plant. Many experimental works have been carried out to evaluate the fretting characteristics of nuclear engineering materials under different test parameters [11-13]. The main parameters affecting fretting wear are reported to be displacement amplitude, normal force, fretting cycles, environment and material composition [14-17]. Microstructure is also a critical factor influencing fretting wear. Under the fretting fatigue condition, the initiation and propagation of fatigue cracks are strongly dependent on grain size and distribution of carbides at grain boundary [18]. In fretting wear, wear volume is strongly related to subsurface deformation and fracture, which are dependent on the magnitude and distribution of local strain within a small volume of material in the subsurface[19]. Although the plastic deformation and fracture of material are closely linked to material microstructure, few studies have been focused on the relationship between the fretting wear and microstructure, which will be investigated in the Inconel 600 in present work.

2. Experimental To obtain different grain sizes, a series of heat treatments were conducted at the temperatures from 950 ℃ to 1150 ℃ at a regular interval of 50 ℃ for 10 minutes in Inconel 600 alloy. During heat treatment, all specimens were vacuum sealed in the quartz tubes to prevent oxidation. After heat treatment, the specimens were electrolytically etched in 10% oxalate solution at 5 V for 25 S to reveal the structural 3   

features. Grain size measurements were performed using a linear intercept method, and the value of grain size was approximately the average size of 300 grains. Bulk hardness was measured using a micro-Vickers hardness tester with a load of 20 g for 15 s. An Optimol SRV IV oscillating friction and wear tester with a ball-on-flat contact configuration was employed to carry out the fretting wear test. The test system was composed of electronics measurement, load unit, linear drive and test chamber. The obtained measurement data were recorded and saved via a computer system for subsequent analysis. The whole test cycle was controlled by software and can be adjusted to specific requirements. Test chamber installed with oscillation block was shown in Fig.1. The upper specimen was 304 stainless steel (abbreviated to 304SS) balls with a hardness of HV~200 and a diameter of 10 mm, and the lower specimen was cut to a dimension of 15 mm×15 mm×2.1 mm then conglutinated to a stainless base. The roughness (Ra) of 304SS ball was 2.0 and the flat specimens were polished to 3.0. The fretting wear tests were conducted under dry condition at room temperature (25-30 ℃) in air with a relative humidity of 30-40%. The fretting oscillation parameters were set as 100 N in normal load, 90 µm in displacement amplitude, 20 Hz in frequency and 30 minutes in test duration. The radius and initial Hertz contact stress was 151 µm and 2176 MPa, respectively. The specimens were ultrasonically cleaned in acetone before fretting test. After the fretting wear test, loose debris were removed by gentle brushing, compressed air and ultrasonic cleaning in acetone before the measurement of wear 4   

volume and profiles of worn scars using a 3D surface profilermeter (America ADE MicroXAM-3D). The morphologies of worn surface and cross-section were investigated using optical microscopy (OM) and scanning electron microscopy (SEM) at an operating voltage of 20 kV with energy dispersive spectroscopy (EDS). In order to characterize the cross-sectional morphology of wear scar, fretting scars were sectioned parallel to the direction of fretting at the edge using a diamond wire saw. The fretting direction corresponded to horizontal direction in all figures. Then the samples were mounted in a thermoplastic resin and polished to 3.0 in Ra finish with progressively finer diamond sprays.

3. Results Fig.2 shows the microstructures of as-received specimen and the ones after different solution treatments. It is found, for the as-received specimen in the mill-annealed condition, all samples exhibit fully recrystallized equiaxed grained structure, and there is no significant difference between the metallurgical structure of as-received specimen and solution treated one at 950 ℃. As the temperature increases to 1000 ℃, the grain begins to grow. When the temperature continues to increase, the grain size shows a significant increase. Fig.3 shows the variations of grain size and bulk hardness values with the solution temperature. It can be seen that the average grain size of as-received specimen is 17.4 µm, and slowly increases to 17.9 µm at a temperature of 950 ℃. When the temperature is further increased, a nearly linear increase in the grain size is 5   

found. A grain size of 100.4 µm is finally seen at the solution temperature of 1150 ℃. By comparison, the bulk hardness decreases with increasing solution temperature. The average hardness of as-received specimen is 235 HV, however it sharply decreases to 180 HV after solution treatment at 1000 ℃, and then slowly decreases to 160 HV after solution treatment at 1150 ℃. Fig.4 plots the friction coefficient as a function of cycle number at different solution temperatures. It can be found that the fiction coefficients in all samples obviously increase in the initial period and tend to a relative steady-state value with small fluctuation. In order to determine the effects of grain size and bulk hardness on wear behaviors of Inconel 600, the variation of average friction coefficient and wear volume with grain size and hardness are plotted in Fig.5 (a) and (b), respectively. It is found that the average friction coefficients in all samples are about 0.9, respective of grain size and hardness. By comparison, the wear volume nearly linearly increases with grain size (see Fig.5 (a)), but decreases nearly exponentially with hardness, as shown in Fig.5 (b). When the grain size increases from 17.4 µm to 100.4 µm, the wear volume increases from 48×106 µm3 to 84×106 µm3. While when the hardness increases from 160 HV to 180 HV, and wear volume greatly decreases from 84×106 µm3 to 52×106 µm3. Then there is a slow decrease in the wear volume (46×106 µm3) when the hardness further increases to 235 HV. These results indicate that small grain and high bulk hardness correspond to high wear resistance. Fig.6 shows the optical images of worn scars formed in the as-received specimen 6   

and the ones after solution treatment at different temperatures. It is found that all morphologies of worn scar are similar except for the size. The area of worn scar increases with increase of solution temperature. And worn scar size formed at solution temperature of 1150 ℃ is significantly larger than that of as-received specimen. Fig. 7 is EDS analysis of point F in Fig.6 (f). It is found that the worn surface contain about 15 wt% oxygen, which is believed that the worn surface is covered with oxide layer. The others worn surfaces have a similar result. In order to make a further detailed investigation on worn scar, three typical SEM surface morphologies of wear scars in as-received sample and solution treated ones at temperatures of 1000 and 1150 ℃ are presented in Fig.8. The locations of worn surface in Fig.8 respectively correspond to the marks A, C and F in Fig.6. It is found that no crack is observed in Fig.8, which indicates that the worn scar surfaces are completely covered by a continuous condense third body layer. Fig.9 shows cross-sectional SEM morphologies of wear scars formed in the as-received specimens and ones after solution treatment at different temperatures. The cross-sections in Fig.9 are located at A, B, C, D, E and F in Fig.6, respectively. It is found that many cracks are observed in the subsurface about 3~5 µm beneath the surface and propagate along the direction paralleled to the surface. In the case of solution temperature ≤ 1000 ℃, the length of main crack is usually about 15 µm surrounded by some secondary cracks, and the length of main crack distinctly increases to more than 30 µm at the solution temperature higher than 1000 ℃. In order to make clear of the evolution of microstructure in the subsurface of 7   

worn scars, the cross-section corroded micrographs of worn scars at different solution temperatures are shown in Fig.10. The cross-sections in Fig.10 are positioned at G, H, I, and J in Fig.6 respectively. The stratified structure are observed in the cross-section of worn scar as shown in Fig.10. They are third body layer (TBL), tribological transformed structure (TTS) layer，general deformation layer (GDL) and matrix respectively with

increase of depth from the topsurface. The TBL is composed of

metallic and oxidation particles with a thickness of 2~3 µm. The thickness of TTS layer is about 15 µm, in which the grain boundary can’t be observed. The detailed characterization of TTS layer has been conducted in our previous research and others [20, 21]. The GDL includes some deformed grains paralleled to the fretting direction with a thickness of 10 µm. Obviously, the elongated grains confirm the shearing plastic deformation takes place in the subsurface as response to the fretting process. It is noticed that with further increase of grain size, TTS layer is not observed in the specimens as the solution temperature is higher than 1000 ℃, as shown in Fig.10 (c) and (d).

4. Discussion It is well known that the growth of grain mainly depends on the migration of large angle boundary. The higher the solution temperature, the faster mobility of grain boundary is. And the hardness of metal strongly depends on the microstructure. As shown in Fig.3, it was believed that at the temperature of 950 ℃,the specimen was in the incubation period, a significant growth of grain size did not take place, however, 8   

the hardness dramatically decreased. While at the temperature of ≥1000 ℃, the movement of grain boundary became active, and grains began to grow with increase of the solution temperature, and hardness showed slow decrease. Previous study indicates the friction coefficient can be influenced by the real contact area, contacting state and wear debris, etc [5]. It is well known that debris particles formed during fretting are momentarily trapped in the contact zone and cannot easily escape, which play a significant role in the friction and wear processes in dry condition [22]. From OM and SEM surface morphologies of worn scar at different solution temperatures (see Fig. 6 and 8), typical annular scar of partial slip regime was not observed, the worn scars were completely covered by condensed third body layer, and as a result the average friction coefficient didn’t change with the grain size and hardness as shown in Fig.5. The increase of friction coefficient curve at the initiate period was attributed to work-hardening, while the fluctuation was due to the formation of wear particles and their ejection from the contact area. From Fig.5 (a) and (b), it was found that wear volume nearly linearly increased with grain size, but decreased nearly exponentially with bulk hardness. In the present study, the relationship of wear volume (Y) on grain size (X1) or bulk hardness (X2) could be considered individually as follows:

Y1 =39.1+0.455 X 1

(1)

Y2 =44.8+4.95 × 105 e( − X 2 /17.1)

(2)

It is well known that grain size and hardness are dependent through yield strength and flow stress [23]. From Fig.3 the dependence of hardness on grain size in 9   

present study can be described as follows: X 2 = 165.9 + 357.3( − X1 /9.4)

(3)

Then the Eq. (2) can be changed to: Y2 = 44.8 + 4.95 × 105 e( −165.9 −357.3e

( − X1/9.4)

/17.1)

(4)

In order to distinguish the effect of grain size and hardness on wear volume, a regression equation can be expressed as:

Y = K1 × Y1 + K 2 × Y2

K1> 0, K2> 0, K1+K2=1, Y1
(5)

By fitting, the multiple non-linear regression equation can be described as: Y = 39.1 + 0.445 X 1 + 0.668e( −20.9 e

( − X1/9.4)

)

(6)

Compared with Eq. (1), (4) and (6), it is calculated K1=0.978 and K2=0.022. The maximum value of K2×Y2 is 1.65 μm3 while the minimum value of K1×Y1 is 45.9 μm3 in present study. It is apparent that the value of K2×Y2 is very small and almost can be neglected compared with K1×Y1 value in all conditions. Therefore, it can be deduced that wear volume mainly depends on grain size; by comparison, bulk hardness was not a major factor in role of the wear volume. To predict the fretting wear behavior of metallic structures, it is essential to consider the elastic-plastic response of materials [23]. Fig.11 shows the typical cross-sectional profiles and SEM images of the corresponding region. It was found that there was a ridge at the edge of wear scar, and the region of ridge at 1150 ℃ was larger than that of as-received specimen, while at the outboard of wear scar edge corresponding to the ridge many slip bands were observed. The formation of ridge and slip bands were presumed to ratcheting effect due to the repeated pushing of 10   

spherical specimen against the edge, leading to localized plasticity. It is apparent that yield strength and grain size follow the Hall-Petch relationship:

σ s =σ o + K y d −1 2

(7)

Where σs was the yield strength, σ0 the resistance to deformation within grain, Ky the constant, d the grain size. The smaller grain size corresponded to the higher yield strength. Thus, on one hand, the lower yield strength of specimen in the condition of large grain size led to larger deformed area, and eventually resulted in large worn scar area. As a result, the lower contact stress formed on the worn scar corresponded to specimen with a larger grain size. With increase of grain size, this effect became obvious. On the other hand, Alloy 600 has a single phase with stable face-centered cubic (fcc) structure, and it is preferable to maintain the initial phase [24]. Therefore, it was expected that the dislocation structure was accelerated to accommodate the large stains generated during fretting wear and caused the change of microstructure in the subsurface of worn scar [25]. It was found from Fig.10 (a) and (b) that TTS layer, which was characterized by nanocrystallite [21], formed in the subsurface of worn scar in the small grain size condition. However, there was no such a TTS layer in the Fig.10 (c) and (d) corresponded to the larger grain size condition. It is proposed that the strain-induced nanocrystallites in Inconel 600 is due to the refinement of coarse grain caused by generation of mechanical microtwins, and subsequently, as the

formation of mechanical microtwins becomes too difficult

dislocation activity subdivides the nanometer-thick twin-matrix lamellae into equiaxed nanograins [26]. The research about the grain size effect on deformation 11   

twinning indicated that deformation twinning became easier with increasing grain size in the coarse-grain size range (larger than 1 µm) [27]. However, as mentioned above, lower contact stress corresponded to larger grain size. Due to low contact stress, the interaction of the microtwins with dislocations during the formation of nanocrystallite seemed to be hindered, which was not conductive to the formation of TTS layer at larger grain size. It has been reported that the specimens with nanocrystalline surface layer on pure bulk Cu exhibit a markedly enhanced fretting wear resistance relative to the coarse-grained counterpart, which is attributed to its high hardness [28, 29]. The study on fretting wear of Inconel 690 has also shown the protective nature of TTS layer. [13] Therefore, the lower wear volume in the small grain specimens can be attributed to the formation of TTS layer. There were many delamination cracks in the subsurface with the depth of 3~5µm from worn surface, as shown in Fig.9. The thickness of TTS in Fig.10 (a) and (b) was about 15 µm, while that of GDL layer in Fig.10 (c) and (d) was about 10 µm. It was apparent that the delamiantion cracks only propagated in TTS layer or GDL layer. From the research of Rigney, the nano-grain boundary of TTS layer can present suitable pathway for delamination crack propagation [25]. The main crack propagation in TTS layer was shorter together with secondary micro-crack than that of GDL layer.

12   

5. Conclusions The fretting tests on Inconel 600 under different solution annealing temperatures were conducted at room temperature in dry environment using a SRV IV fretting test rig. The effects of grain size and bulk hardness on the fretting behavior of Inconel 600 were analyzed. The observation and analysis led to the following conclusions: 1.

Solution annealing temperature had a strong effect on grain size and bulk

hardness of Inconel 600 alloy. The grain size increased with increase of the temperature, whereas the hardness decreased. 2.

The average friction coefficients in all samples are about 0.9, respective of grain

size and hardness. Small grain size and high hardness produced high wear resistance. The wear volume increased with increase of grain size, however, the bulk hardness played little role in wear volume. 3.

The specimen with smaller grain size possessed lower contact stress, which was

conductive to the formation of TTS layer in the subsurface of worn scar than that with larger grain size. 4.

The specimen with smaller grain size produced shorter delamination microcrack

in TTS than that with larger one.

Acknowledgements This work was supported by the National Key Basic Research Program of China (973 Program) (NO.2011CB610504).

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2009;267; 270–82. [10] Rubiolo PR, Young MY. On the factors affecting the fretting-wear risk of PWR fuel assemblies. Nucl Eng Des 2009; 239; 68–79. [11] Lee YH, Kim HK, Kim HD, Park CY, Kim IS. A comparative study on the fretting wear of steam generator tubes in korean power plants. Wear 2003; 255; 1198–208. [12] Kim DG, Lee YZ. Experimental investigation on sliding and fretting wear of steam generator tube materials. Wear 2001; 250; 673–80. [13] Chung I, Lee M, An experimental study on fretting wear behavior of cross-contacting Inconel 690 tubes. Nucl Eng Des 2011; 241; 4103–10. [14] Lee YH, Kim IS. The effect of subsurface deformation on the wear behavior of steam generator tube materials. Wear 2002; 253; 438–47. [15] Busquet M, Descartes S, Berthier Y. Formation conditions of mechanically modified superficial structures for two steels. Tribol Int 2009; 42; 1730–43. [16] Li LQ, Etsionc I, Talkeb FE. The effect of frequency on fretting in a micro-spherical contact. Wear 2011; 270; 857–65. [17] Tuckarta W, Iurmana L, Forlererb E, Influence of microstructure on tribologically mixed layers. Wear 2011; 271; 792–801. [18] Peng JF, Song C, Shen MX, Zheng JF, Zhou ZR, Zhu MH. An experimental study on bending fretting fatigue characteristics of 316L austenitic stainless steel. Tribol Int 2011; 44; 1417–26. [19] Terekhina S, Salvia M, Fouvry S. Contact fatigue and wear behaviour of 15   

bismaleimide polymer subjected to fretting loading under various temperature conditions. Tribol Int 2011; 44; 396–408. [20] Sauger E, Fouvry S, Ponsonnet L, Kapsa Ph, Martin JM, Vincent L. Tribologically transformed structure in fretting. Wear 2000; 245; 39-52. [21] Li J, Lu YH. Effects of displacement amplitude on fretting wear behaviors and mechanism of Inconel 600 alloy. Wear 2013; 304; 223–30. [22] Diomidis N, Mischler S. Third body effects on friction and wear during fretting of steel contacts. Tribol Int 2011; 44; 1452–60. [23] Fouvry S, Kapsa Ph, Vincent L. An elastic–plastic shakedown analysis of fretting wear. Wear; 247; 41–54. [24] Stiller K, Nilsson JO, Norring K. Structure, chemistry and stress corrosion cracking of grain boundaries in alloys 600 and 690. Metall Mater Trans A 1996; 27; 327–41. [25] Rigney DA, Glaeser WA. The significance of near surface microstructure in the wear process. Wear 1978; 46; 241–50. [26] Tao NR, Wu XL, Sui ML, Lu J, Lu K. Grain refinement at the nanoscale via mechanical twinning and dislocation interaction in a nickel-based alloy. J Mater Res 2004; 19; 1623–29. [27] Zhu YT, Liao XZ, Wu XL, Narayan J. Grain size effect on deformation twinning and detwinning, J Mater Res 2013; 48; 4467-75. [28] Zhang YS, Han Z, Lu K. Fretting wear behavior of nanocrystalline surface layer of copper under dry condition. Wear 2008; 265; 396-401. 16   

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Figure captions Fig.1. The test chamber of the SRV IV with installed oscillation block. Fig.2. Microstructures of the specimen (a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150 ℃. Fig.3. The grain size and bulk hardness values as a function of the solution temperature. Fig.4. The friction coefficient as a function of cycle number at different solution temperatures. Fig.5. The average friction coefficient and wear volume as function of (a) grain size and (b) hardness. Fig.6. OM images of worn scars in the specimen (a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150 ℃. Fig.7. EDS analysis of point F in Fig.6 (f). Fig.8. The SEM surface morphologies of worn scars (a) as-received, (b) 1000 and (c) 1150 ℃. Fig.9. The SEM cross-sectional morphologies of worn scars (a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150 ℃. Fig.10. The SEM cross-sectional corroded morphologies of worn scars (a) as-received, (b) 950, (c) 1000, (d) 1150 ℃. 17   

Fig.11. (a) typical cross-sectional profiles of as-received and 1150 ℃, (b) enlarged view of the ridge of 1150 ℃.

FIG.1. J.Li et al.

Fig.1. The test chamber of the SRV IV with installed oscillation block.

18   

FIG.2. J.Li et al.

Fig.2. Microstructures of the specimen (a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150 ℃. 19   

FIG.3. J.Li et al.

220

80

200

60

180

40

160

20 AS

950

1000

1050

1100 o

Solution temperature ( C)

Hardness value (HV)

Grain size (μm)

240

Grain size Hardness value

100

1150

  

Fig.3. The grain size and bulk hardness values as a function of the solution temperature.

20   

1.2 0.9 0.6 0.3 0.0 AS

Friction coefficient

FIG.4. J.Li et al.

o

950 C o

1000 C o

1050 C o

1100 C o

1150 C 0

10000

20000

30000

40000

Cycle number Fig.4. The friction coefficient as a function of cycle number at different solution temperatures.

21   

FIG.5. J.Li et al.

0.9

6

3

90

75

0.6

Wear volume 0.3 Friction coefficient Linear fit of Wear volume

60

45

Friction coefficient

Wear volume ( 10 μm )

(a)

0.0

20

40

60

80

Grain size (μm)

100

0.9

6

3

80

Wear volume Exponential fit of Wear volume Friction coefficient

70 60

0.6

0.3 50

Friction coefficient

Wear volume ( 10 μm )

(b)

0.0

40 160

180

200

220

240

Hardness (HV)                 Fig.5. The average friction coefficient and wear volume as function of

(a) grain size and (b) hardness.

22   

FIG.6. J.Li et al.

 

 

Fig.6. OM images of worn scars in the specimen 23   

(a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150℃.

FIG.7. J.Li et al.

Fig.7 EDS analysis of point F in Fig.6 (f)

24   

FIG.8. J.Li et al.

Fig.8. The SEM surface morphologies of worn scars (a) as-received, (b) 1000 and (c) 1150 ℃.

25   

FIG.9. J.Li et al.

Fig.9. The SEM cross-sectional morphologies of worn scars (a) as-received, (b) 950, (c) 1000, (d) 1050, (e) 1100 and (f) 1150 ℃.

26   

FIG.10. J.Li et al.

Fig.10. The SEM cross-sectional corroded morphologies of worn scars (a) as-received, (b) 950, (c) 1000, (d) 1150 ℃.

27   

FIG.11. J.Li et al.

(a) Hight of scar (μm)

30 15 0 -15 -30 -45 AS received Grain size 17.4μm -60 0

1000

o

1150 C Grain size 100.4μm 2000

3000

Distance across scar (μm)

    Fig.11. (a) Typical cross-sectional profiles of as-received and 1150 ℃,

(b) enlarged view of the ridge of 1150 ℃.

28   

HIGHLIGHT

z Effects of grain size and bulk hardness on fretting behavior of Inconel 600 alloy were studied.

z Smaller grain size and higher bulk hardness produced higher wear resistance.

z The grain size had a great effect on fretting wear resistance in the gross slipping regime.

z The bulk hardness played little role in fretting wear resistance. z The smaller grain was conductive to formation of TTS layer, and produced shorter cracks.

29