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The influence of grain size on the upper shelf impact energy of a 9Ni steel
Scripta METALLURGICA et MATERIALIA
Vol. 27, pp. 881-854, 1992 Printed in the U.S.A.
Pergamon Press Ltd. All rights reserved
T H E I N F L U E N C E O F G R A I N S I Z E ON T H E UPPER SHELF IMPACT E N E R G Y OF A 9Ni S T E E L A.L. Wojcieszynski W.M. Garrison, Jr. A.W. Thompson Dept. of Metallurgical Engineering and Materials Science Carnegie Mellon University Pittsburgh, PA 15213 USA
(Received March 27, 1992) (Revised July 21, 1992) Introduction Experimental results indicate that large grain sizes will reduce the cleavage fracture stress and increase the ductile to brittle transition temperature of martensitic, bainitic and ferritic steels (1). However, there does not appear to be a similar consensus on the effects of grain size on upper shelf toughness. In part this situation has arisen because grain size is typically increased in steels by increasing the austenitizing temperature, and higher austenitizing temperatures can reduce the volume fractions and alter the spacings of void initiating particles such as carbides and sulfide inclusions (2,3). Therefore, while increasing the austenitizing temperature can, in many instances, result in improved upper shelf toughness (4-6), it is very difficult to extract from such results any effect of grain size on toughness due to possible changes in particle dispersions. To unambiguously assess the role of grain size and toughness, it would be desirable to vary the grain size without influencing the inclusion volume fraction and spacing and without complications due to undissolved carbides. The purpose of this paper is to report the results of experiments which suggest that increasing grain size can have a significant and detrimental effect on upper shelf toughness, as measured here by the Charpy impact energy. Materials, Heat Treatments and Procedures The material used in this work was a 0.07wt.%C/9wt.%Ni/0.2wt.%Cr steel (7). This steel was selected because of the low carbon and chromium content, which would minimize the formation of carbides, because it is known to have reasonably high toughness in the as-quenched condition after austenitizing at 790°C and because it remains wholly austenitic to temperatures as low as 700°C. This latter characteristic is useful in maintaining inclusion volume fraction and spacing while changing grain size. The material was produced by vacuum induction melting followed by vacuum arc remelting. The ingot was subsequently upset and cross forged to a bar 5 cm thick and 8.9 cm wide. The chemistry of the material is given in Table 1. Four heat treatments were used. Heat treatment 1 consisted of air cooling after 1 h at 900°C followed by austenitizing at 790°17 for 1 h and a water quench; this heat treatment results in a relatively small prior austenite grain size. Heat treatment 2 consisted of air cooling after 1 h at 900°C followed by 1 h at 1260°C and a water quench. This heat treatment was used to determine what effects reducing inclusion volume fraction coupled with a large grain size would have on toughness. Heat treatment 3 consisted of several steps. The material was air cooled after 1 h at 900°C; the material was then held at 1260°C for 1 h, quenched directly into a salt pot at 750°C, held for 15 h, and then water quenched. This heat treatment results in the large prior austenite grain size of heat treatment 2, but a larger inclusion volume fraction. Heat treatment 4 was identical to heat treatment 3 except the material was given a final grain refining treatment consisting of 1 h at 750°C followed by a water quench.
851 0956-716X/92 $8.00 + .00 Copyright (c) 1992 Pergamon Press
Ltd.
852
IMPACT ENERGY
Vol.
27, No.
Table 1 Alloy Chemistry*
C
Ni
Cr
Mn
Mo
0.077
9.29
0.26
0.37
0.01
Si 0.008
S
P
A1
02**
N2**
0.002
0.003
0.005
7
4
* in wt%; ** in wt ppm Table 2 Mechanical Properties Condition
H.T. H.T. H.T. H.T.
1 2 3 4
oy (MPa)
UTS (MPa)
ef
Cv (J)
822 819 828 850
1128 1090 1095 1149
1.29 1.22 0.74 1.17
167 249 26 83
ay - yield strength; UTS - ultimate tensile strength; e f - the strain to fracture; Cv = Charpy impact energy
Table 3 Microstructural Data Condition
H.T. H.T. H.T. H.T.
1 2 3 4
f
0.00027 0.00004 0.00016 0.00014
Ro
XO
(lain)
(~tm)
0.119 0.095 0.113 0.115
1.59 2.41 1.96 1.98
grain size (btm) 13.3 151 151 13.8
f - inclusion volume fraction; R0 - average inclusion radius calculated from R0 = (x/4)H(d) where H(d) is the harmonic mean of apparent diameters on polished cross-sections (12); X0 - inclusion spacing calculated from X0 = 0.89 R0f-1/3 (13).
The properties measured were the room temperature smooth axisymmetric tensile properties and the Charpy impact energy. The tensile specimens had gage diameters and lengths of 6.35 mm and 25.4 mm respectively and were tested at a crosshead speed of 0.5 mm/min. Grain sizes were obtained according to the linear intercept method (8). Inclusion volume fractions and average radii were determined from 5000X SEM back scatter images of polished cross sections. Results and Discussion The mechanical properties measured for the four heat treatments are listed in Table 2. The four heat treatments result in very similar strength levels. However, the strains to fracture and Charpy impact energies were considerably lower after heat treatment 3 than after the other three heat treatments. Microstructural information is
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Vol.
27, No.
FIG. 1
7
IMPACT ENERGY
Fracture Surfaces of Charpy impact specimens of material produced by (a) heat treatment 1, (b) heat treatment 2, (c) heat treatment 3 and (d) heat treatment 4.
summarized in Table 3. The measured inclusion volume fraction for material receiving heat treatment 1 was 0.00027 while the inclusion volume fraction after heat treatment 2 was much lower, 0.00004; this decrease in inclusion volume fraction was due to dissolution of MnS at the austenitizing temperature of 1260°C. After heat treatment 3 the inclusion volume fraction increased to 0.00016, due to precipitation of MnS particles during the hold at 750°C for 15 h. Heat treatment 4 refined the grain size achieved by heat treatment 3, and the inclusion volume fraction after heat treatment 4 was essentially the same as after heat treatment 3. Our primary interest was to compare the toughnesses measured for heat treatments 1 and 3, which had Charpy impact energies of 167J and 26J, respectively. The fracture surfaces associated with the two heat treatments were microvoid coalescence (Fig. 1), with no evidence of overheating effects on the fracture surface of heat treatment 3. The fracture surfaces are very similar in terms of void size and spacing in some regions but there are regions on the fracture surface of heat treatment 1 where the voids are much larger than observed for heat treatment 3; these regions appear to be associated with the walls of the zig-zag fracture path. Taken at face value the data indicate that the low toughness of heat treatment 3 compared to heat treatment 1 is due to the larger grain size. However, the inclusion volume fraction measured for heat treatment 3 was only half that measured for heat treatment 1. That the inclusion volume fraction measured after heat treatment 3 was only 0.00016 could stem from two causes. First, it is possible that not all of the sulfur has precipitated during the hold at 750°C. Second, it is possible that after heat treatment 3 there are small sulfides not visible on polished cross sections at 5000X; if the latter is true, it is possible that all of the sulfur has reprecipitated during the hold at 750°C and that the inclusion volume fraction is the same as for heat treatment 1 but the inclusion size is smaller than for heat treatment 1. If the latter possibility is true, the low impact energy of heat treatment 3 could be due to a much smaller inclusion spacing than calculated and not to the large grain size.
853
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IMPACT ENERGY
Vol.
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To refine the grain size obtained after heat treatment 3, specimens receiving heat treatment 3 were austenitized at 750°C; this resulted in a grain size of 13.8 larn and an increase in the Charpy impact energy from 26J to 83J. The microstructural data in Table 3 indicate that reaustenitizing at 750°C had little effect on the sulfide distribution. The fracture mode after heat treatment 4 was not transgranular but was a ductile intergranular fracture following the grain boundaries established at 1260°C, and the void nucleating particles on these grain boundaries were MnS particles; this fracture mode is typical of overheating and the improved toughness of heat treatment 4 is consistent with studies of overheating (9,10) which show that the ductile intergranular fracture mode is, at least for lower strength steels, not associated with particularly low Charpy impact energies. That the ductile intergranular fracture mode is observed only after the grain-refining heat treatment is also consistent with studies of overheating which suggest that without the grain refining step there is a competition between the transgranular and grain boundary fracture paths, which is influenced by the ratio of the grain boundary and matrix inclusion spacings, with large ratios favoring the transgranular fracture mode (11). After heat treatment 3 the matrix spacing was apparently too small to permit the intergranular fracture. The microstructures after heat treatments 3 and 4 were identical except for grain size, and refining the grain size eliminated the transgranular fracture associated with a Charpy impact energy of 26J and permitted the material to achieve the conditions necessary for the ductile intergranular fracture which resulted in a Charpy impact energy of 83J. If the grain boundary sulfides could be eliminated, the toughness of the grain refined structure should be higher than 83J but whether it would increase to the 167J observed for heat treatment 1 cannot be determined. It is believed that these results show that the low toughness observed for heat treatment 3 is, significantly, the result of the large grain size. The volume fraction of MnS particles which form on the grain boundaries established at 1260°C is very small. The voids on the intergranular fracture surfaces produced by heat treatment 4 are spaced about 1 I.tm apart and the particles, when visible, have a diameter of about 0.1 larn. The relationship Sv = 2/L, where Sv is the grain boundary area per unit volume and L is the linear intercept grain size (12), suggests the volume fraction of particles at these grain boundaries is about 0.000002. This is considerably less than the inclusion volume fraction observed for heat treatments 3 and 4; thus it is believed that the inclusion volume fractions observed for heat treatments 3 and 4 reflect intra-grain sulfide precipitation during the 15 hour hold at 750°C. Conclusion The results indicate that high austenitizing temperatures can enhance the Charpy impact energy and such heat treatments are associated with increases in grain size and dissolution of MnS particles. Such experiments do not permit an evaluation of the effect of grain size. The purpose of this work was to assess the influence of grain size on the upper shelf Charpy impact energy in the absence of other microstructural changes such as inclusion volume fraction and spacing. The results indicate that increasing grain size can have a significant and detrimental influence on the upper shelf Charpy impact energy. References 1. G.G. Chell and D.A. Curry, "Mechanics and Mechanisms of Ductile Fracture", in Developments in Fracture Mechanics, Vol. 2, G.G. Chell, Ed., p. 101, Applied Science Publishers (1981). 2. K.J. Handerhan and W.M. Garrison, Jr., Metall. Trans. 19A, 2989 (1988). 3. S. Lee, L. Ma~,no and J.R. Asaro, Metall. Trans. 16A, 1633 (1984). 4. V.F. Zackay, E.R. Parker and W.E. Wood, Conf. Proc. Third ICSMA, Institute of Metals, London, Vol. 2, p. 175-179 (1973). 5. R.O. Ritchie and J.F. Knott, Metall. Trans. 5, 782 (1974). 6. R.O. Ritchie and R.M. Horn, Metall. Trans. 9A, 331 (1978). 7. C.W. Marschall, R.F. Hehemann and A.R. Troiano, Trans. ASM 55, 135 (1962). 8. E.E. Underwood, Ouantitative Stereoloev, p. 81, Addison Wesley (1970). 9. E.E. Underwood, Ouantitative Stereology, p. 82, Addison Wesley (1970). 10. R.N. O'Brien, D.H. Jack and J. Nutting, Effects of MnS distributions on the fracture properties of heattreated low-alloy steels, from Heat Treatment '76, Proc. of 16th Int. Heat Treatment Conference, The Metals Society, London, p. 161 (1977). 11. T.J. Baker and R. Johnson, J. Iron Steel Inst., 211,783 (1973). 12. J.E. Hilliard, Stcreolo~y: Principles and Practices, Northwestern University (1981). 13. C.W. Corti, P. Cotterill, A. Fitzpatric, Int. Metals Rev. 19, 77 (1979).
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Vol. 27, pp. 881-854, 1992 Printed in the U.S.A.
Pergamon Press Ltd. All rights reserved
T H E I N F L U E N C E O F G R A I N S I Z E ON T H E UPPER SHELF IMPACT E N E R G Y OF A 9Ni S T E E L A.L. Wojcieszynski W.M. Garrison, Jr. A.W. Thompson Dept. of Metallurgical Engineering and Materials Science Carnegie Mellon University Pittsburgh, PA 15213 USA
(Received March 27, 1992) (Revised July 21, 1992) Introduction Experimental results indicate that large grain sizes will reduce the cleavage fracture stress and increase the ductile to brittle transition temperature of martensitic, bainitic and ferritic steels (1). However, there does not appear to be a similar consensus on the effects of grain size on upper shelf toughness. In part this situation has arisen because grain size is typically increased in steels by increasing the austenitizing temperature, and higher austenitizing temperatures can reduce the volume fractions and alter the spacings of void initiating particles such as carbides and sulfide inclusions (2,3). Therefore, while increasing the austenitizing temperature can, in many instances, result in improved upper shelf toughness (4-6), it is very difficult to extract from such results any effect of grain size on toughness due to possible changes in particle dispersions. To unambiguously assess the role of grain size and toughness, it would be desirable to vary the grain size without influencing the inclusion volume fraction and spacing and without complications due to undissolved carbides. The purpose of this paper is to report the results of experiments which suggest that increasing grain size can have a significant and detrimental effect on upper shelf toughness, as measured here by the Charpy impact energy. Materials, Heat Treatments and Procedures The material used in this work was a 0.07wt.%C/9wt.%Ni/0.2wt.%Cr steel (7). This steel was selected because of the low carbon and chromium content, which would minimize the formation of carbides, because it is known to have reasonably high toughness in the as-quenched condition after austenitizing at 790°C and because it remains wholly austenitic to temperatures as low as 700°C. This latter characteristic is useful in maintaining inclusion volume fraction and spacing while changing grain size. The material was produced by vacuum induction melting followed by vacuum arc remelting. The ingot was subsequently upset and cross forged to a bar 5 cm thick and 8.9 cm wide. The chemistry of the material is given in Table 1. Four heat treatments were used. Heat treatment 1 consisted of air cooling after 1 h at 900°C followed by austenitizing at 790°17 for 1 h and a water quench; this heat treatment results in a relatively small prior austenite grain size. Heat treatment 2 consisted of air cooling after 1 h at 900°C followed by 1 h at 1260°C and a water quench. This heat treatment was used to determine what effects reducing inclusion volume fraction coupled with a large grain size would have on toughness. Heat treatment 3 consisted of several steps. The material was air cooled after 1 h at 900°C; the material was then held at 1260°C for 1 h, quenched directly into a salt pot at 750°C, held for 15 h, and then water quenched. This heat treatment results in the large prior austenite grain size of heat treatment 2, but a larger inclusion volume fraction. Heat treatment 4 was identical to heat treatment 3 except the material was given a final grain refining treatment consisting of 1 h at 750°C followed by a water quench.
851 0956-716X/92 $8.00 + .00 Copyright (c) 1992 Pergamon Press
Ltd.
852
IMPACT ENERGY
Vol.
27, No.
Table 1 Alloy Chemistry*
C
Ni
Cr
Mn
Mo
0.077
9.29
0.26
0.37
0.01
Si 0.008
S
P
A1
02**
N2**
0.002
0.003
0.005
7
4
* in wt%; ** in wt ppm Table 2 Mechanical Properties Condition
H.T. H.T. H.T. H.T.
1 2 3 4
oy (MPa)
UTS (MPa)
ef
Cv (J)
822 819 828 850
1128 1090 1095 1149
1.29 1.22 0.74 1.17
167 249 26 83
ay - yield strength; UTS - ultimate tensile strength; e f - the strain to fracture; Cv = Charpy impact energy
Table 3 Microstructural Data Condition
H.T. H.T. H.T. H.T.
1 2 3 4
f
0.00027 0.00004 0.00016 0.00014
Ro
XO
(lain)
(~tm)
0.119 0.095 0.113 0.115
1.59 2.41 1.96 1.98
grain size (btm) 13.3 151 151 13.8
f - inclusion volume fraction; R0 - average inclusion radius calculated from R0 = (x/4)H(d) where H(d) is the harmonic mean of apparent diameters on polished cross-sections (12); X0 - inclusion spacing calculated from X0 = 0.89 R0f-1/3 (13).
The properties measured were the room temperature smooth axisymmetric tensile properties and the Charpy impact energy. The tensile specimens had gage diameters and lengths of 6.35 mm and 25.4 mm respectively and were tested at a crosshead speed of 0.5 mm/min. Grain sizes were obtained according to the linear intercept method (8). Inclusion volume fractions and average radii were determined from 5000X SEM back scatter images of polished cross sections. Results and Discussion The mechanical properties measured for the four heat treatments are listed in Table 2. The four heat treatments result in very similar strength levels. However, the strains to fracture and Charpy impact energies were considerably lower after heat treatment 3 than after the other three heat treatments. Microstructural information is
7
Vol.
27, No.
FIG. 1
7
IMPACT ENERGY
Fracture Surfaces of Charpy impact specimens of material produced by (a) heat treatment 1, (b) heat treatment 2, (c) heat treatment 3 and (d) heat treatment 4.
summarized in Table 3. The measured inclusion volume fraction for material receiving heat treatment 1 was 0.00027 while the inclusion volume fraction after heat treatment 2 was much lower, 0.00004; this decrease in inclusion volume fraction was due to dissolution of MnS at the austenitizing temperature of 1260°C. After heat treatment 3 the inclusion volume fraction increased to 0.00016, due to precipitation of MnS particles during the hold at 750°C for 15 h. Heat treatment 4 refined the grain size achieved by heat treatment 3, and the inclusion volume fraction after heat treatment 4 was essentially the same as after heat treatment 3. Our primary interest was to compare the toughnesses measured for heat treatments 1 and 3, which had Charpy impact energies of 167J and 26J, respectively. The fracture surfaces associated with the two heat treatments were microvoid coalescence (Fig. 1), with no evidence of overheating effects on the fracture surface of heat treatment 3. The fracture surfaces are very similar in terms of void size and spacing in some regions but there are regions on the fracture surface of heat treatment 1 where the voids are much larger than observed for heat treatment 3; these regions appear to be associated with the walls of the zig-zag fracture path. Taken at face value the data indicate that the low toughness of heat treatment 3 compared to heat treatment 1 is due to the larger grain size. However, the inclusion volume fraction measured for heat treatment 3 was only half that measured for heat treatment 1. That the inclusion volume fraction measured after heat treatment 3 was only 0.00016 could stem from two causes. First, it is possible that not all of the sulfur has precipitated during the hold at 750°C. Second, it is possible that after heat treatment 3 there are small sulfides not visible on polished cross sections at 5000X; if the latter is true, it is possible that all of the sulfur has reprecipitated during the hold at 750°C and that the inclusion volume fraction is the same as for heat treatment 1 but the inclusion size is smaller than for heat treatment 1. If the latter possibility is true, the low impact energy of heat treatment 3 could be due to a much smaller inclusion spacing than calculated and not to the large grain size.
853
854
IMPACT ENERGY
Vol.
27, No.
To refine the grain size obtained after heat treatment 3, specimens receiving heat treatment 3 were austenitized at 750°C; this resulted in a grain size of 13.8 larn and an increase in the Charpy impact energy from 26J to 83J. The microstructural data in Table 3 indicate that reaustenitizing at 750°C had little effect on the sulfide distribution. The fracture mode after heat treatment 4 was not transgranular but was a ductile intergranular fracture following the grain boundaries established at 1260°C, and the void nucleating particles on these grain boundaries were MnS particles; this fracture mode is typical of overheating and the improved toughness of heat treatment 4 is consistent with studies of overheating (9,10) which show that the ductile intergranular fracture mode is, at least for lower strength steels, not associated with particularly low Charpy impact energies. That the ductile intergranular fracture mode is observed only after the grain-refining heat treatment is also consistent with studies of overheating which suggest that without the grain refining step there is a competition between the transgranular and grain boundary fracture paths, which is influenced by the ratio of the grain boundary and matrix inclusion spacings, with large ratios favoring the transgranular fracture mode (11). After heat treatment 3 the matrix spacing was apparently too small to permit the intergranular fracture. The microstructures after heat treatments 3 and 4 were identical except for grain size, and refining the grain size eliminated the transgranular fracture associated with a Charpy impact energy of 26J and permitted the material to achieve the conditions necessary for the ductile intergranular fracture which resulted in a Charpy impact energy of 83J. If the grain boundary sulfides could be eliminated, the toughness of the grain refined structure should be higher than 83J but whether it would increase to the 167J observed for heat treatment 1 cannot be determined. It is believed that these results show that the low toughness observed for heat treatment 3 is, significantly, the result of the large grain size. The volume fraction of MnS particles which form on the grain boundaries established at 1260°C is very small. The voids on the intergranular fracture surfaces produced by heat treatment 4 are spaced about 1 I.tm apart and the particles, when visible, have a diameter of about 0.1 larn. The relationship Sv = 2/L, where Sv is the grain boundary area per unit volume and L is the linear intercept grain size (12), suggests the volume fraction of particles at these grain boundaries is about 0.000002. This is considerably less than the inclusion volume fraction observed for heat treatments 3 and 4; thus it is believed that the inclusion volume fractions observed for heat treatments 3 and 4 reflect intra-grain sulfide precipitation during the 15 hour hold at 750°C. Conclusion The results indicate that high austenitizing temperatures can enhance the Charpy impact energy and such heat treatments are associated with increases in grain size and dissolution of MnS particles. Such experiments do not permit an evaluation of the effect of grain size. The purpose of this work was to assess the influence of grain size on the upper shelf Charpy impact energy in the absence of other microstructural changes such as inclusion volume fraction and spacing. The results indicate that increasing grain size can have a significant and detrimental influence on the upper shelf Charpy impact energy. References 1. G.G. Chell and D.A. Curry, "Mechanics and Mechanisms of Ductile Fracture", in Developments in Fracture Mechanics, Vol. 2, G.G. Chell, Ed., p. 101, Applied Science Publishers (1981). 2. K.J. Handerhan and W.M. Garrison, Jr., Metall. Trans. 19A, 2989 (1988). 3. S. Lee, L. Ma~,no and J.R. Asaro, Metall. Trans. 16A, 1633 (1984). 4. V.F. Zackay, E.R. Parker and W.E. Wood, Conf. Proc. Third ICSMA, Institute of Metals, London, Vol. 2, p. 175-179 (1973). 5. R.O. Ritchie and J.F. Knott, Metall. Trans. 5, 782 (1974). 6. R.O. Ritchie and R.M. Horn, Metall. Trans. 9A, 331 (1978). 7. C.W. Marschall, R.F. Hehemann and A.R. Troiano, Trans. ASM 55, 135 (1962). 8. E.E. Underwood, Ouantitative Stereoloev, p. 81, Addison Wesley (1970). 9. E.E. Underwood, Ouantitative Stereology, p. 82, Addison Wesley (1970). 10. R.N. O'Brien, D.H. Jack and J. Nutting, Effects of MnS distributions on the fracture properties of heattreated low-alloy steels, from Heat Treatment '76, Proc. of 16th Int. Heat Treatment Conference, The Metals Society, London, p. 161 (1977). 11. T.J. Baker and R. Johnson, J. Iron Steel Inst., 211,783 (1973). 12. J.E. Hilliard, Stcreolo~y: Principles and Practices, Northwestern University (1981). 13. C.W. Corti, P. Cotterill, A. Fitzpatric, Int. Metals Rev. 19, 77 (1979).
7