Effects of Induction Hardening and Prior Cold Work on a Microalloyed Medium Carbon Steel

17th ASM Heat Treating Society Conference Proceedings Including the 1st International Induction Heat Treating Symposium

J.L. Cunningham American Axle and Manufacturing, Rochester Hills, Michigan

D.J. Medlin, G. Krauss Colorado School of Mines, Golden, Colorado

ABSTRACT of heat treatment processing (2). Material costs are reduced due to the lowering of alloy additions and processing costs The torsional strength and microstructural response to are reduced due to the elimination of heat treating steps. induction hardening ofa 10V45 steel with prior cold work was evaluated. The vanadium-microalloyed 1045 (10V45) steel was characterized in three conditions: as-hot-rolled, 18% cold-reduced, and 29% cold-reduced. Two of these evaluations, 10V45 as-hot-rolled and 10V45-18%, were subjected to stationary and progressive induction hardening to three nominal case depths: 2mm, 4mm, and 6mm. All specimens were subsequently furnace tempered at 190 deg. C for 1 hour. The martensitic case microstructures contained residual lamellar carbides due to incomplete dissolution of the pearlitic carbides in the prior microstructure. Torsional overload strength, as measured by maximum torque capacity, is greatly increased by increasing case depth, and to a lesser extent by increasing prior cold work level. Maximum torque capacity ranges from 2520 N- m to 3170 N-m, depending upon induction hardening processing. Changing induction hardening processing from stationary (single-shot) to progressive (scan) had little effect on torque capacity.

MONOTONIC TORSIONAL STRENGTH of induction hardened shafts is greatly influenced by the depth and hardness of the case, as well as the hardness of the core, as was shown by Ochi and Koyasu (I). Figure 1 shows that as case depth and carbon content increase, torsional strength increases. Likewise, as hardness increases, whetller in the core or in the case, torsional strength increases, as shown in Figure 2. This effect is also shown for varying case depths. Cost reduction measures in manufacturing industries, particularly within the automotive industry, has led to reduced material consumption and elimination of processing steps where possible. With the addition of small Case hardening depthi radius, ti r amounts ofvanadium, niobium, and/or titanium, microalloyed steels provide high strength with a minimum Figure 1. Torsional strength as a function of case depth for various grades of steel, from Ochi and Koyasu ( I ).

575 concerns regarding phase transformation: dissolution and

- precipitation of second-phaseparticles, and austenitization 10 0- mechanism at fast heating times. Both issues are well ~ understood from an equilibrium perspective, but less for non-equilibrium conditions. Since induction hardening is .!: +' not a slow heating or slow cooling process, the solubility of t)D C not only the microalloy carbonitrides, but also the pearlitic (1) Lo +' carbides, is greatly dependent on heating time, not just VI temperature. Of major concern is the complete dissolution ~ C of the microalloyingelements and other carbonitride O "iii formers during induction heating. Lo O The purpose of this paper is to examine benefits 1- associated with the use of induction hardened microalloyed (a) steel, with and without cold work. Some information regarding carbide retention during induction hardening is also presented. Case hardness (HV) Experimental Procedure .,!U I I I 0- Materials. A commercially produced 10V45 (SAE 2000 ~ 1045, modified with vanadium) was provided by Inland '--" Steel in 33 mm round bars. The chemical composition of ..c: +' the steel is given in Table I. The 10V45 was cold-drawn to bD 1800 c two levels of cold work, 18% and 29%, resulting in final Q) tlr=7 diameters of29.9 mm and 27.8 mm, respectively. These L +' three materials are designated 10V45-0, 10V45-18, and U) t/r=O.35-0.4 1600 10V45-29 for the as-hot-rolled, 18% cold-drawn, and 29% !U I I I C cold-drawn bars, respectively.

.Q 300 400 U) L O Core hardness ( HV) (b) ~

Figure 2. Torsional strength as a function of (a) case hardness and (b ) core hardness for various case depths, where t is case depth and r is bar radius. Taken from Ochi and Kayasu (I).

Microalloyed, medium carbon steels gain their strength from the precipitation of carbonitrides during cooling after austenitization. Strengths typical of highly tempered Hardness Profiles. Hardness of all material conditions martensitic steels are obtained by pearlite formation and were measured on both the Vickers scale and Rockwell C precipitation strengthening. Toughness is low, but may be scale. At least three sections of each material, increased by grain size refinement or by limiting alloy carbon approximately 13 mm in height, were ground to a 6 ~ content (2-5). Since the strengthening effects are exhibited in diamond finish. Vickers microhardness profiles along the the direct -cooled, ferritic-pearlitic microstructures, heat radii of the sections were measured on a LECO M-400A treatment subsequent to rolling or forging is not necessary Hardness Tester with a 1000 9 load and a 55x objective lens. (2). Point measurements of Rockwell C hardness were measured High core strength provides a means of increasing on a LECO R-600 Hardness Tester with a 1 kg load. performance of induction hardened shafts and microalloying Vickers hardness data were converted to Rockwell C scale is one method used to increase the core strength. Also, since using standardized conversion tables (7). strength increases directly with amount of cold work (6), it may be possible to use cold-reduction concurrently witll the Metallography. Specimens of all material conditions microalloying to provide a core with high strength. were polished to a I JUll diamond finish and observed on an In the process of induction hardening there are two

576 Ausjena Neophot 211ight metallograph. Both longitudinal Table 11-Full-factorial experimental test matrix for the and transverse oriented specimens were observed. Ferrite indllction hardened conditions used in this study. volume fraction, ferrite grain size, and decarburization depth Run Material Method Case Depth, were measured by standard techniques (8-10). All specimens mm were etched in a 2% nital solution. lOV45-0 scan 2.0 Tensile Testing. Tensile tests were conducted on an Instron 1125 screw-type mechanical testing machine, with 2 IOV45-0 scan 4.0 analog-to-digital data acquisition. Round tensile specimens were machined from the center of the as-hot-rolled and cold- 3 lOV45-0 scan 6.0 drawn bars. Gage length was 25.4 mm (1.00 in.). Tests were 4 lOV45-O single shot 2.0 run in triplicate and at room temperature with a crosshead speed of 1.27 mm/min. (0.05 in./min.). 5 lOV45-o single shot 4.0

Electron Microscopy. Electron microscopy was 6 lOV45-0 single shot 6.0 conducted on a JEOL JX-840 scanning electron microscope. 7 lOV45-l8 scan 2.0 Metallographic specimens were viewed with a 1 ~ diamond polish and a 15:6:1 H2O : dodecylbenzesulfonic acid, sodium 8 lOV45-l8 scan 4.0 salt: picric acid etching solution, heated to approximately 80 deg. C. 9 lOV45-18 scan 6.0

Machining of Test Shafts. One hundred test shafts were 10 lOV45-18 single shot 2.0 made from each of the three initial material conditions: 11 IOV45-18 single shot 4.0 10V45-0, 10V45-18, and 10V45-29. The shaft is approximately 330 mm (13 in.) in length and 25 mm (1 in.) 12 lOV45-18 single shot 6.0 in diameter. All case depths were measured at the smallest diameter on the shaft (henceforth referred to as the reference region). Table III -Equipment setup for the six induction hardening conditions used in this study. Induction Hardening. Two material conditions were selected, 10V45-0 and 10V45-18, to undergo induction Scan Case Dwell Power, kW Freq,

kHz hardening heat treatment. Induction hardening of the test Rate Depth, Time, shafts was conducted with both stationary (single-shot) and mm sec progressive (scan) methods. An experimental test matrix was defined which tested the 2 4.0 200 (50%) 8.0-9.0 effects of material, hardening method, and case depth on Single 4 4.8 200 (50%) 8.0-9.0 torsional overload strength. The test matrix is shown in Shot Table II. The matrix is full-factorial, testing the two material 6 4.8 224 (56%) 7.5-8.6 conditions, the two hardening methods, and three nominal case depths of 2 mm, 4 mm, and 6 mm. Torsional overload 2 84 (21%) 8.2 strength is the response variable. 0.8 The equipment setup for the differing hardened in/sec 4 0.1 110 (27.5%) 9.0 conditions is given in Table III. All conditions were run on a 6 136 (34%) 5.5 400 kW, 3110 kHz invertor. It was necessaryto adjust tlle invertor for the 6 mm casesby adding capacitance. Peak temperatures were not measured due to equipment limitations. All shafts were quenched to 24 deg. C (76 deg. shaft (undercut). All induction hardened shafts were F) with a 2% concentration ofHoughton Aqua Quench 251 subjected to the same furnace temper treatment: 190 deg. C polymer quenching media. Quenchant flow rate was 7.6 Lis (375 deg. F) for I hour (II). @ 276 MPa and 5.1 Lis @ 207 MPa (120 gpm @ 40 psi and 80 gpm @ 30 psi) for the single-shot and scanned shafts, Torsion Testing. Torsional overload strength was respectively. To confirm correct setup, case depths were chosen as the response variable to the experimental verified with a Brinell glass at the reference region of the induction hardening matrix, given in Table II. Two shafts of each hardened condition (24 total) were tested. Twist rate

577

Figure 7. SEM image showing ghost pearlite in the Figure 8. SEM image showing ghost pearlite in the induction hardened 2mm case microstructure. induction hardened 6mm case microstructure.

Examples of the case microstructures are shown in the SEM micrographs of Figures 7 and 8. These micrographs not only reveal prior-austenitic grain size, but also the presence of lamellar carbides, revealed by a picric acid etch, in what are assumed, from hardness measurements, to be fully martensitic microstructures. These carbides closely resemble the prior- pearlitic microstructure, and have been termed "ghost pearlite" by Medlin et a/. (12), according to industry terminology. Changes in microstructure due to coil geometry are less obvious, but changes in structure due to changes in case depth are readily observed. The ghost pearlite, present in the 2 mm cases, gives way to a more martensitic structure, which becomes nearly fully martensitic in the 6 mm cases. Likewise, prior-austenite grain size increases as case depth increases, from approximately 10 ~ in the 2 mm case to 0 5 O 0.5 approximately 30 ~ in the 6 mm case. Temperatures achieved by induction hardening were not measured, but Time (s) reasonable measurements of time were available. From this Figure 9. Estimated time-temperature diagram showing the information time-temperature cycles to produce the various differences between the six induction hardened conditions. case depths by single shot and scan induction hardening were estimated and are shown schematically in Figure 9. At deeper case depths, the effect of cold-work becomes even less significant, since the mechanical effects of the Torsional Strength. Figure 10 summarizes the effect of deeper case begin to override the effect of prior work case depth, cold work and coil geometry on maximum torque hardening. capacity .Increased case depth has the expected effects of increasing both the yielding torque and the maximum torque Residual Stress. Figure 11 shows the residual stress of the shaft (1). A similar trend is seen with increased cold- profiles for the single-shot and scan induction hardening reduction. As cold work increases, maximum torque of the processing. The residual stress measurements were shaft increases, but to a smaller extent compared to the effect determined by standard x-ray diffraction techniques. All the of case depth. profiles exhibit similar characteristics in which the Coil geometry is shown to have little effect on maximum maximum residual compressive stress is between 0.05-0.10 torque capacity .When the induction hardening process is rnrn below the surface. The 10V45-18 specimen that was changed from single-shot to scan, very little, if any, difference induction hardened by the scan method had the highest in maximum torque is noticed. This shows that the two very residual compressive stress of almost 780 MPa. The other different hardening processescan give similar results in specimens had maximum residual compressive stresses torsional behavior. ranging from 620 to 680 MPa.

580

coil geometry .The major effect of case depth was expected from previous literature (I). Discussion is now limited to the slight effects of cold work, coil geometry and microstructure. The effect of coil geometry, or relatively low sensitivity of the induction hardened microstructure to coil geometry , cannot be construed to say that the single-shot hardening process of this study is equivalent to the scan hardening process of this study. All that can be said is that these two processeswere well matched to produce shafts with equivalent strength properties and case depths. Cold reduction has the expected effect of increasing torque capacity of induction hardened shafts. However, this effect is minor. Cold work does produce shafts which are consistently stronger than currently produced shafts, but that effect may not be separable from the effect of microalloying, which can also account for the increased performance with respect to as-produced shafts.

Conclusions

Mechanical and metallurgical properties of the IOV45 steel have been determined in both the as-received conditions and the induction hardened conditions. As- received conditions include as-hot-rolled, 18% cold-drawn, and 29% cold-drawn bars. Induction hardened conditions include the as-hot-rolled and 18% cold-drawn steels, with 2 mm, 4 mm, and 6 mm case depths, hardened by single-shot and scan induction coils. Evaluation of the properties which resulted from the induction hardening processes has led to Figure 12. Schematic representation of the austenization the following conclusions: mechanism proposed by Speich and Szirmae for two different I. Prior cold work, at the levels studied, has a slight times, ti and ~, and for three different austenite orientations, effect on hardened case microstructure, maximum case Al. A2, and A3, showing interlamellar transformation (23). hardness, or maximum torque capacity . 2. Prior cold work has the effect of producing higher The scan induction hardening cycles in this study case hardness, when compared to as-hot-rolled produced better dissolution of the pearlite than did the single- microstructures hardened with the same induction shot cycles, for the same case depth. Again, this is assumed hardening cycle. to be due to the increased temperature required to drive an 3. For the six given induction hardening setups, full equivalent case with an order-of-magnitude shorter dwell dissolution of the pearlitic prior-microstructure does not time. This is not to say that all scan hardening is better than occur, resulting in a case microstructure of high hardness all single-shot hardening, since the two processesare not which is not fully martensitic. Arrays of lamellar carbides, directly comparable. The magnetic physics involved (13-16) described as "ghost pearlite," are present in the hardened are quite different. However, a comparison of the two case microstructure. processes,as given in this study, dramatically shows the 4. Case microstructure, as a result of induction temperature dependence of the solution kinetics referred to hardening, can be assumed to be primarily a function of earlier (12,17,18-21). austenitizing temperature at the very short heating times of induction hardening. Torsional Strength Response. The torsional strength of 5. Changing of induction coil geometry from single- the induction hardened shafts, represented by maximum shot to scan, while not directly comparable processes, torque capacity , has been shown to be primarily a function of illustrated the strong temperature dependence of the case depth, with lesser effects by cold work and changes in mechanism of austenitization. However, changing of coil geometry does not significantly effect maximum torque

582 capacity. 9 ASTM Standard E112-84, ASTM Annual Book of 6. Maximum torque capacity increases significantly with Standards, Vol. 3.03, 1984, 120-152. increased case depth. 10 ASTM Standard E1077-91, ASTM Annual Book of Acknowledgments Standards, Vol. 3.01,1995,735-744.

This research program was perfonned as the M.S. thesis 11 Luckett, D. D. , personal communication, Induction of one of the authors (Justin L. Cunningham) within the Bar Services, Inc., Warren, Michigan, 1996. and Forging Steel Research Program of the Advanced Steel Processing and Products Research Center (ASPPRC), a NSF 12 Medlin, D. I., G. Krauss, and S. W. Thompson. ls1 Industrial/University Cooperative Research Center, at the International Conference on Induction Hardeninl! of Colorado School of Mines. The authors gratefully Gears and Critical COmDOnents,ASM, Indianapolis, acknowledge the support of many individuals from the 1995, 57-65. following corporations for helpful discussion, provision of steels, cold work processing, induction hardening, and 13 Giancoli, D. Phvsics for Scientists and Engineers, testing: General Motors, Chrysler Corporation, Ford Motor Second Edition, Prentice Hall, Englewood Cliffs, NJ, Company, Inland Steel, Republic Engineered Steels, North 1988,628-631. Star steel, Nelson Steel, Michigan Bar, Induction Services, Inc., The Timken Company, and the Colorado School of 14 Leathennan, A. F. and D. E. Stutz. Metal Treat., 21 Mines. (2), 1970,3-6,8-12.

References 15 Stevens, N. Materials Handbook, 9th Edition, vol.4, ASM, Materials Park, OH, 451-483. Ochi, To and Yo Koyasuo "Strengthening of Surface Induction Hardened Parts for Automotive Shafts Subject 16 Osbom, H. B., Jr. "Induction Heating and Hardening," to Torsional Load," SAE Technical Paper Series, Parts A & B, ASM Home Study and Extension Courses, #940786, 1994. #C6L11A-B,1977.

Grassl, K., S. W. Thompson, and G. Krauss. "New 17 Yakov1eva, I. L., V. M. Schast1ivtsev, T. I. Options for Steel Selection for Automotive Applications," Tabatchikova, D. A. Mirzaev, and A. L. Osintseva. ~ SAE Technical Paper Series, #890508, 1989. Phxsics of Metals and Meta110{!raohv,vo1. 79, 1995, 576-580. 3 Landgraf, R. M. and A. M. Sherman. Micro-Allo,ying '75: Proceedings of the SvrnDosiurn on HSLA Steels, 1-3 18 Speich, G. R. and A. Szinnae. Transactions of the Oct., 1975, 498-502. Metallurgical SocietY of AIME, vol. 245, 1969, 1063- 1074. 4 Cline, R. S, and J. McClain. Fundamentals of Microalloying Forging Steels, ed. by G. Krauss and S. K. 19 Speich, G. R., V. A. Demarest, and R. L. Miller. Banelji, TMS/AIME, 1987,339-349. Metallurgical Transactions A, vol. 12, 1981, 1419- 1428. 5 Held, J. F. Fundamentals of Microalloxin!! For!!in!! ~, ed. by G. Krauss and SoK. Banerji, TMS/ AIME, 20 Hillert, M., K. Nilsson, and L. -E. Torndahl. Journal of 1987,175-188. the Iron and Steel Institute, vol. 209, 1971,49-66.

6 Mischke, C. R. Standard Handbook of Machine Design, 21 Schastlivtsev, V. M., I. L. Yakovleva, T. I. ed. by I. E. Shigley and C. R. Mischke, McGraw-Hill, Tabatchikova, and D. A. Mirzaev. Journal de Physigue 1986,8.1-8.33. IY, Colloque C8, supplement au Journal de Physigue ill, vol. 5, December 1995,531-536. '7 ASTM Standard EI40-88, ASTM Annual Book of Standards, Vol. 3.01, 1991. 22 J.L. Cunningham, "Effects of Induction Hardening and Prior Cold Work on a Microalloyed Medium Carbon 8 ASTM Standard E562-83, ASTM Annual Book of Steel", M.S. Thesis, MT-SRC-O96-024, Colorado Standards, Vol. 3.03, 1984,518-524. School of Mines, Golden, Colorado, August 1996.

583 23 Speich, G.R. and A. Szirmae, "Fonnation of austenite from Ferrite and Ferrite-Carbide Aggregates", Transactions of the Metallurgical SocietY of AIME, vol. 245, 1969, p. 1063-1074.

584