Volume 07 - Powder Metal Technologies and Applications Part 8 pdf

Fig. 16 Redox curves for chromium and silicon in 316L in H 2 at atmospheric pressure. Source: Ref 17 What happens after sintering, that is, during cooling, is even more important. Figure 16 shows two scenarios. For both scenarios, the sintering temperature is 1200 °C (2192 °F). In the first scenario, the dew point of the sintering atmosphere is -40 °C (-40 °F); in the second scenario, it is -60 °C (-76 °F). In the first scenario, the stainless steel part crosses the SiO 2 /Si redox curve (Fig. 16) upon cooling at 1070 °C (1960 °F). At this temperature, the rate of oxidation of silicon is quite rapid and, therefore, rapid cooling is necessary to prevent or minimize, the formation of silicon oxides on the surface of the stainless steel part. Figure 17 shows a scanning electron microscopy (SEM) of such oxide precipitates for 316L. These precipitates do not cause the part to discolor, and they are visible only under a microscope. When tested in aqueous FeCl 3 , in accordance with ASTM G 46 (20 °C), such parts exhibit inferior corrosion resistance due to pitting. Higher-alloyed stainless steel powder such as SS-100 (20Cr-17Ni-0.8Si-5Mo) appear to be more immune to this type of corrosion (Ref 10). Fig. 17 Spheroidal silicon oxide particles formed on 316L part on cooling With the second scenario in Fig. 16, the lower dew point of -60 °C (-76 °F) causes the parts to cross the SiO 2 /Si redox curve in the cooling zone of the furnace at the much lower temperature of 890 °C (1632 °F). At that temperature, the rate of silicon oxidation either is very slow, or any oxides formed at that temperature have little effect on the corrosion resistance of the part. Figure 18 shows tentative critical cooling rate-dew point combinations as a function of dew point for three hydrogen sintered austenitic stainless steels. Upper critical cooling temperatures, that is, the lowest high temperatures where rapid cooling is to commence, are shown in Fig. 19. Fig. 18 Effect of cooling rate and dew point upon corrosion resistance (5% aqueous NaCl) of hydrogen sintered stainless steels. Dashed curves representing maximum corrosion resistance are tentative. Corrosion resistances shown in parentheses are percentages of maximum corrosion resistance for given grade and density. Fig. 19 Upper critical cooling temperature and iso-corrosion resistance curves (5% aqueous NaCl) for H 2 sintered 316L (schematic) Processing to the right of the cooling rate-dew point curves produces maximum corrosion resistance. Processing to the left results in rapid deterioration of corrosion resistance as shown schematically in Fig. 19 for 316L. It is clear from these relationships that for maximum corrosion, resistance, sintering in hydrogen requires very low dew points and/or rapid cooling after sintering. Mechanical properties of hydrogen sintered stainless steels are given in the article "Powder Metallurgy Stainless Steels" in this Volume. Sintering in Vacuum. In the early years of commercial sintering of stainless steel parts, vacuum furnaces were said to be good alternatives to other types of sintering because of their low consumption of gas. After years of experience, parts producers learned that in addition to high initial capital cost, vacuum furnaces also were costly to maintain. Nevertheless, it is clear from the previous section on sintering in hydrogen that with the typical furnaces (belt, pusher, and walking beam) used presently in the industry, the number one property of stainless steel, superior corrosion resistance, is for most presently used P/M stainless steels not attainable to a sufficient degree. Therefore, future P/M opportunities that require excellent corrosion resistance cannot be realized until furnace manufacturers construct furnaces that are capable of lower dew points and parts producers equip their furnaces with (already available) rapid cooling devices. This is where vacuum furnaces are used. With a state of the art vacuum furnace, it is much easier to maintain a low dew point and to obtain rapid cooling than it is with a typical atmosphere furnace. Nevertheless, certain precautions are necessary. For maximum corrosion resistance of vacuum sintered stainless steel, surface depletion of chromium due to high-vapor pressure and the presence of original surface oxides must be minimized. Sintering under a partial pressure of nitrogen or argon of 1500 m of mercury effectively reduces chromium losses. Reference 9 shows that after high temperature sintering (>1205 °C, or 2200 °F), where chromium losses are more severe, holding the parts prior to cooling for a short time at a lower temperature, or by increasing the partial pressure of argon to 1 at the lower temperature, significantly improves corrosion resistance. Both measures allow the parts to replenish (from the interior) surface chromium that was lost at the high-sintering temperature. In spite of the high-oxygen contents of water atomized stainless steel powders, vacuum sintered stainless steel parts usually are bright. This is because some of the surface oxides, during sintering, diffuse into the interior. In the absence of an external-reducing gas atmosphere, vacuum-sintered stainless steel parts have relatively low carbon and oxygen contents due to the reaction between carbon and oxygen (or oxides) particularly at high-sintering temperatures, to form carbon monoxide (Fig. 20) (Ref 18 and 19). There is evidence, however, that the typical amounts of carbon present in a stainless steel part after delubrication are insufficient for removing most of the original oxide particles present on the outer surfaces of a part. These unreduced original oxide particles give rise to pitting corrosion. Admixing small amounts of graphite to the stainless steel powder, overall oxide reduction, particularly at the higher-sintering temperatures, is greatly enhanced and is also sufficient to reduce surface oxides. Alternatively, a small partial pressure of H 2 should accomplish the same. Fig. 20 Oxygen versus carbon contents of vacuum and atmosphere sintered P/M austenitic stainless steels of varying compositions When making graphite additions to a stainless steel powder, it should be kept in mind that the carbon content of the sintered part will increase. Thus, the optimum graphite addition is the maximum addition that produces no chromium carbide precipitates in the cooling zone of the sintering furnace. It depends, among other factors, on the composition of the stainless steel, the oxygen content of powder, the sintering temperature, and the cooling rate employed. Vacuum-sintered stainless steels should be rapidly cooled in a non-oxidizing gas to prevent the formation of deleterious surface oxides. Cooling in nitrogen will result in the formation and precipitation of chromium nitrides on the surfaces of the parts. The attendant chromium depletion will cause the parts to have very low-corrosion resistance. Mechanical properties of vacuum sintered stainless steels are shown in the article "Powder Metallurgy Stainless Steels" in this Volume. Sintering in H 2 -N 2 Gas Mixtures. Sintering at 1120 °C (2050 °F) in dissociated ammonia was the most widely used method for sintering stainless steels in the 1960s and 1970s. Dissociated NH 3 was not only less expensive than H 2 , but it also increased the strength of the sintered parts, although at some reduction in ductility, to levels comparable to wrought stainless steels of the same composition, at densities of 85 to 90% of theoretical. The strengthening is the result of nitrogen absorption during sintering. The amount of nitrogen absorbed follows known phase equilibria in accordance with Sievert's law, that is, nitrogen absorption is proportional to the square root of the partial pressure of nitrogen in the sintering atmosphere. (See the article "Corrosion-Resistant Powder Metallurgy Alloys.") The relationships for 304L are shown in Fig. 21 (Ref 20), showing both the amount of nitrogen absorbed as a function of sintering temperature and sintering atmosphere (dissociated ammonia and nitrogen) and the strength increase due to nitrogen absorption. Fig. 21 (a) Effect of nitrogen content on ultimate tensile strength and elongation of 304L stainless steel. (b) Effect of sintering temperature on amount of absorbed nitrogen for 304L. Source: Ref 20 The problems with nitrogen absorption and chromium nitride (Cr 2 N) precipitation during cooling (after sintering), and sensitization and loss of corrosion resistance, when sintering is done in dissociated ammonia, are described in Ref 21. They were not appreciated for many years, in part because corrosion resistance demands were modest and/or corrosion resistance was not assessed. Later, with increasing demands for improved corrosion resistance, and as more quantitative information on the effect of sintering in dissociated ammonia became available, recommendations were made to limit nitrogen absorption to some 3000 ppm in austenitic stainless steels. While this limitation seemed to satisfy some corrosion resistance requirements in an acidic environment, it was unsatisfactory for parts tested in aqueous NaCl. Even with small amounts of Cr 2 N precipitates, rust spots would form in a short time. Figure 22 shows examples of Cr 2 N precipitates in sintered austenitic stainless steels. Fig. 22 Chromium nitride precipitates in 316L (a) sintered at 1150 °C (2100 °F) in dissociated NH 3 ; 4500 ppm N 2 ; Cr 2 N precipitates along grain boundaries (1), and within grains (2). (b) sintered at 1120 °C (2050 °F) in dissociated NH 3 and slowly cooled; 6500 ppm N 2 ; Cr 2 N in lamel lar form near surface (1) and as grain boundary precipitates in the interior (2). In an early study, Sands et al. (Ref 22) pointed out that 316L sintered in dissociated NH 3 required a cooling rate of 200 °C/min for preventing nitrogen absorption and precipitation of Cr 2 N. More recently, Frisk et al. (Ref 23) determined in a laboratory study that sintering of 316L in dissociated NH 3 at 1250 °C (2280 °F) required cooling rates of >450 °C/min (Fig. 23). The higher critical cooling rate of Frisk et al. can probably be ascribed to their much lower dew point (-100 °C versus -40 to -60 °C for Sands), which allows for more rapid nitrogen absorption during cooling as illustrated in Fig. 24 (Ref 24) for the bright annealing of stainless steel strip. The deleterious reactions of increasing nitrogen absorption with decreasing dew point, and of increasing oxidation with increasing dew point, leave a relatively narrow dew point window for sintering in dissociated ammonia. Fig. 23 Effect of cooling rate on presence of chromium nitrides in microstructure of 316L parts sintered at 1250 °C in dissociated NH 3 . Source: Ref 23 Fig. 24 Effect of dew point on nitrogen absorption and oxidation of 316L shim, disk, and ba r stock annealed for 15 min at 1038 °C (1900 °F) in 30%H 2 -70%N 2 . It took 2.3; 2.8; and 4.7 min respectively to cool the three materials from 1038 °C (1900 °F) to 538 °C (1000 °F). Source: Ref 24 These cooling rate requirements are even higher than those for sintering in hydrogen to prevent oxidation during cooling, and the cooling-rate dew point relationship appears to be reversed. It is, therefore, not surprising that stainless steel parts producers are increasingly shifting towards hydrogen sintering at the expense of sintering in dissociated NH 3 . Sintering in an atmosphere of 10%N 2 -90%H 2 may be a more practicable compromise that greatly reduces the high- cooling rate requirements of dissociated NH 3 to more manageable levels, while still benefiting substantially from the solid solution strengthening obtainable with the lower nitrogen concentration. Good corrosion resistance for such conditions were reported by Larsen (Ref 25) and Mathiesen (Ref 26). The positive effect of nitrogen on corrosion resistance, as documented for wrought stainless steels, is expected to apply equally to sintered stainless steels. However, this beneficial effect has not yet been well documented for. sintered stainless steels, probably because of the overshadowing negative effect of excessive nitrogen absorption on the surface of parts from the sintering atmosphere during cooling. Mechanical and other properties for 316L and 434L sintered in dissociated NH 3 are given in the article "Powder Metallurgy Stainless Steels" in this Volume. References cited in this section 9. E. Klar and P.K. Samal, Optimization of Vacuum Sintering Parameters for Improved Corrosion Resistance of P/M Stainless Steels, Advances in Powder Metallurgy and Particulate Materials, Vol 7, Metal Powder Industries Federation, 1994, p 239-251 10. E. Klar and P.K. Samal, Effect of Density and Sintering Variables on the Corrosion Resistance of Austenitic Stainless Steels, Advances in Powder Metallurgy & Particulate Materials, Vol 3, Metal Powder Industries Federation, 1995, p 11-3 to 11-17 17. E. Maahn, S.K. Jensen, R.M. Larsen, and T. Mathiesen, Factors Affecting the Corrosion Resistance of Sintered Stainless Steel, Advances in Powder Metallurgy & Particulate Materials, Vol 7, 1994, p 253-271 18. C. Lall, Fundamentals of High Temperature Sintering: Application to Stainless Steels and Soft Magnetic Alloys, Int. J. of Powder Metall., Vol 27 (No. 4), 1991, p 315-329 19. E. Klar, M. Svilar, C. Lall, and H. Tews, Corro sion Resistance of Austenitic Stainless Steels Sintered in Commercial Furnaces, Advances in Powder Metallurgy & Particulate Materials, Vol 5, Metal Powder Industries Federation, 1992, p 411-426 20. N. Dautzenberg, Paper No. 6.18, Eigenschaften von Sinters tählen aus Wasserverdüsten Unlegierten und Fertiglegierten Pulvern, 2nd European Symposium on Powder Metallurgy, 1968, EPMA, Vol II 21. M.A. Pao and E. Klar, On the Corrosion Resistance of P/M Austenitic Stainless Steels, Proceedings of the International Powder Metallurgy Conference (Florence, Italy), Associazone Italiano di Metallurgia, 1982 22. R.L. Sands, G.F. Bidmead, and D.A. Oliver, The Corrosion Resistance of Sintered Stainless Steels, Modern Developments in Powder Metallurgy, Vol 2, H.H. Hausner, Ed., Plenum Press, 1966, p 73-85 23. K. Frisk, A. Johanson, and C. Lindberg, Nitrogen Pick up During Sintering of Stainless Steel, Advances in Powder Metallurgy & Particulate Materials, 1992, Vol 3, Metal Powder Industries Federation, p 167-179 24. R.H. Shay, T.L. Ellison, and K.R. Berger, Control of Nitrogen Absorption and Surface Oxidation of Austenitic Stainless Steels in H-N Atmospheres, Progress in Powder Metallurgy, Vol 39, H.S. Nayar, S.M. Kaufman, and E.E. Meiners, Ed., Metal Powder Industries Federation, 1983, p 411-430 25. M. Larsen, "Debindering and Sintering in Powder Metallurgy Processes," Ph.D. thesis, Technical University of Denmark, 1994 (in Danish) 26. T. Mathiesen, "Corrosion Properties of Sintered Stainless Steels," Ph.D. thesis, Techn ical University of Denmark, 1993 (in Danish) Production Sintering Practices Sintered Density For many years, inferior corrosion resistance of sintered stainless steel parts has been and continues to be mistakenly attributed to the presence of pores in accordance with the mechanism of crevice corrosion. That this hypothesis is untenable follows from widely available evidence (Ref 7) that sintered stainless steel parts of identical composition and similar pore volumes, pore sizes, and pore shapes, but sintered under varying conditions, may have corrosion resistances (in 5% aqueous NaCl) that can vary by two orders of magnitude. Furthermore, a comparison of wrought and sintered (85% of theoretical density) type 316L for susceptibility to crevice corrosion in 10% FeCl 3 , in accordance with ASTM G 48, showed that the wrought part was actually more severely attacked than the porous part (Ref 7, 27). More recent studies have shown that most cases of inferior corrosion resistance of sintered stainless steels are due to incorrect sintering, as previously described. Also, for many years, there existed controversy as to the effect of sintered density upon corrosion resistance. Corrosion testing in an acidic environment, such as dilute H 2 SO 4 , HCl, and HNO 3 , always showed that corrosion resistance improved with increasing density. Testing in a neutral salt solution, however, showed the corrosion resistance to decrease with increasing density (Ref 17, 28, 29, 30, 31), when corrosion testing was done in long-term immersion or salt spray tests, whereas higher density was found to be beneficial in short-term potentiodynamic polarization tests (Ref 32). Maahn and Mathiesen (Ref 8) attributed the failure to observe this important relationship between corrosion resistance and density in short-term, potentiostatic anodic polarization tests to the lack of time for the time-consuming build up of localized attack within the pores, in analogy to the mechanism of crevice corrosion. By using slow stepwise polarization, the expected relationship, that is, a decrease of the stepwise initiation potential with increasing density (equivalent to deteriorating corrosion resistance), was observed (Ref 33). Recent studies, with austenitic stainless steels (Ref 10, 17, 31) showed that the corrosion resistance in 5% aqueous NaCl (by immersion) can be reduced by up to two orders of magnitude due to the presence of porosity. The negative effect appears at sintered densities of 6.7 g/cm 3 , reaches a minimum corrosion resistance between 6.9 and 7.2 g/cm 3 , depending on pore morphology, and thereafter disappears at densities of 7.4 g/cm 3 (Fig. 25). To the left of the minimum, corrosion resistance decreases with decreasing pore size due to increasing oxygen depletion and failure to maintain the passive layer. To the right of the minimum, corrosion resistance improves as pores become closed off and inaccessible to the surface. This type of corrosion can be reduced by impregnating the pores with a resin, by metallurgical modification of the pore surfaces, or by the use of higher-alloyed stainless steels, particularly those containing higher concentrations of molybdenum. Fig. 25 Effect of density on corrosion resistance (B rating) of pressed, sintered, repressed, and annealed 317L parts Another approach to avoiding the problem of long-term corrosion in a neutral salt solution due to the presence of certain size pores is to make use of the various forms of liquid-phase sintering (transient, persistent, and supersolidus) and to achieve sintered densities >7.4 g/cm 3 . For austenitic stainless steels, silicon additions of several percent (Ref 34, 35) or smaller amounts of boron (Ref 17), have been used. For ferritic stainless steels phosphorus additions (Ref 35) have been used. The large shrinkage occurring during liquid-phase sintering is often accompanied by increasing loss of dimensional stability. Sizing is usually employed to establish dimensional accuracy. Also, depending on the composition of the liquid- phase sintering additive, secondary phases may be present after sintering, which can have a negative effect on corrosion resistance (Ref 17). References cited in this section 7. E. Klar, Corrosion of Powder Metallurgy Materials, Corrosion Metals Handbook, 9th ed., ASM International, 1987, p 823-845 8. E. Maahn and T. Mathiesen, "Corrosion Properties of Sintered Stainless Steel," p resented at UK Corrosion 91 (Manchester), 1991 10. E. Klar and P.K. Samal, Effect of Density and Sintering Variables on the Corrosion Resistance of Austenitic Stainless Steels, Advances in Powder Metallurgy & Particulate Materials, Vol 3, Metal Powder Ind ustries Federation, 1995, p 11-3 to 11-17 17. E. Maahn, S.K. Jensen, R.M. Larsen, and T. Mathiesen, Factors Affecting the Corrosion Resistance of Sintered Stainless Steel, Advances in Powder Metallurgy & Particulate Materials, Vol 7, 1994, p 253-271 27. D. Ro and E. Klar, Corrosion Behavior of P/M Austenitic Stainless Steels, Modern Developments in Powder Metallurgy, Vol 13, H.H. Hausner and P.W. Taubenblat, Ed., Metal Powder Industries Federation, 1980, p 247-287 28. M. Svilar and H.D. Ambs, P/M Martensitic Stainless Steels: Processing and Properties, Advances in Powder Metallurgy, Vol 2, 1990, p 259-272 29. S.K. Chatterjee, M.E. Warwick, and D.J. Maykuth, The Effect of Tin, Copper, Nickel, and Molybdenum on the Mechanical Properties and Corrosion Resistance of Sintered Stainless Steel (AISI 304L), Modern Developments in Powder Metallurgy, Vol 16, E.N. Aqua and C.I. Whitman, Ed., Metal Powder Industries Federation, 1984, p 277-293 30. F.M.F. Jones, The Effect of Processing Variables on the Properties of Type 316L Powder Compacts, Progress In Powder Metallurgy, Vol 30, Metal Powder Industries Federation, 1970, p 25-50 31. R.M. Larsen, T. Mathiesen, and K.A Thorsen, The Effect of Porosity and Oxygen Content on the Corrosion Resistance of Sintered 316L, Powder Metallurgy World Congress (Paris), Vol III, Les Editions de Physique, 1994, p 2093-2096 32. G. Lei, R.M. German, and H.S. Nayar, Corrosion Control in Sintered Austenitic Stainless Steels, Progress in Powder Metallurgy, Vol 39, H.S. Nayar, S.M. Kaufma n, and K.E. Meiners, Ed., Metal Powder Industries Federation, 1984, p 391-410 33. T. Mathiesen and E. Maahn, Corrosion Behavior of Sintered Stainless Steels in Chloride Containing Environments, 12th Scandinavian Corrosion Congress (Helsinki), 1992, p 1-9 34. W.F. Wang and Y.L. Su, Powder Metallurgy, Vol 29, 1986 35. N.S. Mikkelsen and M. Jensen, "Sinterwerkstoff", European Patent Publication 0564778 A1, 1993 Production Sintering Practices Sintering of High-Speed Steels and Tool Steels Powder metallurgy processing offers several advantages to costly and highly alloyed tool steel materials. These advantages include uniform and finer microstructure, improved grindability, improved cutting performance, and capabilities of high-speed steels and tool steel alloys that cannot be made by conventional ingot metallurgy. The use of conventional press-and-sinter technology offers the additional advantage of net shape or near-net shape capability. Cutting tools, bearings, and wear parts are being produced commercially by fully dense sintering. Wear parts also are produced by conventional P/M techniques to densities of 80 to 90%. Current commercial fully dense sintering uses high green strength and compressible water-atomized tool steel powders that are compacted in rigid dies using uniaxial pressing or in flexible rubber molds using cold isostatic pressing to make green tools and parts. These parts are then sintered in a microprocessor-controlled vacuum furnace near the solidus temperature of the alloy to virtually full density (at least 98% and frequently 99+% of theoretical). The flexibility of this process allows production of pressed and sintered tool and die parts in various forms. Generally, parts are pressed to 70 to 85% of theoretical density before sintering to full density. Pressing to lower green densities tends to require longer times at temperature to obtain full density and results in coarser microstructures. Pressing at higher pressures, which is required for increased green densities, results in increased tool wear and breakage. Parts with high green density also may require extra care in sintering. The surface can sinter rapidly to high density and entrap gases from the center of the part. Production Sintering Practices Sintering Mechanisms Important contributions to sintering arise from diffusion and viscous flow. Diffusion rates increase with increasing temperature because of the increased number of vacancies that promote diffusion of substantial alloying elements. Sintering temperature must be held very close to the solidus temperature to attain full density in a reasonable time. A sintering temperature 5.5 to 11 °C (10 to 20 °F) above the solidus reportedly forms a small amount of liquid, which allows high-diffusion rates for enhanced sintering. These results are based on relatively rapid densification at high temperatures. However, metallographically, there is no resemblance between a successfully sintered high-speed steel and a typical liquid-phase sintered material, such as tungsten carbide or tungsten heavy metal. Additional research is required to fully understand high-temperature sintering. Factors Affecting Sintering. Sintering is a series of complex processes, of which densification is only one phase. As green parts are heated to the sintering temperature, gases can be adsorbed (nitrogen, hydrogen, or oxygen) or evolved (nitrogen, hydrogen, or carbon monoxide). Admixed carbon dissolves and homogenizes, while carbides dissolve and grow. Grain growth also occurs as pores shrink and virtually disappear. Sintering Time and Temperature. Increasing the sintering temperature decreases the amount of time required to achieve full density. Higher temperatures also reduce the time between reaching full density and oversintering. These relationships are shown schematically in Fig. 26. A sintering curve for M2 high-speed steel is shown in Fig. 27. Table 13 provides typical properties of M2 tool steel for various sintering temperatures. Table 13 Sintering data for M2 tool steel Testing was conducted on five lots. All compacts were pressed at 827 MPa (60 tsi) and sintered for 60 min. Sintered properties: At 1240 °C (2260 °F) At 1250 °C (2280 °F) At 1260 °C (2300 °F) Total carbon (a) , % Added graphite, % Density, g/cm 3 Characteristics Density, g/cm 3 Characteristics Density, g/cm 3 Characteristics Lot 1 0.77 0.0 7.19 MLP, FG 7.29 FG, SP 7.47 MSP, VFG 0.87 0.1 7.25 VFG, LP 8.03 FP, LP 8.04 FP, SP 0.97 0.2 7.48 VFG, LP 8.11 FP 8.09 FP 1.07 0.3 8.01 FG 8.09 FP, SP 8.07 E, LP Lot 2 0.79 0.0 7.00 VFG, MP 7.25 FG, MP 7.77 MP 0.89 0.1 7.18 VFG, MP 7.76 MP 8.08 FP 0.99 0.2 7.58 MP, VFG 8.08 FP 8.06 FP, LG 1.09 0.3 8.07 VFP 8.00 LG, SP 8.05 E, LP, FP Lot 3 0.86 0.0 7.17 VFG, MP 7.47 MP 8.09 FP 0.96 0.1 7.29 VFG, MP 8.08 MLP 8.08 FP, LP 1.06 0.2 7.84 VFG, MP 8.09 SP 8.06 E, LP, LG 1.16 0.3 8.07 FG, VFP 8.07 SP, E 8.07 E, LG [...]... 6 6-6 7 6 7-6 8 6 8- 6 9 DPH 79 0 -8 15 81 5 -8 40 84 0 -8 70 81 5 -8 40 84 0 -8 70 87 0-9 05 90 5-9 40 79 0 -8 15 81 5 -8 40 84 0 -8 70 79 0 -8 15 81 5 -8 40 84 0 -8 70 87 0-9 05 79 0 -8 15 81 1 -8 40 84 0 -8 70 84 0 -8 70 87 0-9 05 90 5-9 40 94 0-9 80 87 0-9 05 90 5-9 40 94 0-9 80 Hardening temperature(a) °C °F 1170 21 38 1 180 2156 1200 2192 1170 21 38 1 180 2156 1200 2192 1220 22 28 1 180 2156 1200 2192 1200 2192 1 180 2156 1200 2192 1200 2192 1210 2210 1 180 2156 1 180 2156 1200... 2190 1000 183 0 1000 183 0 1000 183 0 Carbon activity 0. 6-0 .25 0. 6-0 .25 0. 6-0 .25 0. 6-0 .25 0. 6-0 .25 0. 6-0 .25 0. 5-0 .2 0. 5-0 .2 0. 5-0 .2 Hydrogen, % 99.94 4-9 9.976 99.4 5-9 9.47 89 .9 6 -8 9. 98 99 .8 4-9 9.93 99.2 7-9 9.40 89 .7 9 -8 9 .87 99.5 6-9 9 .82 99.0 5-9 9.29 89 .4 3 -8 9.34 (a) Carbon monoxide, % None 0.5 10 None 0.5 10 None 0.5 10 Water, ppm None 3. 4 -8 60. 0-1 44.0 None 12. 2-2 9.2 220. 0-5 70.0 None 114. 0-2 90.0 2 080 . 0-5 200.0 Carbon... ksi 48 7 T6 165 Green strength MPa psi 3.1 450 Tensile strength(a) MPa ksi 110 16 T4 601AB Green density % g/cm3 85 2.29 Temper 186 27 172 25 3 T1 169 24.5 145 24 2 T4 210 30.5 179 26 3 T6 2 48 36 2 48 36 0 5 5-6 0 HRH 8 0 -8 5 HRH 7 0-7 5 HRE 6 0-6 5 HRH 8 0 -8 5 HRH 7 5 -8 0 HRE 6 5-7 0 HRH 8 5-9 0 HRH 8 0 -8 5 HRE 5 5-6 0 HRH 6 5-7 0 HRH 5 5-6 0 HRE 5 5-6 0 HRH 7 0-7 5 HRH 6 5-7 0 HRE 6 0-6 5 HRE 7 0-7 5 HRE 8 0 -8 5 HRE 180 Coldformed parts... hardening and tempering temperatures, and the average Snyder-Graaf intercept grain size for several high-speed steels Table 15 Recommended conditions for salt bath hardening of sintered high-speed steels Alloy M2 M3 type 2 M4 M15 M35 T15 T42 (a) (b) Requires hardness HRC 6 3-6 4 6 4-6 5 6 5-6 6 6 4-6 5 6 5-6 6 6 6-6 7 6 7-6 8 6 3-6 4 6 4-6 5 6 5-6 6 6 3-6 4 6 4-6 5 6 5-6 6 6 6-6 7 6 3-6 4 6 4-6 5 6 5-6 6 6 5-6 6 6 6-6 7 6 7-6 8 6 8- 6 9 6 6-6 7 6 7-6 8. ..Lot 4 0 .87 0.97 1 .07 1.17 Lot 5 1.00 1.10 1.20 1.30 0.0 0.1 0.2 0.3 7.22 7.47 8 .07 8. 08 VFG, MP FG, MP FP FP 7 .87 8. 08 8.06 8. 06 MP FP, SP FP, SP E, SP 8. 09 8 .07 8. 06 8. 05 FP, LP SP E, MSP E, MSP 0.0 0.1 0.2 0.3 7 .82 8. 08 8 .07 8. 06 VFG FG VFP LG 8. 09 8 .07 8. 06 8. 06 FP, SP FP, SP LG, E LG, E 8. 06 8. 05 8. 05 8. 04 E, FP, LG E, EP, LG E, FP, LG E, FP, LP M, many;... sintered M2, M35, and T15 high-speed steels Property Density, g/cm3 Ultimate tensile strength(a), MPa (ksi) Grade M2 8. 0 5 -8 .2 75 0 -8 00 M35 8. 0 5 -8 .2 77 0 -8 20 T15 8. 1 5 -8 .3 77 0 -8 30 Elongation(a), % Ultimate tensile strength(b), MPa (ksi) (a) (b) 1 2-1 4 75 0-2 000 (10 8- 2 90) 6-9 77 0-2 000 (11 2-2 90) 3-6 77 0-2 000 (11 2-2 90) Fully annealed Depending on heat treatment Fig 30 Tempering curves for high-speed steels at... ppm None 4.7 × 1 0-3 -1 .1 × 1 0-2 1. 7-4 .0 None 3 × 1 0-2 -7 .2 × 1 0-2 12. 2-2 9.2 None 0.3 4-0 .85 135. 0-3 38. 0 Methane, % 0.05 6-0 .023 0.05 5-0 .023 0.04 5-0 .019 0.2 4-0 .10 0.2 4-0 .10 0.1 9-0 . 08 0.4 4-0 . 18 0.4 4-0 . 18 0.3 6-0 .14 At pressure of 1 atm Typical commercial practice for hydrogen sintering carbon control is based on partial reaction of the incoming gas (hydrogen plus impurity levels of water vapor and oxygen), which... 413 13 227 33 147 21.3 7.3 T2 T4 T6 T8 2 38 236 274 280 33.9 34.3 39 .8 40.6 216 1 48 173 250 31.4 21.5 25.1 36.2 2.3 8 8.7 3 7 0-7 5 HRE 7 5 -8 0 HRE 8 5-9 0 HRE 7 0-7 5 HRE 8 0 -8 5 HRE 9 0-9 5 HRE 5 5-6 0 HRH 7 0-7 5 HRH 4 5-5 0 HRE 80 HRE 70 HRE 85 HRE 87 HRE Tensile properties determined using powder metal flat tension bar (MPIF standard 1 0-6 3), sintered 15 min at 620 °C (1150 °F) in nitrogen Aluminum preforms can be... 1 180 2156 1200 2192 1170 21 38 1 180 2156 1200 2192 1220 22 28 1160 2120 1 180 2156 1210 2210 Tempering temperatures(b) °C °F 570 10 58 560 1040 550 1022 570 10 58 570 10 58 560 1040 550 1022 570 10 58 560 1040 540 1004 560 1040 580 1076 570 10 58 560 1040 580 1076 560 1040 560 1040 570 10 58 565 1049 550 1022 520 9 68 570 10 58 570 10 58 560 1040 Intercept grain size 15 15 15 14 10.5 8 8 17 17 17 15 12 12 12 17... strain) 180 13 13 (a) 90 90 2.49 2.49 5.4 5.4 780 780 97.0 92.4 92.4 2. 58 2.70 2.56 2.56 3 245 35.6 205 29 .8 3.5 323 46 .8 322 46.7 0.5 T1 209 30.3 181 26.2 3 262 38 214 31 5 332 48. 1 327 47.5 2 T1 160 23.2 75 10.9 10 194 28. 2 119 17.2 8 T6 180 2000 92.9 24.6 T4 202AB Compacts 13 .8 1200 170 T6 2.64 8. 3 29.2 T4 95 2.50 201 T6 30 90 T1 T4 413 13 227 33 147 21.3 7.3 T2 T4 T6 T8 2 38 236 274 280 33.9 34.3 39.8 . 6 4-6 5 81 5 -8 40 1200 2192 560 1040 8 M4 6 5-6 6 84 0 -8 70 1200 2192 540 1004 8 6 3-6 4 79 0 -8 15 1 180 2156 560 1040 17 6 4-6 5 81 5 -8 40 1200 2192 580 1076 17 6 5-6 6. Density, g/cm 3 8. 0 5 -8 .2 8. 0 5 -8 .2 8. 1 5 -8 .3 Ultimate tensile strength (a) , MPa (ksi) 75 0 -8 00 77 0 -8 20 77 0 -8 30 Elongation (a) , % 1 2-1 4 6-9 3-6 75 0-2 000 77 0-2 000 77 0-2 000 Ultimate. 6 3-6 4 79 0 -8 15 1170 21 38 570 10 58 . . . 6 4-6 5 81 5 -8 40 1 180 2156 560 1040 . . . M2 6 5-6 6 84 0 -8 70 1200 2192 550 1022 . . . 6 4-6 5 81 5 -8 40 1170 21 38 570 1058