Handbook Heat Treating (1991) WW Part 2 docx

Fig. 11 Comparison of strength levels achievable by means of continuous and batch annealing of solution- strengthened and HSLA steels. Source: Ref 21 The HSLA steels are sensitive to the hot-mill coiling temperature, and a low-temperature coiling practice is preferred to maximize precipitation strengthening. The thermal profile on a continuous-annealing line for microalloyed HSLA steels is similar to that for the solution-strengthened steels. However, the HSLA steels require higher annealing temperatures in order to ensure complete recrystallization because the carbonitride particles of niobium, titanium, and vanadium retard recrystallization (Fig. 12). Fig. 12 Variation of recrystallization-finish temperature with alloy content in solution-strengthened and HSLA steels. Source: Ref 22 Yield strength levels ranging from about 280 to 550 MPa (40 to 80 ksi) are possible and practical with solution- strengthened and microalloyed HSLA steels. Yield ratio, that is, the ratio of yield strength to tensile strength, is about 0.8. Like the plain carbon steels, these steels exhibit bake-hardening characteristics. Recovery-annealed steels, also known as stress-relief annealed steels, can be produced by a low-temperature annealing treatment (Fig. 13) of about 565 °C (1050 °F). High strength levels are achievable by preventing recrystallization (substructure strengthening), while some improvement in ductility can be realized because of the recovery of the cold- worked structure. The substructure strengthening is of the order of 340 MPa (50 ksi), and further increases in strength can be achieved with additions of phosphorus or silicon and also niobium, in a manner similar to that used with the fully recrystallized steels (Ref 24). Yield strengths of commercially available steels range from 600 to 800 MPa (90 to 120 ksi), with a yield ratio of about 0.95. Fig. 13 Tensile properties as a function of anneal (soak) temperature for a 0.06% P, 0.50% Si steel; 70% cold reduction. Source: Ref 23 Dual-Phase Steels. Annealing of dual-phase steels involves soaking in the intercritical or two-phase (ferrite-plus-austenite) region, followed by the transformation of some of the austenite into martensite. The martensite is responsible for the higher strength levels, especially tensile strength, of these steels (Ref 25, 26). To promote the austenite-to-martensite transformation, a critical level of hardenability is needed, depending on the cooling rate. Lower hardenability, from a reduced amount of manganese (and/or molybdenum, chromium) in the steel, can be tolerated with a higher cooling rate (Fig. 14). As described below, there are several types of dual-phase steels, determined by the thermal profile on the annealing line following the intercritical anneal. Fig. 14 Effect of cooling rate from the intercritical temperature on the manganese required to form dual-phase microstructures. Source: Ref 1 Quenching from the Intercritical Temperature. The most economical dual-phase steels can be produced by water quenching low-hardenability (0.3 to <1% Mn) steels directly from the intercritical-annealing temperature (>740 °C, or 1365 °F). The overaging treatment tempers the martensite phase and lowers the solute carbon content in the ferrite. Overaging is conducted at 400 °C (750 °F) if ductility is to be maximized or at 260 °C (500 °F) if bake hardenability is to be maximized (Ref 25, 26, 27). The overaging treatment can result in the return of the yield-point elongation, necessitating subsequent temper rolling. Different tensile strength levels (400 to 1200 MPa, or 60 to 170 ksi) are realized by altering the volume fraction of the martensite phase through changes in the carbon content, manganese content, and quenching temperature (Ref 25, 26). Dual-phase steels of this type are characterized by a yield ratio of about 0.7 and, while the r - value is generally low, it can be improved somewhat by resorting to high hot-mill coiling temperatures, in conjunction with high soak temperatures on the annealing line (Ref 3). Dual-Phase Steels with Increased Hardenability. Dual-phase steels (tensile strengths of 400 to 1000 MPa, or 60 to 140 ksi) having a low yield ratio (about 0.5), along with a superior tensile strength-ductility combination and strain-hardening behavior (high n-value) can be produced by lowering the quenching temperature to about 450 °C (840 °F) (Fig. 15). Because lowering the quench temperature requires gas-jet cooling (~20 °C/s, or 35 °F/s) between the soak and quench stages, the hardenability must be increased by the addition of manganese (Fig. 16). Although the manganese level is generally about 1.6%, partial substitution with molybdenum (for example, 1.3% Mn plus 0.3% Mo) offers some advantages (Ref 29). These types of dual-phase steels can be processed even on a gas-jet cooling line, with similar alloying levels. Overaging is generally restricted to less than about 150 °C (300 °F) and, because the steels are continuous yielding in the as-annealed condition, no temper rolling is required. Fig. 15 Mechanical properties of a 1.5% Mn steel versus quenching temperature. RT, room temperature. Source: Ref 28 Fig. 16 Relation between quench temperature and the manganese content required to obtain a dual- phase microstructure. A, overaging (tempering) required; B, no tempering required. Source: Ref 3 Fully Martensitic Steel. A related category of steels comprises the fully martensitic steels produced by annealing and water quenching from above the A 3 critical temperature. Ultrahigh tensile strength levels, ranging from 900 to 1500 MPa (130 to 210 ksi), are realized using relatively lean compositions (0.08 to 0.25% C, 0.45% Mn). To ensure a martensitic structure, some boron (10 ppm) is generally added. These steels have very limited ductility. Bendability can be improved by low-temperature (<260 °C, or 500 °F) overaging. Both the yield and tensile strengths of martensitic steels are primarily determined by the carbon content. The 0.2% yield strength of low-carbon martensite increases with increasing carbon, as shown by (Ref 30): σ 0.2 (MPa) = 413 + 17.2 × 10 5 % wtC (Eq 1) Advantages of Continuous Annealing. Paint performance, particularly as it relates to adhesion and corrosion resistance, is dependent on sheet surface cleanliness. The presence of in-line electrolytic cleaning/scrubbing in modern annealing lines, followed by annealing at high temperatures, provides a cleaner surface than is attained for batch-annealed sheet, which is not usually cleaned electrolytically. Continuous-annealed and batch-annealed sheets are compared with regard to surface carbon contamination in Fig. 17. Fig. 17 Comparison of surface carbon contamination on batch- and continuous-annealed sheets. Source: Ref 31 Because a single strand of sheet is annealed on the continuous lines, the control of temperature is better than for batch annealing, and more uniform properties are obtained along the coil length. This is especially true of high-strength steel grades in which the strengthening components precipitation (HSLA steels), substructure (recovery-annealed steels), and martensite (dual-phase steels) are dependent on the annealing temperature. For example, in the case of a 420 MPa (60 ksi) yield strength HSLA grade, standard deviations of yield strength are reported to be 10 MPa (1.5 ksi) with continuous annealing, compared to 22 MPa (3 ksi) with batch annealing (Ref 23). For a 700 MPa (100 ksi) yield strength recovery- annealed grade, the standard deviations are 12 MPa (1.7 ksi) with continuous annealing and 31 MPa (4.5 ksi) with batch annealing. Steels for Tinplate Applications The conventional tinplate continuous-annealing lines involve soaking at 650 to 700 °C (1200 to 1300 °F), followed by slow gas-jet cooling (~10 °C/s, or 20 °F/s) to the ambient. The T4 (Rockwell Hardness, HR 30T = 61 ± 3) and T5 (HR 30T = 65 ± 3) tempers are being produced on these lines using plain carbon aluminum-killed chemistries. The production of T2 (HR 30T = 53 ± 3) and T3 (HR 30T = 57 ± 3) tempers by continuous annealing have necessitated several chemistry restrictions and process modifications (Ref 32, 33, 34, 35). The optimum carbon level is 0.02 to 0.07%, with total nitrogen restricted to less than 0.003%. Hotmill coiling is restricted to below 630 °C (1165 °F) to prevent deterioration of corrosion resistance due to the presence of coarse carbides (Ref 32). The effects of pertinent continuous- annealing parameters on black-plate hardness (0.035% C, 0.003% N steel) are shown in Fig. 18. Rapid cooling (40 to 70 °C/s, or 70 to 125 °F/s) from 700 °C (1300 °F), followed by overaging at 400 to 450 °C (750 to 840 °F) for 60 s, is necessary to reduce the carbon concentration solute and, consequently, the hardness. The rapid cooling is achieved by means of high-speed gas-jet cooling systems (Ref 9, 34). Fig. 18 (a) Effect of annealing-cycle parameters on black- plate hardness of a 0.035% C steel. (b) Heat cycle parameters and optimum conditions determined from plots shown in (a). Source: Ref 33 Improved hardness distribution in the continuous-annealed T3 product is shown in Fig. 19 (Ref 33). Other advantages over batch-annealed product include improved corrosion resistance as a result of enhanced surface cleanliness and the prevention of surface defects caused by the surface enrichment of carbon and manganese. Fig. 19 Comparison of the hardness distribution of T3 tinplate produced by (a) batch annealing and (b) continuous annealing. N , number of specimens; X, mean; S, standard deviation. Source: Ref 33 Steels for Enameling Applications Specific requirements for steels to be used in porcelain-enameled appliance parts, other than good formability, surface cleanliness, and flatness; include freedom from carbon boil and absence of fish scaling (Ref 1, 36, 37). Carbon boiling is associated with the presence of coarse carbides near the sheet surface that react with the enamel frit and produce carbon monoxide/dioxide bubbles and pits in the enamel surface. The formation of such carbides is unlikely with continuous annealing. While a low carbon level of 0.015 to 0.02% is desirable from a formability stand-point (Fig. 7), a higher carbon level is preferred in enameling steels to provide resistance to fish scaling, which is the expulsion of enamel fragments caused by the diffusion of hydrogen from the steel. Alternate means of improving the fish scaling resistance include the introduction of BN particles by the addition of boron (30 to 60 ppm) or of TiN/TiS particles by the use of an IF-type chemistry. The second alternative also provides excellent formability. References cited in this section 1. P.R. Mould, in Metallurgy of Continuous-Annealed Sheet Steel, B.L. Bramfitt and P.L. Mangonon, Ed., TMS-AIME, 1982, p 3-33 2. T. Obara et al., Kawasaki Steel Tech. Rep., No. 12, July 1985, p 25-35 3. K. Matsudo et al., in Technology of Continuously Annealed Cold-Rolled Sheet Steel, R. Pradhan, Ed., TMS- AIME, 1985, p 1-36 4. K. Matsudo et al., Nippon Kokan Tech. Rep. (Overseas), No. 38, 1983, p 10-20 9. F. Yanagishima et al., Iron Steel Eng., May 1983, p 36-44 13. W.B. Hutchinson, Int. Met. Rev., Vol 29, 1984, p 25-42 14. N. Takahashi et al., in Metallurgy of Continuous-Annealed Sheet Steel, B.L. Bramfitt and P.L. Mangonon, Ed., TMS-AIME, 1982, p 133-153 15. R. Pradhan and J.J. Battisti, in Hot-and Cold-Rolled Sheet Steels, R. Pradhan and G. Ludkovsky, Ed., TMS- AIME, 1988, p 41-56 16. Y. Tokunaga and H. Kato, in Metallurgy of Vacuum-Degassed Steel Products, R. Pradhan, Ed., The Minerals, Metals & Materials Society, 1990, p 91-108 17. K. Osawa et al., in Metallurgy of Vacuum-Degassed Steel Products, R. Pradhan, Ed., The Minerals, Metals & Materials Society, 1990, p 181-195 18. K. Yamazaki et al., in Microalloyed HSLA Steels, ASM International, 1988, p 327-336 19. K. Nakaoka et al., in Formable HSLA and Dual-Phase Steels, A.T. Davenport, Ed., TMS- AIME, 1977, p 126-141 20. M. Kurosawa et al., Kawasaki Steel Tech. Rep., No. 18, 1988, p 61-65 21. R. Pradhan, J. Heat Treat., Vol 2 (No. 1), 1981, p 73-82 22. R. Pradhan, in Metallurgy of Continuous-Annealed Sheet Steel, B.L. Bramfitt and P.L. Mangonon, Ed., TMS-AIME, 1982, p 203-227 23. R. Pradhan et al., Iron Steelmaker, Feb 1987, p 25-30 24. R. Pradhan, in HSLA Steels: Technology and Applications, American Society for Metals, 1984, p 193-201 25. R. Pradhan, in Technology of Continuously Annealed Cold-Rolled Sheet Steel, R. Pradhan, Ed., TMS- AIME, 1985, p 297-317 26. I. Gupta and P H. Chang, in Technology of Continuously Annealed Cold-Rolled Sheet Steel, R. Pradhan, Ed., TMS-AIME, 1985, p 263-276 27. K. Nakaoka et al., in Structure and Properties of Dual-Phase Steels, R.A. Kot and J.W. Morris, Ed., TMS- AIME, 1979, p 330-345 28. K. Matsudo et al., Nippon Kokan Tech. Rep. (Overseas), No. 29, Sept 1980, p 1-9 29. R. Pradhan and J.J. Battisti, Paper presented at the Second NKK- CAL Family Meeting, Düsseldorf, West Germany, 1989 30. G.R. Speich and H. Warlimont, J. Iron Steel Inst., Vol 206, 1968, p 385-392 31. P. Paulus et al., Paper presented at ATS Steelmaking Conference, Paris, Dec 1986 32. T. Obara et al., in Technology of Continuously Annealed Cold-Rolled Sheet Steel, R. Pradhan, Ed., TMS- AIME, 1985, p 363-383 33. H. Kuguminato et al., Kawasaki Steel Tech. Rep., No. 7, March 1983, p 34-43 34. Recent Development in CAPL Technology: CAPL for Tinplate, Nippon Steel Tech. Rep., Oct 1988 35. T. Asamura et al., Nippon Steel Tech. Rep., No. 29, April 1986, p 45-52 36. A. Yasuda et al., in Hot- and Cold-Rolled Sheet Steels, R. Pradhan and G. Ludkovsky, Ed., The Minerals, Metals & Materials Society, 1988, p 273-285 37. S.T. Furr and O. Ehrsam, in Hot- and Cold-Rolled Sheet Steels, R. Pradhan and G. Ludkovsky, Ed., The Minerals, Metals & Materials Society, 1988, p 287-312 Quenching of Steel Charles E. Bates, Southern Research Institute; George E. Totten, Union Carbide Chemicals and Plastics Company Inc.; Robert L. Brennan, E.F. Houghton & Company Introduction QUENCHING refers to the process of rapidly cooling metal parts from the austenitizing or solution treating temperature, typically from within the range of 815 to 870 °C (1500 to 1600 °F) for steel. Stainless and high-alloy steels may be quenched to minimize the presence of grain boundary carbides or to improve the ferrite distribution but most steels including carbon, low-alloy, and tool steels, are quenched to produce controlled amounts of martensite in the microstructure. Successful hardening usually means achieving the required microstructure, hardness, strength, or toughness while minimizing residual stress, distortion, and the possibility of cracking. The selection of a quenchant medium depends on the hardenability of the particular alloy, the section thickness and shape involved, and the cooling rates needed to achieve the desired microstructure. The most common quenchant media are either liquids or gases. The liquid quenchants commonly used include: • Oil that may contain a variety of additives • Water • Aqueous polymer solutions • Water that may contain salt or caustic additives The most common gaseous quenchants are inert gases including helium, argon, and nitrogen. These quenchants are sometimes used after austenitizing in a vacuum. The ability of a quenchant to harden steel depends on the cooling characteristics of the quenching medium. Quenching effectiveness is dependent on the steel composition, type of quenchant, or the quenchant use conditions. The design of the quenching system and the thoroughness with which the system is maintained also contribute to the success of the process. Fundamentals of Quenching and Quenchant Evaluation Fundamentally, the objective of the quenching process is to cool steel from the austenitizing temperature sufficiently quickly to form the desired microstructural phases, sometimes bainite but more often martensite. The basic quenchant function is to control the rate of heat transfer from the surface of the part being quenched. Quenching Process The rate of heat extraction by a quenching medium and the way it is used substantially affects quenchant performance. Variations in quenching practices have resulted in the assignment of specific names to some quenching techniques: • Direct quenching • Time quenching • Selective quenching • Spray quenching • Fog quenching • Interrupted quenching Direct quenching refers to quenching directly from the austenitizing temperature and is by far the most widely used practice. The term direct quenching is used to differentiate this type of cycle from more indirect practices which might involve carburizing, slow cooling, reheating, followed by quenching. [...]... (650 °F) 20 5 °C (400 °F) °C °F °C/s °F/s °C/s °F/s °C/s °F/s W/m2 · K Btu/ft2 · h · °F 27 80 32. 6 58.6 14.6 26 .2 7.3 13 .2 93 82. 9 16 52. 5 32 90 32. 3 58.1 14.6 26 .3 7 .2 13.0 9 024 .6 1589.4 38 100 31.0 55.8 14.3 25 .8 7.0 12. 5 7645.4 1346.5 49 120 24 .3 43.7 14.4 25 .9 6.8 12. 2 3504 617.1 60 140 9.3 16.8 14.3 25 .7 6.3 11.3 754.0 1 32. 8 71 160 5.6 10.1 13.7 24 .6 5.9 10.6 417 73.5 82 180 4.7 8.5 13.5 24 .2 5.4 9.7... °C (1 625 °F) to 355 °C (670 °F), s At 27 °C (80 °F) At 120 °C (25 0 °F) °C Conventional °F Ni-ball Chromized Ni-ball Ni-ball Chromized Ni-ball 375 22 .5 27 .2 105 195 380 17.8 27 .9 107 170 340 16.0 24 .8 4 50 145 29 0 7.0 (a) 5 94 170 335 9.0 15.0 6 107 190 375 10.8 17.0 7 110 185 370 12. 7 19.6 8 120 190 375 13.3 17.8 9 329 23 5 455 19 .2 27.6 18.4 22 .1 10 719 24 5 475 26 .9 29 .0 25 .1... coefficient 1.1 5000 880 50 2. 1 9000 1600 100 2. 7 120 00 21 00 0.76 130 Btu/ft2 · h · °F 0.51 55 W/m2 · K 150 2. 8 120 00 21 00 0.00 0 0 .2 1000 180 0 .25 6500 1100 150 2. 4 10500 1850 0.00 0 0.5 20 00 350 50 1.0 4500 790 100 1.1 5000 880 150 1.5 6500 120 0 0.00 0 0.8 3500 620 0 .25 50 1.3 6000 1100 0.51 100 1.5 6500 120 0 0.76 110 1.5 0.76 43 100 0.51 25 % polyvinyl pyrrolidone 440 0 .25 140 25 00 0.76 60 0.6 0.51 Fast... 100 1 .2 5000 880 Air 27 80 0.00 0 0.05 20 0 35 2. 54 500 0.06 25 0 44 5.08 1000 0.08 350 62 Source: Ref 12 The data in Table 3 show that unagitated water at 32 °C (90 °F) has a Grossmann number of about 1 and an interface heat transfer coefficient of about 5000 W/m2 · K The effective interface heat transfer coefficient produced by 32 °C (90 °F) water increased to 9000 and 12, 000 W/m2 · K (1600 and 21 00... from 13, 25 , 38, and 50 mm (0.5, 1.0, 1.5, and 2. 0 in.) diameter cylindrical AISI type 304 stainless steel probes from the cooling rate, as illustrated in Fig 9 Table 2 Model parameters used to calculate the H-factor Probe diameter Model parameters mm in A B C D 12. 7 0.5 0.0 028 02 0.1857 × 10-7 1 .20 1 2. 846 25 .4 1.0 0.0 023 48 0 .25 64 × 10-9 1.508 4.448 38.1 1.5 0.0 023 09 0.57 42 × 10-9 1.749 5.076 50.8 2. 0 0.003706... results using pure nickel balls Type of quenching oil Quenching duration from 885 °C (1 625 °F) At 65 °C (150 °F) At 120 °C (25 0 °F) At 175 °C (350 °F) At 23 0 °C (450 °F) Conventional 14 -22 14 -22 Fast 7-14 7-14 Martempering without speed improvers 18-34 18-34 22 -38 ≈ 47 Martempering with, speed improvers 14 -20 13-18 16 -22 ≈ 33 High-speed motion-picture techniques have been used to reveal the influence... 8 120 190 375 13.3 17.8 9 329 23 5 455 19 .2 27.6 18.4 22 .1 10 719 24 5 475 26 .9 29 .0 25 .1 30.4 11 25 50 300 575 31.0 32. 0 31.7 32. 8 12 337 23 0 450 15.3 (b) 12. 8 (b) 13 713 24 5 475 16.4 17.9 14.0 15.6 14 Martempering, with speed improvers 190 3 Martempering, without speed improvers 1 02 2 Fast 1 24 50 300 570 19.7 17.0 15.1 15.4 (a) SUS, Saybolt universal seconds (b) Not available Table 6 Comparison... Unagitated water at 55 °C (130 °F) had an interface heat transfer coefficient of about 1000 W/m2 · K (180 Btu/ft2 · h · °F) A water velocity of 0.76 m/s (2. 5 ft/s) increased the interface heat transfer coefficient of the 55 °C (130 °F) water to 10,500 W/m2 · K (1850 Btu/ft2 · h · °F) The benefit of knowing or experimentally determining interface heat transfer coefficients produced by specific quenchants... 0.718/(5.11 · H exp 1 .28 ) (Eq 6) E3/4R = D exp 1.05/(8. 62 · H exp 0.668) (Eq 7) E1/2R = D exp 1.16/(9.45 · H exp 0.51) (Eq 8) E1/4R = D exp 1.14/(7.7 · H exp 0.44) (Eq 9) EC = D exp 1.18/(8 .29 · H exp 0.44) (Eq 10) These equations are valid for: 20 mm . 21 . R. Pradhan, J. Heat Treat., Vol 2 (No. 1), 1981, p 73- 82 22 . R. Pradhan, in Metallurgy of Continuous-Annealed Sheet Steel, B.L. Bramfitt and P.L. Mangonon, Ed., TMS-AIME, 19 82, p 20 3 -22 7. Air (b) None 2 0.9-1.0 0 .25 -0.30 0. 02 Mild 2- 2 .2 1.0-1.1 0.30-0.35 . . . Moderate . . . 1 .2- 1.3 0.35-0.40 . . . Good . . . 1.4-1.5 0.4-0.5 . . . Strong . . . 1.6 -2. 0 0.5-0.8 . . . Violent. Table 2 Model parameters used to calculate the H-factor Probe diameter Model parameters mm in. A B C D 12. 7 0.5 0.0 028 02 0.1857 × 10 -7 1 .20 1 2. 846 25 .4 1.0 0.0 023 48