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Heat treatment of multi-element low alloy wear-resistant steel
Materials Science and Engineering A 396 (2005) 206–212
Heat treatment of multi-element low alloy wear-resistant steel Hanguang Fu a, 1 , Qiang Xiao b, ∗ , Hanfeng Fu c, 2 a
Department of Mechanical Engineering, Tsinghua University, Welding Building, Room 204, Beijing 100084, PR China b Department of Mechanical Engineering, University of Wyoming, P.O. Box 3295, Laramie, WY 82071, USA c First Manufacturing Steel Plant, Shanghai Baoshan Iron and Steel Group Corporation, Shanghai 201900, PR China Received 10 September 2004; received in revised form 8 January 2005; accepted 8 January 2005
Abstract The effect of heat treatment on the performances of multi-element low alloy wear-resistant steel (MLAWS), which is used to make the rolling mill torii liner, was investigated. The results show that the hardness and tensile strength increase as the quenching temperature is increased from 840 to 900 ◦ C. However, the hardness decreases rapidly as the quenching temperature is increased beyond 900 ◦ C, while the temperature has little influence on the tensile strength when it exceeds 900 ◦ C. No clear influence on the impact toughness has been observed unless the quenching temperature is beyond 920 ◦ C. As the tempering temperature exceeds 300 ◦ C, tiny ε carbides separate out from martensite and bainite complex structures and cause the carbon content in the complex structures to decrease. This results in the toughness to increase significantly. The best wear resistance can be obtained by tempering at 350 ◦ C. The optimum heat treatment of MLAWS comprises quenching at 900–920 ◦ C and tempering at 350–370 ◦ C. © 2005 Elsevier B.V. All rights reserved. Keywords: Rolling mill torii; Liner; Heat treatment; Mechanical performances; Wear resistance
1. Introduction In order to protect the stand, the liners have been installed on the torii of roughing mill [1]. The interval between the liner and the bearing chock of Baosteel ϕ 1300 mm roughing mill is 1.3 mm. When the interval becomes too large the liner needs to be replaced. Under rolling conditions, wear of the liner is normally very severe due to the poor service condition of the roughing mill [2]. Therefore, the wear-resistant explosive complex liners have been developed, which consist of Q235 steel plate and multi-element low alloy wearresistant steel (MLAWS) (for details, refer to patent number CN03125740.1) [3,4]. The influence of quenching and tempering temperatures on low alloy steel properties has been widely studied. It was found that the wear resistance of 40SiMnCrMoVRE steel increased as the quenching temperature is increased, and its ∗
Corresponding author. Tel.: +1 307 766 3187; fax: +1 307 766 2695. E-mail addresses: [email protected], [email protected] (Q. Xiao), [email protected] (H. Fu). 1 Tel.: +86 10 62773640; fax: +86 10 62773637. 2 Tel.: +86 21 66142017; fax: +86 21 66142017. 0921-5093/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2005.01.023
maximum value was reached at 950 ◦ C. When the tempering temperature was below 350 ◦ C, the wear resistance decreased slightly as the tempering temperature is increased. As the tempering temperature exceeded 350 ◦ C, the wear resistance increased rapidly, and the maximum wear resistance was reached at 400 ◦ C [5]. The typical heat treatment for medium carbon low alloy Cr–Mo–Ni–V steels involves a twostage austenitizing, followed by tempering and oil quenching [6,7]. Unfortunately, few reports are presently available for high carbon low alloy steel. Therefore, in this paper, we study the influences of quenching temperature, quenching cooling modes and tempering temperature on the hardenability, mechanical performances and wear resistance of high carbon MLAWS. As a result, we obtain an optimal heat treatment process.
2. Experimental MLAWS was melted in an arc furnace. Table 1 shows the composition of the MLAWS. The heat treatment was conducted on a RJX-18-13 resistance furnace. For impact
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Table 1 Chemical composition of sample (wt.%) Element
Content
C Si Mn Cr Ni Mo B Y K Na
0.6–1.0 0.5–1.0 0.8–1.2 0.5–1.5 0.4–0.6 0.4–0.8 0.005–0.015 0.05–0.10 0.05–0.15 0.05–0.15
toughness (αk ) tests, the specimen was of 10 mm × 10 mm × 55 mm without notch. Impact toughness (αk ) tests were performed on a JB30A Charpy impact testing machine. The size of the specimen for fracture toughness (K1c ) test was 10 mm × 10 mm × 55 mm, with a notch of 0.2 mm in depth and 2 mm in width. A 1251-Instron machine was used for the measurement of fracture toughness (K1c ) under force control [8]. The hardness was measured on an HR-150A Rockwell hardness tester. The tensile strength (σ b ), elongation (δ) and section reduction (ψ) tests were performed on a WE-30 universal testing machine [9]. The specimen size was Φ 10 mm × 130 mm. Wear test was carried out at room temperature with a MM200 wear apparatus under the load of 200 N [10]. YG6 hard alloy was used as the anti-wear material, and each test was run for 30 min. The reciprocal of wear amount per unit time represents the wear resistance of materials. The grain sizes and hardenability of austenite were measured through cementation process and end-quench test, respectively. In order to obtain the reasonable heat treatment parameters, the critical points of MLAWS were measured using the expanding method on a Gleeble-1500 thermal stress–strain simulation test machine [11,12].
3. Results and discussion 3.1. Effect of different cooling patterns on the hardenability The critical points of MLAWS are listed in Table 2. Ac1 is the lower critical point, Ac3 is the upper critical point, Ms is the starting temperature of martensitic transformation, and Mf is the finishing temperature of martensitic transformation. The grain structures at different heating temperatures are shown in Fig. 1. The grain sizes of the microstructure are Table 2 Critical points of material (◦ C) Ac1 Ac3 Ms Mf
782 845 294 98
Fig. 1. Grain microstructures at different heating temperature (held for 6 h): (a) 840 ◦ C; (b) 920 ◦ C; (c) 960 ◦ C.
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Fig. 3 shows the effects of quenching temperature on the tensile strength and hardness of MLAWS. The hardness
increases as the quenching temperature is increased from 840 to 900 ◦ C. However, the hardness decreases rapidly as the quenching temperature is increased beyond 900 ◦ C. This is in correlation with the content of carbon and alloy elements which is dissolved in high temperature austenite. At low quenching temperature, the contents of carbon and alloy elements are relatively small. This results in a quenching structure of bainite and martensite, plus some pearlite with low hardness and poor wear resistance [15,16]. When the quenching temperature is above 900 ◦ C, the contents of carbon and alloy elements dissolved in high temperature austenite increase significantly, and the stability of high temperature austenite increases, so the quenching structure contains low hardness retained austenite [5,17]. Therefore, the hardness drops rapidly [18,19]. In addition, it can be seen from Fig. 3 that the effect of quenching temperature on the tensile strengthen of MLAWS is similar to that on the hardness when quenching temperature is below 900 ◦ C, while the quenching temperature has little influence on the tensile strength when it exceeds 900 ◦ C. As discussed above, the carbon and alloy elements dissolved in high temperature austenite increase with increasing quenching temperature. Accordingly, the quenching structure contains more carbon and alloy elements, which strengthen quenching structures and improve the tensile strength [20]. As quenching temperature exceeds 900 ◦ C, the dissolved carbon and alloy elements continue to increase as the temperature increases, and the quenching structure continues to strengthen. At the same time, the retained austenite appears steadily in the quenching structure because the stability of matrix increases. As a result of this comprehensive function, the tensile strength does not change significantly [21]. Effects of quenching temperature on the impact toughness and the fracture toughness of MLAWS are shown in Fig. 4. When the quenching temperature is below 920 ◦ C, it has no significant influence on the impact toughness because the grain sizes do not change much [22,23]. As the quenching temperature increases above 920 ◦ C, the crystal grains grow up consistently, as shown in Fig. 1, and
Fig. 3. Effect of quenching temperature on strength and hardness of MLAWS.
Fig. 4. Effect of quenching temperature on impact toughness and fracture toughness of MLAWS.
Fig. 2. Effect of quenching cooling patterns on hardenability of MLAWS.
in grade 8–9 when the specimens are heated to 920 ◦ C and held for 6 h. This indicates that the grain sizes of austenite are fine and have a very good stability when heated at 920 ◦ C. The influence of cooling patterns on the hardenability of MLAWS is shown in Fig. 2. Since there are many alloy elements in MLAWS that can improve the hardenability of the material, the hardenability of MLAWS is excellent. However, under the condition of water and oil cooling, it is likely to produce crack due to high carbon content, poor heat conductivity and high temperature plasticity [13,14]. Therefore, fog cooling is adopted as the cooling pattern at all stages. 3.2. Effect of quenching temperature on mechanical properties
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Fig. 5. Effect of tempering temperature on strength and hardness.
cause the impact toughness to decrease slightly. Meanwhile, it can be seen from Fig. 4 that the quenching temperature has no clear influence on the fracture toughness of MLAWS. 3.3. Effect of tempering temperature on mechanical properties For tempering temperature influence investigation, MLAWS was quenched at 900 ◦ C and tempered for 2 h at different temperatures. The influence of tempering temperature on the hardness, tensile strength, impact toughness, fracture toughness, elongation and section reduction are shown in Figs. 5–7. Fig. 5 shows that the hardness does not change when the tempering temperature is below 150 ◦ C. Afterwards, the hardness decreases gradually as the temperature is increased. The hardness drops by about 5 HRC after tempering at 450 ◦ C when compared to as-quenched. The hardness tends to drop more rapidly when the tempering temperature
Fig. 6. Effect of tempering temperature on impact toughness and fracture toughness.
209
Fig. 7. Effect of tempering temperature on elongation and section reduction.
is above 450 ◦ C. In this case, the structure of MLAWS was transformed into pearlite from martensite because of high temperature tempering. Therefore, the hardness of MLAWS reduced more rapidly [24]. Similar behavior can be observed for the tensile strength, which also decreases with increase in tempering temperature. The slight difference between these results is that the tensile strength curve is smoother than that of hardness curve as the tempering temperature increases above 150 ◦ C. Fig. 6 shows the influence of tempering temperature on the impact toughness and the fracture toughness of MLAWS. Since the phase transformation stress does not fully dispel [25,26], its toughness is relatively low when tempering temperature is below 294 ◦ C. As the tempering temperature exceeds 300 ◦ C, tiny ε carbides separate out from the complex structures of martensite and bainite, and cause the carbon content in the complex structures to decrease, thereby increasing the toughness significantly [27–29]. When tempering temperature rises to 400 ◦ C, the amount of ε carbides reduces slowly, the decomposition of retained austenite occurs, and the brittle cementite separates out [26,30,31]. All these led to the reduction of impact toughness [32]. After that, a large amount of cementite begin to segregate from the complex structures, which decreases the carbon content in martensite and bainite, and results in the toughness of matrix to increase. Although the amount of cementite increases further, the integration effects show that the toughness of MLAWS begins to increase greatly as the tempering temperature increases above 450 ◦ C (Fig. 5). Meanwhile, the overall trend of the fracture toughness change is similar to the impact toughness. It varies with the increase of tempering temperature as shown in Fig. 6. Effects of tempering temperature on the elongation and section reduction of MLAWS are shown in Fig. 7. When tempering temperature is below 400 ◦ C, both the elongation and section reduction are relatively low. As tempering temperature increases above 400 ◦ C, the elongation and the section
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And W c = K1 Wf = K2
Fig. 8. Effect of tempering temperature on wear resistance.
reduction improve by a large margin because of the formation of pearlite in the structure [33,34]. 3.4. Effect of tempering temperature on the wear resistance Fig. 8 shows the influence of tempering temperature on the wear resistance of MLAWS after quenching at 900 ◦ C and tempered for 2 h at different temperatures. When tempering under 300 ◦ C, the effect of temperature on wear resistance is not very clear. However, the wear resistance improves significantly as tempering temperature exceeds 300 ◦ C, and reaches the maximum value at 350 ◦ C. After that, the wear resistance drops consistently as the temperature rise. The relationship between the wear resistance and tempering temperature is related to the hardness and the impact toughness of MLAWS. In the wear of a material, the wear rate consists of two parts [35,36], namely W t = Wc + W f
(1)
P H
(2) P
(εf H)2
(3)
where Wt is the wear rate of materials, Wc is the wear rate caused by cutting, Wf is the wear rate caused by fatigue, P is the pressure, H is the hardness, and εf is the fracture strain at single axis tensile, which reflects the toughness. From the above equations, it can be noted that the wear of a material mainly depends on the hardness and the toughness of the material. When tempering at 350 ◦ C, both the toughness and the hardness of MLAWS are relatively high, so the maximum value of the wear resistance was reached. With the further increase of tempering temperature, the toughness of MLAWS clearly improved, but the hardness drops by a large margin. Therefore, the wear resistance drops as a result of the combination effect. In addition, the special relation between wear resistance and tempering temperature has something to do with the elastic limit of MLAWS. It is well known that carbon steel and alloy steel containing 0.6–0.8% of carbon have supreme elastic limit after quenching and tempering at an intermediate temperature [5,37]. They can be flexibly replied and does not produce plasticity deformation after bearing greater deformation when used for making spring. The carbon content of MLAWS is 0.6–1.0%. There is a supreme elastic limit as well after quenching and tempering at intermediate temperature. The possible reasons for its wear resistance being higher than that of low temperature tempering and high temperature tempering are [38]: (1) Improving the elastic limit can make steel produce greater elasticity deformation at the receiving load and reduce the amount of plasticity deformation, while the elasticity deformation can be replied after the external
Fig. 9. Microstructure of MLAWS after heat treatment.
H. Fu et al. / Materials Science and Engineering A 396 (2005) 206–212
force is released, which will not cause wear and improve the wear resistance. (2) From the viewpoint of energy, improving the elastic limit of the steel can improve elastic specific power. With same wear load, there is more energy that can turn into elastic energy, and reduces the energy used for plasticity deformation, crack initiation and propagation. Therefore, it is favorable to the improvement of the wear resistance. Based upon the above analysis, the optimum heat treatment process of MLAWS is concluded as follows: fog quenching at 900–920 ◦ C + tempering at 350–370 ◦ C. Fig. 9 shows the microstructure of MLAWS after heat treatment. It consists of tempering martensite, bainite and retained austenite. This structure has excellent comprehensive mechanical properties and a good wear resistance. The above heat treatment techniques were applied to the wear-resistant compound liner for industry testing. The complex liner was used in Baosteel ϕ 1300 mm roughing mill torii. The result of on-line service investigation indicated that the wear for this liner was only 0.5 mm after 4 months, while the wear for common heat treatment liner was 2.0 mm after 4 months. It appears that the service life of complex liner improves by more than three times.
4. Conclusions This study was carried out to investigate the effect of heat treatment on the performances of multi-element low alloy wear-resistant steel used to make the rolling mill torii liner. This work supports the following conclusions: (1) For quenching temperature less than 900 ◦ C, the hardness and tensile strength increase as the quenching temperature increase. However, the hardness decreases rapidly as the quenching temperature exceeds 900 ◦ C, while the temperature has little influence on the tensile strength when it exceeds 900 ◦ C. No clear influence on the impact toughness has been observed unless the quenching temperature is beyond 920 ◦ C. (2) For the tempering temperature below 150 ◦ C, the hardness does not change. Afterward, the hardness decreases gradually as the temperature increases. The overall trend of the tensile strength changes is similar to the hardness. The impact and fracture toughness are relatively low when tempering temperature is below 300 ◦ C. As the tempering temperature exceeds 300 ◦ C, the toughness increases significantly. The toughness begins to increase greatly as the tempering temperature increases from 450 ◦ C. (3) When quenching at 900–920 ◦ C and tempering at 350–370 ◦ C, the microstructure with tiny martensite, bainite and particle carbides can be obtained. It has excellent comprehensive mechanical properties and a good wear resistance.
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(4) For satisfactory use of rolling mill torii liner, the optimum heat treatment technique of MLAWS comprises fog quenching at 900–920 ◦ C and tempering at 350– 370 ◦ C. (5) The wear amount of the complex liner is only 0.5 mm after using in the rolling mill torri for 4 months. The service life of complex liner improves by more than three times. References [1] R.C. Schrama, Lubr. Eng. 50 (1994) 438. [2] D. Viens, K. Albano, C. Churchill, Proceedings of the 39th Mechanical Working and Steel Processing Conference on Iron and Steel Society of AIME, Indianapolis, IN, USA, 19–22 October, 1997, pp. 467–474. [3] Z.J. Huang, H.G. Fu, patent number CN03125740.1. [4] H.G. Fu, Z.J. Huang, Special Steel 25 (2004) 46. [5] P.Q. Dai, S.X. Huang, Heat Treat. Met. 12 (1998) 19. [6] T. Wada, W.C. Hagel, Metall. Trans. 9A (1978) 691. [7] Z.C. Liu, J.P. Yan, L.P. Zhao, G. Wang, Q. Li, Y.X. Tian, J. Baotou Univ. Iron Steel Technol. 20 (2001) 30. [8] H.G. Fu, D.M. Fu, D.N. Zou, J.D. Xing, J. Wuhan Univ. Technol. Mater. Sci. Ed. 19 (2004) 48. [9] X.H. Cui, O.C. Jiang, S.Q. Wang, Z.M. He, Acta Metall. Sinica (English Letters) 13 (2000) 1053. [10] S.O. Wang, Q.C. Jiang, X.M. Sui, Z.M. He, Tribology 19 (1999) 33. [11] H.Y. Li, N.M. Xiao, L.P. Meng, X.J. Xie, J. Cent. South Univ. Technol. 33 (2002) 495. [12] F. Cao, X.N. Chen, Heat Treat. Met. 28 (2003) 24. [13] Y. Mikita, I. Nakabayashi, Trans. Jpn Soc. Mech. Eng. Part A 54 (1988) 1246. [14] B.J. Yin, X.H. Li, Hot Working Technol. 6 (2003) 46. [15] R. Yang, L.J. Li, Y.C. Li, Iron Steel 34 (1999) 41. [16] E. Anelli, M.C. Cesile, P.E. Di Nunzio, Mater. Sci. Technol. Meeting (2003) 453–466. [17] S.F. De, R.C. Rodriguez, C.A. Ratnapuli, Bottrelcoutinho, Metalurgia 40 (1984) 195. [18] Y. Tomita, K. Morioka, Mater. Char. 38 (1997) 243. [19] D.Y. Wei, J.L. Gu, H.S. Fang, B.Z. Bai, Z.G. Yang, Int. J. Fatigue 26 (2004) 437. [20] B.B. Vinokur, A. Vinokur, V.E. Shtessel, Metall. Mater. Trans. A 27A (1996) 2852. [21] J. Zeman, S. Rolc, J. Buchar, J. Pokluda, ASTM Special Tech. Publication 1074 (1990) 396. [22] S.I. Yokoyama, K. Kubota, H. Sasaki, Y. Minamino, J. Iron Steel Inst. Jpn. 90 (2004) 263. [23] C.L. Davis, J.E. King, Mater. Sci. Technol. 9 (1993) 8. [24] M. Gojic, L. Kosec, P. Matkovic, J. Mater. Sci. 33 (1998) 395. [25] C. Aubry, S. Denis, P. Archambault, A. Simon, F. Ruckstuhl, B. Miege, Metall. Italiana 91 (1999) 33. [26] H.G. Fu, J.Z. Wu, J. Xu, Res. Iron Steel 4 (1994) 17. [27] G.D. Pigrova, Prikladnaya Matematika i Mekhanika 60 (1996) 2. [28] Y. Yomei, T. Hiromichi, T. Yasuhiko, I. Yasumi, Tetsu-To-Hagane 89 (2003) 705. [29] E. Gariboldi, W. Nicodemi, G. Silva, M. Vedani, Mater. Sci. Forum 163–166 (1994) 107. [30] Z.P. Zhang, Y.H. Qi, D. Delagnes, G. Bernhart, Heat Treat. Met. 29 (2004) 27. [31] R.C. Thomson, M.K. Miller, Acta Mater. 46 (1998) 2203. [32] R.D. Griffin, J.A. Griffin, G.M. Janowski, C.B. Moss, C.E. Bates, ASM Proc. Heat Treat. (1998) 320–328.
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[33] Y.F. Wang, J.D. Hu, Z.Y. Wang, X.Z. Bu, Heat Treat. Met. 10 (1991) 11. [34] Y.M. Brunzel, Y.B. Gurevich, I.M. Fomin, Metallovedenie i Termicheskaya Obrabotka Metallov 10 (1993) 10. [35] S. Spuzic, C. Subramanian, K.N. Strafford, Mater. Sci. Forum 189–190 (1995) 429.
[36] Q.S. Bai, Y.X. Yao, P. Bex, G. Zhang, Tribology 23 (2003) 81. [37] Y. Zhang, X.Y. Gong, H.M. Zhao, J. Tianjin Inst. Technol. 19 (2003) 69. [38] E. Hornbogen, Metallurgy 34 (1980) 1079.
Heat treatment of multi-element low alloy wear-resistant steel Hanguang Fu a, 1 , Qiang Xiao b, ∗ , Hanfeng Fu c, 2 a
Department of Mechanical Engineering, Tsinghua University, Welding Building, Room 204, Beijing 100084, PR China b Department of Mechanical Engineering, University of Wyoming, P.O. Box 3295, Laramie, WY 82071, USA c First Manufacturing Steel Plant, Shanghai Baoshan Iron and Steel Group Corporation, Shanghai 201900, PR China Received 10 September 2004; received in revised form 8 January 2005; accepted 8 January 2005
Abstract The effect of heat treatment on the performances of multi-element low alloy wear-resistant steel (MLAWS), which is used to make the rolling mill torii liner, was investigated. The results show that the hardness and tensile strength increase as the quenching temperature is increased from 840 to 900 ◦ C. However, the hardness decreases rapidly as the quenching temperature is increased beyond 900 ◦ C, while the temperature has little influence on the tensile strength when it exceeds 900 ◦ C. No clear influence on the impact toughness has been observed unless the quenching temperature is beyond 920 ◦ C. As the tempering temperature exceeds 300 ◦ C, tiny ε carbides separate out from martensite and bainite complex structures and cause the carbon content in the complex structures to decrease. This results in the toughness to increase significantly. The best wear resistance can be obtained by tempering at 350 ◦ C. The optimum heat treatment of MLAWS comprises quenching at 900–920 ◦ C and tempering at 350–370 ◦ C. © 2005 Elsevier B.V. All rights reserved. Keywords: Rolling mill torii; Liner; Heat treatment; Mechanical performances; Wear resistance
1. Introduction In order to protect the stand, the liners have been installed on the torii of roughing mill [1]. The interval between the liner and the bearing chock of Baosteel ϕ 1300 mm roughing mill is 1.3 mm. When the interval becomes too large the liner needs to be replaced. Under rolling conditions, wear of the liner is normally very severe due to the poor service condition of the roughing mill [2]. Therefore, the wear-resistant explosive complex liners have been developed, which consist of Q235 steel plate and multi-element low alloy wearresistant steel (MLAWS) (for details, refer to patent number CN03125740.1) [3,4]. The influence of quenching and tempering temperatures on low alloy steel properties has been widely studied. It was found that the wear resistance of 40SiMnCrMoVRE steel increased as the quenching temperature is increased, and its ∗
Corresponding author. Tel.: +1 307 766 3187; fax: +1 307 766 2695. E-mail addresses: [email protected], [email protected] (Q. Xiao), [email protected] (H. Fu). 1 Tel.: +86 10 62773640; fax: +86 10 62773637. 2 Tel.: +86 21 66142017; fax: +86 21 66142017. 0921-5093/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2005.01.023
maximum value was reached at 950 ◦ C. When the tempering temperature was below 350 ◦ C, the wear resistance decreased slightly as the tempering temperature is increased. As the tempering temperature exceeded 350 ◦ C, the wear resistance increased rapidly, and the maximum wear resistance was reached at 400 ◦ C [5]. The typical heat treatment for medium carbon low alloy Cr–Mo–Ni–V steels involves a twostage austenitizing, followed by tempering and oil quenching [6,7]. Unfortunately, few reports are presently available for high carbon low alloy steel. Therefore, in this paper, we study the influences of quenching temperature, quenching cooling modes and tempering temperature on the hardenability, mechanical performances and wear resistance of high carbon MLAWS. As a result, we obtain an optimal heat treatment process.
2. Experimental MLAWS was melted in an arc furnace. Table 1 shows the composition of the MLAWS. The heat treatment was conducted on a RJX-18-13 resistance furnace. For impact
H. Fu et al. / Materials Science and Engineering A 396 (2005) 206–212
207
Table 1 Chemical composition of sample (wt.%) Element
Content
C Si Mn Cr Ni Mo B Y K Na
0.6–1.0 0.5–1.0 0.8–1.2 0.5–1.5 0.4–0.6 0.4–0.8 0.005–0.015 0.05–0.10 0.05–0.15 0.05–0.15
toughness (αk ) tests, the specimen was of 10 mm × 10 mm × 55 mm without notch. Impact toughness (αk ) tests were performed on a JB30A Charpy impact testing machine. The size of the specimen for fracture toughness (K1c ) test was 10 mm × 10 mm × 55 mm, with a notch of 0.2 mm in depth and 2 mm in width. A 1251-Instron machine was used for the measurement of fracture toughness (K1c ) under force control [8]. The hardness was measured on an HR-150A Rockwell hardness tester. The tensile strength (σ b ), elongation (δ) and section reduction (ψ) tests were performed on a WE-30 universal testing machine [9]. The specimen size was Φ 10 mm × 130 mm. Wear test was carried out at room temperature with a MM200 wear apparatus under the load of 200 N [10]. YG6 hard alloy was used as the anti-wear material, and each test was run for 30 min. The reciprocal of wear amount per unit time represents the wear resistance of materials. The grain sizes and hardenability of austenite were measured through cementation process and end-quench test, respectively. In order to obtain the reasonable heat treatment parameters, the critical points of MLAWS were measured using the expanding method on a Gleeble-1500 thermal stress–strain simulation test machine [11,12].
3. Results and discussion 3.1. Effect of different cooling patterns on the hardenability The critical points of MLAWS are listed in Table 2. Ac1 is the lower critical point, Ac3 is the upper critical point, Ms is the starting temperature of martensitic transformation, and Mf is the finishing temperature of martensitic transformation. The grain structures at different heating temperatures are shown in Fig. 1. The grain sizes of the microstructure are Table 2 Critical points of material (◦ C) Ac1 Ac3 Ms Mf
782 845 294 98
Fig. 1. Grain microstructures at different heating temperature (held for 6 h): (a) 840 ◦ C; (b) 920 ◦ C; (c) 960 ◦ C.
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Fig. 3 shows the effects of quenching temperature on the tensile strength and hardness of MLAWS. The hardness
increases as the quenching temperature is increased from 840 to 900 ◦ C. However, the hardness decreases rapidly as the quenching temperature is increased beyond 900 ◦ C. This is in correlation with the content of carbon and alloy elements which is dissolved in high temperature austenite. At low quenching temperature, the contents of carbon and alloy elements are relatively small. This results in a quenching structure of bainite and martensite, plus some pearlite with low hardness and poor wear resistance [15,16]. When the quenching temperature is above 900 ◦ C, the contents of carbon and alloy elements dissolved in high temperature austenite increase significantly, and the stability of high temperature austenite increases, so the quenching structure contains low hardness retained austenite [5,17]. Therefore, the hardness drops rapidly [18,19]. In addition, it can be seen from Fig. 3 that the effect of quenching temperature on the tensile strengthen of MLAWS is similar to that on the hardness when quenching temperature is below 900 ◦ C, while the quenching temperature has little influence on the tensile strength when it exceeds 900 ◦ C. As discussed above, the carbon and alloy elements dissolved in high temperature austenite increase with increasing quenching temperature. Accordingly, the quenching structure contains more carbon and alloy elements, which strengthen quenching structures and improve the tensile strength [20]. As quenching temperature exceeds 900 ◦ C, the dissolved carbon and alloy elements continue to increase as the temperature increases, and the quenching structure continues to strengthen. At the same time, the retained austenite appears steadily in the quenching structure because the stability of matrix increases. As a result of this comprehensive function, the tensile strength does not change significantly [21]. Effects of quenching temperature on the impact toughness and the fracture toughness of MLAWS are shown in Fig. 4. When the quenching temperature is below 920 ◦ C, it has no significant influence on the impact toughness because the grain sizes do not change much [22,23]. As the quenching temperature increases above 920 ◦ C, the crystal grains grow up consistently, as shown in Fig. 1, and
Fig. 3. Effect of quenching temperature on strength and hardness of MLAWS.
Fig. 4. Effect of quenching temperature on impact toughness and fracture toughness of MLAWS.
Fig. 2. Effect of quenching cooling patterns on hardenability of MLAWS.
in grade 8–9 when the specimens are heated to 920 ◦ C and held for 6 h. This indicates that the grain sizes of austenite are fine and have a very good stability when heated at 920 ◦ C. The influence of cooling patterns on the hardenability of MLAWS is shown in Fig. 2. Since there are many alloy elements in MLAWS that can improve the hardenability of the material, the hardenability of MLAWS is excellent. However, under the condition of water and oil cooling, it is likely to produce crack due to high carbon content, poor heat conductivity and high temperature plasticity [13,14]. Therefore, fog cooling is adopted as the cooling pattern at all stages. 3.2. Effect of quenching temperature on mechanical properties
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Fig. 5. Effect of tempering temperature on strength and hardness.
cause the impact toughness to decrease slightly. Meanwhile, it can be seen from Fig. 4 that the quenching temperature has no clear influence on the fracture toughness of MLAWS. 3.3. Effect of tempering temperature on mechanical properties For tempering temperature influence investigation, MLAWS was quenched at 900 ◦ C and tempered for 2 h at different temperatures. The influence of tempering temperature on the hardness, tensile strength, impact toughness, fracture toughness, elongation and section reduction are shown in Figs. 5–7. Fig. 5 shows that the hardness does not change when the tempering temperature is below 150 ◦ C. Afterwards, the hardness decreases gradually as the temperature is increased. The hardness drops by about 5 HRC after tempering at 450 ◦ C when compared to as-quenched. The hardness tends to drop more rapidly when the tempering temperature
Fig. 6. Effect of tempering temperature on impact toughness and fracture toughness.
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Fig. 7. Effect of tempering temperature on elongation and section reduction.
is above 450 ◦ C. In this case, the structure of MLAWS was transformed into pearlite from martensite because of high temperature tempering. Therefore, the hardness of MLAWS reduced more rapidly [24]. Similar behavior can be observed for the tensile strength, which also decreases with increase in tempering temperature. The slight difference between these results is that the tensile strength curve is smoother than that of hardness curve as the tempering temperature increases above 150 ◦ C. Fig. 6 shows the influence of tempering temperature on the impact toughness and the fracture toughness of MLAWS. Since the phase transformation stress does not fully dispel [25,26], its toughness is relatively low when tempering temperature is below 294 ◦ C. As the tempering temperature exceeds 300 ◦ C, tiny ε carbides separate out from the complex structures of martensite and bainite, and cause the carbon content in the complex structures to decrease, thereby increasing the toughness significantly [27–29]. When tempering temperature rises to 400 ◦ C, the amount of ε carbides reduces slowly, the decomposition of retained austenite occurs, and the brittle cementite separates out [26,30,31]. All these led to the reduction of impact toughness [32]. After that, a large amount of cementite begin to segregate from the complex structures, which decreases the carbon content in martensite and bainite, and results in the toughness of matrix to increase. Although the amount of cementite increases further, the integration effects show that the toughness of MLAWS begins to increase greatly as the tempering temperature increases above 450 ◦ C (Fig. 5). Meanwhile, the overall trend of the fracture toughness change is similar to the impact toughness. It varies with the increase of tempering temperature as shown in Fig. 6. Effects of tempering temperature on the elongation and section reduction of MLAWS are shown in Fig. 7. When tempering temperature is below 400 ◦ C, both the elongation and section reduction are relatively low. As tempering temperature increases above 400 ◦ C, the elongation and the section
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And W c = K1 Wf = K2
Fig. 8. Effect of tempering temperature on wear resistance.
reduction improve by a large margin because of the formation of pearlite in the structure [33,34]. 3.4. Effect of tempering temperature on the wear resistance Fig. 8 shows the influence of tempering temperature on the wear resistance of MLAWS after quenching at 900 ◦ C and tempered for 2 h at different temperatures. When tempering under 300 ◦ C, the effect of temperature on wear resistance is not very clear. However, the wear resistance improves significantly as tempering temperature exceeds 300 ◦ C, and reaches the maximum value at 350 ◦ C. After that, the wear resistance drops consistently as the temperature rise. The relationship between the wear resistance and tempering temperature is related to the hardness and the impact toughness of MLAWS. In the wear of a material, the wear rate consists of two parts [35,36], namely W t = Wc + W f
(1)
P H
(2) P
(εf H)2
(3)
where Wt is the wear rate of materials, Wc is the wear rate caused by cutting, Wf is the wear rate caused by fatigue, P is the pressure, H is the hardness, and εf is the fracture strain at single axis tensile, which reflects the toughness. From the above equations, it can be noted that the wear of a material mainly depends on the hardness and the toughness of the material. When tempering at 350 ◦ C, both the toughness and the hardness of MLAWS are relatively high, so the maximum value of the wear resistance was reached. With the further increase of tempering temperature, the toughness of MLAWS clearly improved, but the hardness drops by a large margin. Therefore, the wear resistance drops as a result of the combination effect. In addition, the special relation between wear resistance and tempering temperature has something to do with the elastic limit of MLAWS. It is well known that carbon steel and alloy steel containing 0.6–0.8% of carbon have supreme elastic limit after quenching and tempering at an intermediate temperature [5,37]. They can be flexibly replied and does not produce plasticity deformation after bearing greater deformation when used for making spring. The carbon content of MLAWS is 0.6–1.0%. There is a supreme elastic limit as well after quenching and tempering at intermediate temperature. The possible reasons for its wear resistance being higher than that of low temperature tempering and high temperature tempering are [38]: (1) Improving the elastic limit can make steel produce greater elasticity deformation at the receiving load and reduce the amount of plasticity deformation, while the elasticity deformation can be replied after the external
Fig. 9. Microstructure of MLAWS after heat treatment.
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force is released, which will not cause wear and improve the wear resistance. (2) From the viewpoint of energy, improving the elastic limit of the steel can improve elastic specific power. With same wear load, there is more energy that can turn into elastic energy, and reduces the energy used for plasticity deformation, crack initiation and propagation. Therefore, it is favorable to the improvement of the wear resistance. Based upon the above analysis, the optimum heat treatment process of MLAWS is concluded as follows: fog quenching at 900–920 ◦ C + tempering at 350–370 ◦ C. Fig. 9 shows the microstructure of MLAWS after heat treatment. It consists of tempering martensite, bainite and retained austenite. This structure has excellent comprehensive mechanical properties and a good wear resistance. The above heat treatment techniques were applied to the wear-resistant compound liner for industry testing. The complex liner was used in Baosteel ϕ 1300 mm roughing mill torii. The result of on-line service investigation indicated that the wear for this liner was only 0.5 mm after 4 months, while the wear for common heat treatment liner was 2.0 mm after 4 months. It appears that the service life of complex liner improves by more than three times.
4. Conclusions This study was carried out to investigate the effect of heat treatment on the performances of multi-element low alloy wear-resistant steel used to make the rolling mill torii liner. This work supports the following conclusions: (1) For quenching temperature less than 900 ◦ C, the hardness and tensile strength increase as the quenching temperature increase. However, the hardness decreases rapidly as the quenching temperature exceeds 900 ◦ C, while the temperature has little influence on the tensile strength when it exceeds 900 ◦ C. No clear influence on the impact toughness has been observed unless the quenching temperature is beyond 920 ◦ C. (2) For the tempering temperature below 150 ◦ C, the hardness does not change. Afterward, the hardness decreases gradually as the temperature increases. The overall trend of the tensile strength changes is similar to the hardness. The impact and fracture toughness are relatively low when tempering temperature is below 300 ◦ C. As the tempering temperature exceeds 300 ◦ C, the toughness increases significantly. The toughness begins to increase greatly as the tempering temperature increases from 450 ◦ C. (3) When quenching at 900–920 ◦ C and tempering at 350–370 ◦ C, the microstructure with tiny martensite, bainite and particle carbides can be obtained. It has excellent comprehensive mechanical properties and a good wear resistance.
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(4) For satisfactory use of rolling mill torii liner, the optimum heat treatment technique of MLAWS comprises fog quenching at 900–920 ◦ C and tempering at 350– 370 ◦ C. (5) The wear amount of the complex liner is only 0.5 mm after using in the rolling mill torri for 4 months. The service life of complex liner improves by more than three times. References [1] R.C. Schrama, Lubr. Eng. 50 (1994) 438. [2] D. Viens, K. Albano, C. Churchill, Proceedings of the 39th Mechanical Working and Steel Processing Conference on Iron and Steel Society of AIME, Indianapolis, IN, USA, 19–22 October, 1997, pp. 467–474. [3] Z.J. Huang, H.G. Fu, patent number CN03125740.1. [4] H.G. Fu, Z.J. Huang, Special Steel 25 (2004) 46. [5] P.Q. Dai, S.X. Huang, Heat Treat. Met. 12 (1998) 19. [6] T. Wada, W.C. Hagel, Metall. Trans. 9A (1978) 691. [7] Z.C. Liu, J.P. Yan, L.P. Zhao, G. Wang, Q. Li, Y.X. Tian, J. Baotou Univ. Iron Steel Technol. 20 (2001) 30. [8] H.G. Fu, D.M. Fu, D.N. Zou, J.D. Xing, J. Wuhan Univ. Technol. Mater. Sci. Ed. 19 (2004) 48. [9] X.H. Cui, O.C. Jiang, S.Q. Wang, Z.M. He, Acta Metall. Sinica (English Letters) 13 (2000) 1053. [10] S.O. Wang, Q.C. Jiang, X.M. Sui, Z.M. He, Tribology 19 (1999) 33. [11] H.Y. Li, N.M. Xiao, L.P. Meng, X.J. Xie, J. Cent. South Univ. Technol. 33 (2002) 495. [12] F. Cao, X.N. Chen, Heat Treat. Met. 28 (2003) 24. [13] Y. Mikita, I. Nakabayashi, Trans. Jpn Soc. Mech. Eng. Part A 54 (1988) 1246. [14] B.J. Yin, X.H. Li, Hot Working Technol. 6 (2003) 46. [15] R. Yang, L.J. Li, Y.C. Li, Iron Steel 34 (1999) 41. [16] E. Anelli, M.C. Cesile, P.E. Di Nunzio, Mater. Sci. Technol. Meeting (2003) 453–466. [17] S.F. De, R.C. Rodriguez, C.A. Ratnapuli, Bottrelcoutinho, Metalurgia 40 (1984) 195. [18] Y. Tomita, K. Morioka, Mater. Char. 38 (1997) 243. [19] D.Y. Wei, J.L. Gu, H.S. Fang, B.Z. Bai, Z.G. Yang, Int. J. Fatigue 26 (2004) 437. [20] B.B. Vinokur, A. Vinokur, V.E. Shtessel, Metall. Mater. Trans. A 27A (1996) 2852. [21] J. Zeman, S. Rolc, J. Buchar, J. Pokluda, ASTM Special Tech. Publication 1074 (1990) 396. [22] S.I. Yokoyama, K. Kubota, H. Sasaki, Y. Minamino, J. Iron Steel Inst. Jpn. 90 (2004) 263. [23] C.L. Davis, J.E. King, Mater. Sci. Technol. 9 (1993) 8. [24] M. Gojic, L. Kosec, P. Matkovic, J. Mater. Sci. 33 (1998) 395. [25] C. Aubry, S. Denis, P. Archambault, A. Simon, F. Ruckstuhl, B. Miege, Metall. Italiana 91 (1999) 33. [26] H.G. Fu, J.Z. Wu, J. Xu, Res. Iron Steel 4 (1994) 17. [27] G.D. Pigrova, Prikladnaya Matematika i Mekhanika 60 (1996) 2. [28] Y. Yomei, T. Hiromichi, T. Yasuhiko, I. Yasumi, Tetsu-To-Hagane 89 (2003) 705. [29] E. Gariboldi, W. Nicodemi, G. Silva, M. Vedani, Mater. Sci. Forum 163–166 (1994) 107. [30] Z.P. Zhang, Y.H. Qi, D. Delagnes, G. Bernhart, Heat Treat. Met. 29 (2004) 27. [31] R.C. Thomson, M.K. Miller, Acta Mater. 46 (1998) 2203. [32] R.D. Griffin, J.A. Griffin, G.M. Janowski, C.B. Moss, C.E. Bates, ASM Proc. Heat Treat. (1998) 320–328.
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