Ultrahigh-strength low-alloy steels with enhanced fracture toughness

Progrrn

m Materral~ Scrmcr Vol. 42. pp 209-242, 1997 Elsevier Science Ltd. All rights reserved Prmted ,n Great Bntam 0079-6425197 532 00

‘R 1997

Pergamon

PII: SOO79-6425(97)00016-9

ULTRAHIGH-STRENGTH LOW-ALLOY STEELS WITH ENHANCED FRACTURE TOUGHNESS G. Malakondaiah”, M. Srinitlas* and P. Rama Raot *Defence Metallurgical Research Laboratory, Hyderabad, 500 058, India and tJawaharla1 Nerhu Centre for Advanced Scientific Research, Jakkur, Bangalore, 560 064, India CONTENTS 1. INTRODUCTION 1.1. Maraging Steels: Strengthening Through Intermetallic Precipitates 1.2. Secondary Hardening Steels: Strengthenrng Through Fine Alloy Carbides 1.3. Low-alloy Steels: Strengthening Through Transttion Carbides 1.4. Approaches to Achieve Enhanced Fracture Toughness in Lo&r-alloy Steels 1.4.1. Retarned austenite 1.4.1 1. Retained austenite: through high-temperature austenitising (HTA) 1.4.1 2. Retained austenite: through alloymg addrttons 1.4.2.Bainrte-martensite mixed microstructures 1.4.3. Control of sulfide inclusions 2. OUR APPROACH 2.1. Infruence of Alloying Elements on Strength and Toughness of Iron 2.1.1. Tensile data 2.1.2. Fracture toughness data 2.1.3. Modeling fracture toughness data 2.2. Fe-CX Alloys 3. DEVELOPMENT OF NiSiCrCoMo STEEL 3.1. Cobalt-modified NiSiCr Steels 3.2. Comparison with Other Ultrahigh-strength Steels 3.3. Reproducibility of Properties of NiSiCrCoMo Steel on Tonnage Scale 3.4. Machinabrhty and Formability 3.5. Weldability ACKNOWLEDGEMENTS REFERENCES

treatment

209 212 212 213 214 214 214 215 216 216 217 218 218 222 223 225 227 229 234 236 239 239 240 ‘40

1. INTRODUCTION Structural steels capable of developing a minimum yield strength of 1380 MPa (200 ksi) are generally referred to as ultrahigh-strength (UHS) steels. Broadly, there are three classes of ultrahigh-strength steel: (a) high-alloy maraging steels typified by 18Ni(250) steel, (b) high-alloy secondary hardening steels typified by AF1410 and (c) low-alloy steels typified by AISI 4340. The focus of this paper is on low-alloy UHS steels and on the attempts made to enhance their fracture toughness. Following a brief discussion of microstructural features and their influence on the fracture toughness of the three aforementioned types of steel, we shall describe the work carried out in the authors’ laboratory on the development of an UHS NiSiCrCoMo low-alloy steel. 209

210

Progress in Materials Science

Maraging steels are essentially carbon-free martensitic steels which employ substitutional elements to achieve age hardening. Intermetallic compounds [Ni,Mo, N&Ti, Fe,(Mo,Ti)] formed during ageing of essentially carbon-free Fe-Ni martensite impart ultrahigh strength to maraging steels. Secondary hardening steels are a class of quenched and tempered steels which contain sufficient amount of carbide-forming elements (chromium, molybdenum, vanadium). Upon tempering of low-carbon Fe-Ni martensite at temperatures between 773 and 873 K, the coarse cementite particles are replaced by a fine dispersion of the more stable alloy carbides. In contrast to the normal softening that occurs at these tempering temperatures, a fine dispersion of carbides results in hardening-leading to the term ‘secondary hardening’. Secondary hardening as well as maraging steels contain high alloying content. On the other hand, the medium-carbon low-alloy martensitic steels attain ultrahigh strength in a stage-I (low-temperature) tempered condition, with transition (iron) carbides. Here, the alloying additions are such as mainly to impart hardenability while contributing to strength through solid-solution hardening. Steels for several high-technology applications, such as aircraft and aerospace, need to possess ultrahigh strength coupled with high fracture toughness in order to meet the requirement of minimum weight while ensuring high reliability. Yet another property not discussed here but one that needs to be addressed, as brought out in the excellent review by Olson,(‘) is resistance to stress corrosion cracking. Enhancements of fracture toughness, at a yield strength of 1500-1700 MPa, achieved in the three classes of UHS steels under consideration are summarised in Fig. 1 from which it is clear that secondary hardening as well as maraging steels possess high fracture toughness at ultrahigh strength levels (Table 1 may be referred to for the compositions of all steels discussed). On the other hand, the accompanying toughness

HARDENING STEELS

AF1410

I20

Aer Met 100

18Ni (250) 1,(1700MPa)

.

LOWALLOYSTEELS lY60

1970

1980

19YO

2000

YEAR Fig. I. Enhancement of fracture toughness in ultrahigh-strength steels. High-alloy secondary hardening as well as high-alloy maraging steels possess high fracture toughness at ultrahigh strength levels. Significant progress has been made during the past three-and-a-half decades to enhance fracture toughness of low-alloy steels.

0.34

0.49

0.10 max

001

max

0.12 max

0.03 max

Mn

O&O.8 0.65-0.9 0.6-0.9 O.OSSO.2 0.1 max 0.1 max 0.12 max

0 38-0.43 0.4CkO.46 0.42-0.48 0.10-O 14 0.134.17 0.21-0.27 0.03 max

Si

(wt%)

2.89

I 88

18-19

0.4-0.7 9 S-IO.5 9.5-10.5 I I-12 17-19

3 5-4.0

4.6-5.2

MO 0.2-0.3 0.3-0.4s 0.9-1.1 09-1.1 0.9-l 1 1 o-l.3 4.65. I

Ni

steels considered

1.65-2.0 1.65-2.0

17-18

I .02

0.7-0.9 0.7-0.9s 0.9-I .2 I .8-2.2 I .8-2.2 2.5-3.3

Cr

of ultrahigh-strength

0 10 max

0.12 max

0.12 max

0.2-0.35 1.45-1.80 O.lSSO.3 0 I max 0.1 max

1. Compositlons

4340 300M D6ac HY 180 AF1410 AerMetlOO 18Ni(250) Maraging 18Ni(300) Maragmg 18Ni(350) Maraging Garrison NiSiCr steel

Table

C

Designation

_~

_

o.osLO 1 0.05-O. 1

V

m this article

12-13

8-9 5

7.5-8.5 13.5-14.5 13 3-13.5 7-8.5

co

0.3-0.5 Ti 0.054 15 Al 0 5-0.8 Ti 0.05-O. 15 Al 1.6-2.0 TI 0.14 2 Al

Other

212

Progress in Materials Science

levels in the case of UHS low-alloy steels are relatively low. Low-alloy steels offer advantages of availability and low cost, as compared with high-alloy steels such as maraging steels. With enhanced fracture toughness, matching that of maraging steels, low-alloy steels can become an excellent option for high-technology aerospace applications.

1.l . Maraging Steels: Strengthening Through Intermetallic Precipitates The superior toughness of maraging steels compared with low-alloy steels (like AISI 4340) is attributablec2’ to the finer size of intermetallic precipitates, with increased resistance to fracture, as compared with carbides in quenched and tempered low-alloy steels. Fracture in either case occurs by void nucleation, growth and coalescence. Voids nucleate by the fracture of Ti(C,N) inclusions in maraging steels and at the interfaces between MnS inclusions and the matrix in AISI 4340 steel. However, the critical difference is with regard to void growth and coalescence. Fracture of coarse carbide particles causes void sheets that aid linkage of large voids in low-alloy steels while in maraging steels primary voids growing until impingement cause coalescence and final fracture. Of the three popular ultrahigh-strength grades of 18Ni maraging steels, namely 250, 300 and 350 grades, we confine ourselves to 250 grade (Table 1) since the emphasis with the other two grades, developed during 1961 to 1965, was on increased strength.‘” Development of cobalt-free grades with properties comparable to the standard cobalt-bearing gradest4’ and modified heat treatment cycles for ring forgingse) have been the recent developments with regard to 18Ni(250) maraging steel. A modified heat treatment cycle has been developed(5) since the solution-annealing treatment specified for 18Ni(250) maraging steelc6)was found to be inadequate to achieve the required fracture toughness (K,,) of 90 MPa rn’12in heavy forgings with non-uniform cross-section. The modified heat treatment involves subjecting the forgings, prior to the standard treatment, to a high-temperature solution-annealing treatment of 1223 K to dissolve carbides and then quenching rapidly to prevent carbide precipitation at the grain boundaries. The pretreatment results in the formation of fine recrystallised grains and increases fracture toughness by up to 40% (Fig. 1). For heavier forgings, a three-stage solution annealing (1223 K + 1223 K + 1093 K) was necessary to develop the required fracture toughness. Auger electron microscopy performed by Misra et al.(‘) showed that the variation in K,, with heat treatment temperature correlates with variation in area1 density of Ti,CS/TiC precipitates. Higher heat treatment temperatures (> 1243 K) cause dissolution of carbides, leading to higher toughness.

1.2. Secondary Hardening Steels: Strengthening Through Fine Alloy Carbides AFl410 and its derivatives form the class of secondary hardening steels. AFl410 is a higher-strength modification of HY180, with increased carbon and cobalt, developed in the early 1970s. The chemical composition of AF1410 is included in Table 1. On tempering at 783 K, following oil quenching from austenitising temperature, AF1410 develops@,9, a yield strength of 1545 MPa and K,, of 154 MPa m’12.The high toughness is possible because the alloy carbides, in a manner similar to intermetallics in the case

UHS Low-alloy Steels and Enhanced Fracture Toughness

213

of maraging steels, are too fine to nucleate voids. As shown in Fig. 1, lanthanum addition further improves fracture toughness. When the sulfur is gettered as lanthanum oxysulfides and the inclusion volume fraction is about 0.003, AF1410 develops”’ “I K,, of 190 MPa m”‘. AerMet 100”” is the latest generation from the family of Ni-Co-Cr secondary hardening steels that combines high strength (yield strength = 1720 MPa), high fracture toughness (& = 126 MPa m”2; Fig. 1) and exceptional resistance to stress corrosion cracking (&cc = 38.5 MPa ml!‘). Multi-step ageing, previously developed”3’ for AF1410, results”4’ in a further increase in strength of AerMet 100 (hardness increases from R, = 53.5 to Rc = 54.5) while maintaining high fracture toughness. The multi-step ageing treatment consists of an initial high-temperature short-time austenite precipitation treatment followed by a lower-temperature step. The beneficial effect of multi-step ageing treatment was attributed to dispersed-phase transformation toughening arising from precipitated austenite. 1.3. Low-alloy

Steels: Strengthening

Through

Transition

Carbides

In a recent review, Krauss”” provided a detailed account of microstructure, deformation and fracture of low-temperature-tempered martensitic steels. The fine structure of lath martensite that forms in medium-carbon low-alloy steels, characterised by relatively high MS (martensite start) temperatures, consists of high densities of tangled dislocations. Fine intralath transition carbides form during tempering at low temperatures (stage I, up to 523 K), with the morphologies of as-quenched martensite crystals remaining unchanged. [Jack ‘Ih)first identified, by X-ray diffraction, the structure of the transition carbides to be hexagonal and termed the carbide c-carbide (Fe2,C). Subsequently, Hirotsu and Nagakura “‘.“I have identified, by electron diffraction. the structure as orthorhombic and termed the carbide n-carbide (FeC).] Because of the fine transition carbides, martensitic steels tempered in stage I maintain high strength. At intermediate tempering temperatures (stage II; 523-673 K), interlath-retained austenite transforms to carbides, leading to tempered martensite embrittlement with an accompanying decrease in strength. With further increase in tempering temperature (stage III; 523-973 K), transition carbides and retained austenite are replaced by mixtures of ferrite and cementite leading to increased toughness; however, at the cost of strength. To achieve ultrahigh strength levels in medium-carbon low-alloy steels it is, therefore, essential to restrict to stage-I (low-temperature) tempering. AISI 4340, 300M and D6ac (Table 1) are UHS low-alloy steels which have received much attention. Alloy 300M is a silicon-modified (1.6% Si) 4340 steel. D6ac is a low-alloy ultrahigh-strength steel developed for aircraft and missile structural applications and develops high fracture toughness’” when subjected to aus-bay (798 K) quench, following austenitisation treatment at 1198 K. Aus-bay quench results in K,, values of 99 to 104 MPa m’ ’ at a tensile strength of 1650 MPa. Considerable research efforts have been directed, during about the last three decades or so, to enhance the fracture toughness of AISI 4340 and 300M low-alloy steels (Fig. 1). Tomita has reviewed”” the methods employed by several investigators for improving the strength/toughness combination in medium-carbon low-alloy steels. These efforts can be classified microstructurally in terms of strategies based on (1) retained austenite, (2) mixed microstructures and (3) control of non-metallic inclusions.

Progress in Materials Science

214

1.4. Approaches to Achieve Enhanced Fracture Toughness in Low-alloy Steels During the 1960s and 1970s newer steelmaking practices were developed with the primary aim of reducing the oxygen and sulfur contents. These include vacuum arc remelting (VAR) and electroslag refining (ESR). VAR results in a large reduction in the contents of hydrogen and oxygen, and also nitrogen to some extent, while ESR reduces sulfide and oxide inclusions. Application of these steel-refining techniques during production have resulted in minimising directionality in properties and noticeable, but not adequate, improvements in fracture toughness at ultrahigh strength levels,(9) as summarised in Fig. 1. For example, vacuum arc refining of AISI 4340 causes an increase in K,, from 45 MPa m’j2 for the air-melted (AM) condition to around 60 MPa rn”l at a yield strength of 1650 MPa. Incorporation of ductile retained austenite as thin interlath films in fully martensite as well as bainite-martensite mixed microstructures, partial transformation in the bainitic range to develop bainite-martensite mixed microstructures and inclusion shape control through calcium or Si + Ni additions have successfully been employed to achieve further improvements in fracture toughness of ultrahigh-strength low-alloy steels. 1.4.1. Retained austenite

The presence of retained austenite as thin interlath films improves toughness through its crack-arresting ability. Cracks propagating through martensite would be stopped upon intersecting a region of retained austenite. With further loading, the cracks branch out and grow around the austenite area. This manner of crack motion would necessarily involve more energy absorption than straight propagation through martensite plates, leading to improved toughness. 1.4.1.1. Retained austenite: through high-temperature austenitising (HTA) treatment. A higher austenitising temperature of 1473 K results in an increase of up to 90% in &, in the as-quenched condition, as compared with the conventional austenitising treatment at 1143 K (38 MPa ml’* to 72 MPa m’j2) with no loss in strength (1517 MPa yield strength for 1473 K treatment compared with 1551 MPa for 1143 K treatment). The increased fracture toughness of AISI 4340 steel associated with the high-temperature austenitising (HTA) treatment was attributed (*‘I to the presence of interlath-retained austenite. However, the increase in K,, is accompanied by an appreciable decrease in Charpy impact energy. To arrive at an explanation for the discrepancy between Charpy and fracture toughness data, Ritchie and co-workers studied the influence of austenitising temperature on the toughness of AISI 4340 steel in the as-quenched(22) as well as the quenched and tempered(23’ condition. In line with the previous reports, HTA treatment at 1473 K was found to double the K,, (sharp crack) of AISI 4340 steel with a concomitant decrease in Charpy V-notch (rounded notch) impact energy. The increase in K,, after hightemperature (1473 K) austenitisation in 473 K tempered condition(23) is particularly significant as K,, increased from 56 MPa m”* to 92 MPa ml’* as against the enhancement from 38 MPa m”* to 72 MPa m “* for the as-quenched condition.‘**) The marked enhancement in J&, as compared with the conventional 1143 K austenitising treatment, was seen in both the conditions even though fracture is by ductile rupture in the quenched and tempered condition and by cleavage/intergranular mechanism in the as-quenched condition. However, they found no evidence to support the suggestion(2’*24)that the

UHS Low-alloy Steels and Enhanced Fracture Toughness

215

enhancement of fracture toughness at higher austenitising temperatures was due to the presence of films of retained austenite. The amount of retained austenite was observed to be nearly the same (5 to 6%) when the austenitising temperature was increased from 1143 to 1473 K. Further, the retained austenite was so mechanically unstable in the as-quenched condition that less than 1.5% remained untransformed at yield in tensile tests. In a recent study, Haidemenopoulos et al. (13’studied the stability of retained austenite as a function of stress state. The retained austenite, although sufficiently stable under a state of pure shear stress, was unstable under the crack-tip stress state. Ritchie et al.“‘.‘” attributed the discrepancy between Charpy and fracture toughness data to a differing response of the microstructure produced by each austenitising treatment to the notch root radius. Charpy specimens contain a V-notch (root radius, p = 0.25 mm) whereas K,, test specimens contain a fatigue precrack (p-0). Failure by stress-controlled quasi-cleavage/intergranular mechanism (or by strain-controlled ductile rupture mechanism) occurs when the maximum tensile stress (or strain) exceeds the critical fracture stress, cF (or critical fracture strain, c,), over a characteristic distance ahead of the notch tip. Stress-controlled fracture predominates in the as-quenched condition for which the decrease in Charpy V-notch impact energy with increasing austenitising temperature was related to a reduction in oF, whereas the increase in K,, was attributed to an increase in the characteristic distance, through a coarsening of microstructure. On the other hand, for strain-controlled fracture, which dominates in the quenched and tempered condition, the decrease in Charpy V-notch impact energy with increasing austenitising temperature has been attributed to a decrease in ductility or eF while the increase in K,, was associated with an increase in the characteristic distance, apparently brought about by dissolution of void-nucleating particles at high austenitising temperatures. 1.4.1.2. Retained austenite: through alloying additions. To establish the role of retained austenite on mechanical behaviour, with special emphasis on fracture toughness, Narasimha Rao and Thomas (25)studied quaternary alloy (manganese or nickel) additions to Fe-C-Cr base steel. Both manganese and nickel are fee austenite stabilisers and promote retention of austenite. It was observed (25’ that addition of 2 wt% manganese to FeeO.3C4Cr results in 4 to 5% retained austenite while the 5 wt% nickel-modified alloy has 6 to 8% retained austenite. A significant improvement in K,, at a given yield strength was obtained in either case. Addition of 2% manganese increased K,, from 85 to 140 MPa m”’ in the 473 K tempered condition at a yield strength of 1300 MPa. The beneficial effect was attributed to the increased amount and stability of retained austenite. The somewhat inferior toughness properties of the 5% nickel-modified alloy compared with the manganese-modified alloy were attributed ‘z5’to the small but significant fraction (about 10%) of twinning in the former case. Retained austenite also imparts significant improvements to fracture toughness in isothermally transformed, high-silicon-bearing, experimental low-alloy steels FeeO.2C- 2Si--3Mn and Fe-O.4C-2Si4Ni (““)) as well as silicon-modified 4330 ((?“) and 4340 (300M) steels. (‘* -. 29)The presence of a high silicon content in steels that have been isothermally transformed in the bainitic temperature region, encourages retention of ductile high-carbon austenite, rather than the formation of brittle interlath cementite films which have a detrimental effect on both ductility and toughness. 300M steel (composition in Table 1), subjected to partial transformation at 593 K followed by oil

216

Progress in Materials Science

quenching and subsequent tempering at 473 K, showed(28.29)a 25% increase in K,, (68 MPa m “* to 85 MPa m”‘; Fig. l), at matching ultimate tensile strength of 1950-1990 MPa, over the value for the conventional quench-tempered condition. However, the yield strength decreased from 1600 to 1420 MPa. The improvements in ductility and toughness were attributed (28,29) to the thin-film form of mechanically stable austenite and the bainitic ferrite, whose individual ferrite plates are separated by the thin films of austenite, while the increased strength was attributed to the refinement of martensite by the bainite. To summarise, retained austenite through HTA treatment increases K,, by a factor of about two. The marked increase in K,, resulting from HTA is, however, accompanied by a decrease in Charpy V-notch impact energy. It is not therefore advantageous to resort to HTA treatment for applications where K,, and impact energy are equally important. On the other hand, the relatively stable retained austenite introduced through alloy modifications improves K,, significantly without affecting Charpy impact energy. 1.4.2. Bainite-martensite

mixed microstructures

Bainite-martensite mixed microstructures (MM), developed through short-term isothermal treatments, can significantly improve mechanical properties if the ductile second phase appears in a suitable morphology (size, shape and distribution). In the case of 4340 steel, one steel that has been studied extensively with regard to mechanical behaviour of mixed microstructures, it has been reported (3s32)that lower bainite in association with tempered martensite provides a better combination of strength and toughness compared with fully martensitic structures. lO-25% by volume of lower bainite in 4340 steel causes an improvement in K,, from 54 MPa m ‘I2 for the fully martensitic condition to 78 MPa m’12.(3’)On the other hand, the presence of upper bainite was found to be detrimental and results in the appearance of a remarkably elevated fracture transition temperature (a measure of toughness), irrespective of the volume fraction.(30) Lower bainite associated with martensite appears in acicular form and partitions prior austenite grains of the matrix martensite leading to microstructural refinement, while upper bainite associated with martensite appears as masses and fills prior austenite grains of matrix martensite. Contrary to the generally reported behaviour, a vanadium-bearing AISI 4330V steelt3” treated to form upper bainite in a mixed microstructure developed significant improvement in toughness without affecting the strength of fully martensitic structure. Further, no beneficial effect of lower bainite on mechanical properties was observed in the lower bainite-martensite mixed microstructure. The improvement in fracture toughness was to the extent of about 12% in the AISI 4330V steel (81.4 MPa m’12 compared with 72.7 MPa ml/* in the fully martensitic condition at the same yield strength level of 1410 MPa). Short-term isothermal treatments, although beneficial from the point of view of toughness, appear to bear limited industrial significance since steels are not commonly processed under isothermal conditions. 1.4.3. Control of sulfide inclusions Rice and Johnson(34) have shown that the upper-shelf fracture toughness of ultrahighstrength steels is sensitive to inclusion spacing as well as inclusion volume fraction. Cox and Lowos) have demonstrated that reducing the inclusion content increases the

UHS Low-alloy Steels and Enhanced Fracture Toughness

217

resistance to void nucleation and is thereby effective in improving the fracture toughness of 4340 steel. Leslie(36’has also suggested that reducing the sulfur content might be a means of achieving enhanced fracture toughness of UHS steels. In a recent study, Tomitaf3” observed that desulfurisation treatment alone does not always result in improved fracture toughness of 4340 steel. In a separate study, Tomita’-‘“’ also demonstrated that addition of calcium to the desulfurised steel melt is effective in improving the fracture toughness. Calcium feeding (S/Ca = 3) using Ca-Si wires (Ca:Si = 60:30) modified the morphology of the inclusions from stringer MnS to finely dispersed, spherical CaS particles. In comparison with the commercial 4340 steel, an improvement in K,, of 25 MPa ml’* (50 MPa m”’ to 75 MPa ml’?) in the L-T orientation and 30 MPa m’ ’ (39 MPa m”’ to 69 MPa ml’*) in the T-L orientation was observed (Fig. I), at a yield strength of 1600 MPa, in desulfurised (at 0.002 wt% S level) 4340 steel. However, the calcium treatment was found to be not so effective in modifying the morphology of sulfide inclusions in commercial 4340 steel containing higher sulfur, such as 0.016 wt% S. In this situation, calcium treatment raised the fracture toughness of commercial 4340 steel only marginally from 50 MPa ml’* to 55 MPa m’ ‘. A change in morphology of MnS inclusions can also be achieved through a decrement in hot-rolling reduction. (39’Decreasing the hot-rolling reduction from 98% to 80% modified the shape of MnS inclusions in 4340 steel from stringer (aspect ratio = 17.5) to ellipsoid (aspect ratio = 3.8) and the K,, improved by - 20 MPa m”’ without affecting the strength properties. Garrison@‘(‘)achieved control of sulfide size through combined additions of nickel and silicon. He studied the effect of separate as well as combined additions of silicon and nickel on the fracture toughness behaviour of 0.4C-1.5Ni-l.OCr base steel. Separate additions of silicon and nickel did not increase the fracture toughness of the base steel significantly. However, these additions, when made in combination, had a marked effect on toughness. Combined additions of silicon (2 wt%) and nickel (1.5 wt%) increased K,, from 75 to 115 MPa m” at 1650 MPa yield strength. The improved fracture toughness of base + Ni + Si steel was attributed to the increased sulfide spacing resulting from an increase in average sulfide size. The average sulfide radius for the base + Ni + Si steel was 0.71 urn, almost three times the average size measured for base, base + Ni or base + Si steels. Control of sulfide inclusions (CSI) appears to be an attractive means to achieve improvements in K,,. However, for this strategy to be more effective, the steel has got to be cleaner. 2. OUR APPROACH Although numerous ultrahigh-strength low-alloy steels are currently utilised in engineering applications requiring good toughness, relatively little systematic information exists on the effects of alloying elements on their fracture toughness properties. In the case of low-alloy martensitic steels, certain principles of alloy design have been utilised. In general, optimum properties are obtained by using the lowest possible carbon content consistent with the desired strength level. Solid-solution elements are added in various quantities to develop adequate strength and hardenability, with nickel considered necessary to improve toughness and molybdenum to minimise reversible temper embrittlement. Although these basic guidelines are available, no consistent set of

Progress in Materials Science

218

conclusions has been formulated which determine the effect of alloying elements on the fracture toughness. A comprehensive research and development programme was launched to achieve improvements in fracture toughness of low-alloy steels through modifications in alloy chemistry. As part of this programme, the fracture toughness behaviour of iron and its alloys has been studied. Evaluation has been made of fracture behaviour of iron-based binary solid solutions with commonly used alloying elements, namely chromium, cobalt, molybdenum, nickel or silicon as solutes. To establish the influence of alloying additions in the presence of carbon, these studies were extended to Fe-C-X alloys. The toughening effect of cobalt, as revealed by the studies on iron-based solid solutions, was found to be even more pronounced in the presence of carbon. With the understanding thus developed, cobalt additions were made to achieve further improvements in fracture toughness of a NiSiCr steel made and evaluated by Garrison. (40)The work has resulted in the development of a NiSiCrCoMo low-alloy steel possessing a strength/toughness combination quite comparable to that of the high-alloyed 18Ni(250) grade maraging steel. We present, in the following sections, a detailed account of the development of NiSiCrCoMo steel including the technological aspects. 2.1. Influence of Alloying Elements on Strength and Toughness of Iron To study systematically the influence of alloying elements on the fracture behaviour of iron, five commonly used alloying elements in steel, namely cobalt, chromium, nickel, molybdenum and silicon, were chosen. Solid-solution alloys of iron with chromium, cobalt, molybdenum or nickel at 0.5 and 5.0 wt% and silicon at 0.5 and 3.5 wt% concentrations were vacuum induction melted in 80 kg lots using commercial-purity Armco iron and electrolytically pure chromium, cobalt, molybdenum, nickel or silicon. The chemical composition of the binary alloys is presented in Table 2. 2.1.1. Tensile data As a prelude to examining the influence of these solute elements on the fracture toughness of iron, tensile tests were conducted and the data are included in Table 2. It is evident from Fig. 2 that cobalt and chromium cause solid-solution softening while molybdenum, nickel and silicon strengthen iron. Among the elements studied, silicon has Table 2. Chemical composition, tensile and fracture toughness properties of iron binary alloys

Material Armco iron Fe-OSSi Fe-3.5Si Fe-O.SMo Fe-5.OMo Fe-O.SNi Fe-5.ONi Fe-O.SCr Fe-5.OCr Fe-O.SCo Fe-5.OCo

Solute concentration (wt%) 0.58 3.30 0.45 4.40 0.48 4.70 0.42 4.65 0.48 4.80

Grain size (crm)

Yield stress (MPa)

UTS (MPa)

n

(kJ m-*)

118 125 140 125 120 125 80 115 135 130 125

180 225 423 187 212 195 257 118 139 110 130

296 385 541 359 385 335 386 196 262 290 293

0.28 0.21 0.15 0.27 0.22 0.25 0.21 0.28 0.285 0.30 0.35

140 91 42 112 102 59 80 120 145 162 187

XC

UHS Low-alloy Steels and Enhanced Fracture Toughness

219

Si

( ~ 3:. -------__ -_a-_--_--_-.

ke

Cr co

U

1

2

3

4

5

6

ALLOYING ELERIENT CONTENT,Wt.% Fig. 2. Change in yield stress of iron as a function of alloying element content. Cobalt and chromium cause solid-solution softening while molybdenum, nickel or sdicon strengthen iron.

a pronounced hardening effect. The increase in yield strength of iron by molybdenum, nickel or silicon observed in the present study (Table 2 and Fig. 2) is in agreement with that reported earlier.14’43) Cobalt lowers the yield strength of iron, with no observable effect on ultimate tensile strength (UTS); Table 2. On the other hand, addition of chromium lowered the yield strength as well as the UTS of iron. This difference in the effect of cobalt and chromium on the strength properties of iron is reflected in the work-hardening exponent, n, of Fe-Co and Fe-Cr binary alloys (Fig. 6). The alloy softening in both these cases is more marked in dilute binaries (Fig. 2). A reduction in yield strength of iron with cobalt(43.44’and chromium““) additions have been reported in the literature with explanations based on intrinsic as well as extrinsic mechanisms. The intrinsic mechanism pertains to enhanced double-kink nucleation on screw dislocations at dissolved alloying atoms,‘4’.JS46)whereas the extrinsic ‘scavenging’ mechanism is related to the suggestion that the addition of alloying elements decreases the solubility of interstitials.‘43 44” “) Cobalt increases the activity of carbon in ferrite and results in a 75% decrease in the solubility of carbon in Fe-5Co.“‘) Through secondary-ion mass spectroscopy (SIMS), experimental evidence has been obtained for the suggestion that the ‘scavenging’ of interstitials leads to alloy softening. Ion images for Armco iron, Figs 3(a) and (b), indicate that the interstitial carbon is homogeneously distributed in the solution. On the other hand, C and CN- ion images for Fe-SC0 alloy, Figs 3(c) and (d), suggest segregation of carbon and nitrogen at the grain boundaries. To substantiate SIMS observations, the Hall-Petch constants were determined from yield strength data obtained over a wide range of grain size for iron and four binary alloys, namely FeeO.SCo, Fe-SCo, Fe-5Cr and Fe-3.5Si. The scavenging mechanism implies removal of carbon from the grain interior and its segregation to the grain boundaries. If this is operative, it should be seen from the Hall-Petch constants, namely (T,,(a measure of friction stress resisting dislocation movement within the grain) and k, (a measure of

220

Progress in Materials Science

50 urn Fig. 3. SIMS images showing (a,b) uniform distribution of interstitials in Armco iron, and (c,d) scavenging of interstitials to grain boundaries in Fe-SC0 alloy.

grain boundary strength). Variation of yield stress with grain size for Armco iron, Fe-Co alloys, Fe-5Cr and Fe-3.5Si alloys is presented in Fig. 4. The grain size dependence of yield stress (a,) can be expressed as cr, = 119 + 25.4d-“2

(Armco iron)

(1)

(Fe-O.SCo)

(2)

gy = 58 + 37.0d-“2

(Fe-5Co)

(3)

oy = 50 + 26.0d-‘I2

(Fe-5Cr)

(4)

(Fe-3.5Si)

(5)

gY= 39 + 36.8d-“2

6, = 366 + 25.0d-“*

where q is in MPa and d is in mm. Expressions (1) to (3) suggest that cobalt addition causes a significant reduction in cr,,whereas k, increases from 25.4 to 37.0 MPa mm’12. As k, is a known measure of grain boundary strength, the increase in k, with cobalt

UHS Low-alloy

Steels and Enhanced

Fracture

Toughness

221

3505 200 Fe -5.0 Cr 1.50 P

. z

100 -

ii s n LI s 200 -

150 -

100 -

50

0

1

2

(GRAIN DIAMETER

3

4

, d j”2, IIIII?‘~

Fig. 4 HallLPetch plots (grain SE dependence of yield stress) for Armco Iron. Fe4 5Co. FeeSCo, Fe-5Cr and Fe-3.56 alloys. Increase m k,,a measure of grain boundary strength, m Fe-Co alloys substantl,ates SIMS observations (Fig. 3) suggestmg segregation of mterstktlal atoms at the gram boundarles

addition suggests segregation of interstitial atoms at the grain boundaries. Preliminary studies with limited grain size data in the case of Fe-5Cr alloy point to a negligible increase in k, from 25.4 to 26.0 MPa mm’ ’ and therefore no carbon segregation at the grain boundaries, Since chromium is a carbide former, if (Cr,Fe),C, forms, the equivalent amount of carbon in solution would be diminished and solid-solution softening due to chromium addition is again attributable to removal of carbon from ferrite as reported in softening occurs in this case without enhancement the literature.‘” ” 54’ Solid-solution of the grain boundary strength. On the other hand, silicon, a known strong solid-solution strengthener, causes, at a concentration level of 3.5 wt%. a threefold increase in CJ,,. There is no nnfluence of silicon addition on k).

222

Progress in Materials Science

1

0

2

3

4

5

6

ALLOYING ELEMENT CONTENT,wt. % Fig. 5. Influence of solute additions on fracture toughness of Armco iron at a constant grain size of 125 pm. Of the five solutes studied, cobalt causes significant improvement to fracture toughness of Armco iron, chromium at 5% causes no change while nickel, molybdenum or silicon has a deleterious effect of varying degree on toughness.

2.1.2. Fracture toughness data Armco iron, Fe-O.SCr, Fe-SCr, Fe-O.SCo, Fe-SC0 and Fe-O.SMo exhibited stable crack extension while FeO.SNi, Fe-SNi, Fe-O.SSi, Fe-3.5Si and Fe-5Mo showed cleavage instability during loading of precracked fracture toughness specimens.@” The J-R curve technique as well as critical stretch zone width were employed to derive JIC in the former case while in the latter case the value of J at the onset of cleavage instability was taken as JQ and subjected to validity checks. Fracture toughness JIGdata for the 10 binary alloys are presented in Table 2 and plotted as a function of alloying content in Fig. 5. The notable observation is that cobalt imparts to Armco iron a significant improvement in fracture toughness. On the other hand, molybdenum, silicon and nickel have an increasing degree of deleterious effect on Armco iron in that order, with the least effect being noticed with chromium.(55-57) The large decrease in JIGwith silicon and higher molybdenum concentration is explained on the basis of a change in fracture mode from ductile to cleavage as a result of stress concentration ahead of the crack tip reaching the cleavage fracture stress. On the other hand, the loss in fracture toughness with nickel addition is attributed to sulfur segregation at grain boundaries which results in pockets of intergranular fracture. The increase in JIG of Armco iron as a consequence of cobalt addition is primarily attributable to an increase in the strain-hardening exponent, n (Fig. 6). The critical plastic zone size increases with increasing n. Furthermore, increase in II results in slip dispersal and reduces the void growth rate. Increased n as a result of cobalt addition is therefore expected to delay void nucleation and decrease the void growth rate. It thus appears that cobalt addition toughens iron by influencing the energy spent in formation of the plastic zone as well as in the subsequent stages of the ductile fracture process, namely void nucleation, growth and coalescence of ductile fracture. If the lowering of yield strength by cobalt addition to iron were to be the prime reason for the enhancement of JIC, addition

UHS Low-alloy Steels and Enhanced Fracture Toughness

u

1

ALLOYING

223

2

3 4 5 6 ELEMENT CONTENT, wt.%

Fig. 6 Variation of strain-hardening exponent n of Iron with alloying element content. Of the five solutes studled, cobalt increases n of Armco iron while moybdenum, nickel and silicon lower n. On the other hand, chromium appears to have no observable effect on n of iron.

of 0.5% cobalt, which reduced the yield strength by 40% as compared with a 25% decrease with 5% cobalt addition, should have resulted in a larger increase in Jlc. Contrary to these expectations, a 5% cobalt addition caused a higher increase in JIc of iron (35%). Although Fe-Co alloys have lower yield strengths than Armco iron, the flow stress builds up rapidly as a result of higher strain-hardening rate at the early stages of plastic deformation. 21.3. Modeling ,fracture toughness data A new approach has been proposed (j8) for the calculation of Jlc in terms of the critical strain criterion model, originally developed by Rice and Johnson’34’ and Ritchie and Thompson,“” for engineering alloys having a substantial volume fraction of second-phase particles. The sequence of events leading to crack initiation is assumed to be as follows. As soon as the precracked sample is loaded in tension under a monotonically increasing load, extensive blunting of the crack tip occurs. The crack blunting process is completed, leading to a constant stretch zone width (SZW,), before the void nucleation and growth occur at a characteristic distance 1, from the blunted crack tip (Fig. 7). The mechanism of void nucleation ahead of the blunted crack and also the characteristic distance Z, of the void from the crack tip have been determined through a careful examination of the interrupted fracture toughness test specimens of Armco iron, a single-phase material. (@IExperimental evidence was obtained for the nucleation of voids at slip-band intersection and at sites where slip-band impingement on grain boundaries occurred (Fig. 8). To determine the void site ahead of the crack tip, the compact tension toughness test of Armco iron (38 urn grain size) was interrupted at regular displacement intervals, samples mid-sectioned, polished, etched and examined under a scanning electron microscope (SEM). A SEM micrograph corresponding to the displacement level of 2.9 mm is shown in Fig. 9. At this displacement the crack tip

224

Progress in Materials Science

Fig. 7. A schematic

illustration

of the plastic zone ahead of a blunted zones.

crack tip delineating

various

has blunted extensively and developed a nearly semicircular profile. Figure 9 reveals the formation of voids ahead of the blunted crack tip. The distance (k) between the first parallel void array and the crack tip is 125 urn. The SEM micrograph corresponding to the displacement level of 3.3 mm (Fig. 10) reveals that, when compared with Fig. 9, the crack tip coalesced with the voids ahead of it. It is to be noted that the crack front progresses by linking the parallel array of voids progressively leading to the reblunting of the crack tip. Once again parallel void arrays are seen. Measurements of 1, for varied grain size show that I, is a microstructure-dependent parameter and decreases with coarsening of grain size (Fig. 11). The proposed approach not only incorporates the experimentally measured 1, values, but also plastic-flow-related energy dissipation in

Fig.

8. SEM micrograph showing the nucleation of voids by slip-band impingement: intersection of slip bands (A) and intersection with the grain boundary (B).

mutual

UHS Low-alloy Steels and Enhanced Fracture Toughness

225

Fig. 9. SEM micrograph of Armco iron of 38 pm grain size corresponding to a displacement level of 2.9 mm in an interrupted fracture toughness test. Extensive blunting of the crack tip to a near-semicircular profile and void nucleation ahead of it at a characteristic distance 1, are seen.

the Hutchinson-Rice-Rosengren expression

zone (Fig. 7) beyond the process zone leading to the

J,, = {CK&/(n + l)} {[2.3 exp(-2.351,/Q x exp(-1.5o,/a)y+’

+ 3.57 ln(bl,/a,)

+ [6,,/(1,(2.36,,/1,- l)n”+ ‘) exp( -2.351,/Q]“+

‘)

(6)

where C is a numerical constant, KH the strain-hardening coefficient in the relationship g = KHt”, d,, the crack-tip opening displacement, bl, the final radius of the void where b is substantially lower than 1, a, the initial stage of void and a,/o the normalised tensile hydrostatic stress. Theoretically predicted J,c values are compared in Fig. 12 with those derived experimentally. Figure 12 also includes data on Armco iron and nickel derived over the temperature range 298-673 K.(56) It is clear from Fig. 12 that the proposed method estimates reasonably well (to within ) 15%) the ductile initiation fracture toughness, Jlc, of single-phase materials exhibiting a wide range of values (100-300 kJ m-*). 2.2. Fe-C-X

Alloys

The studies on Fe-C-X alloys have shown that the beneficial effect of cobalt on & of iron is even more pronounced in the presence of carbon.@‘) The strength and toughness properties of Fe-O.2C, FeeO.2C-5Co and Fe42C-5Ni alloys are compared in Table 3. The microstructural features, namely ferrite grain size and the volume fraction

226

Progress in Materials Science

Fig. 10. SEM micrograph of Armco iron of 38 pm grain size corresponding to a displacement level of 3.3 mm in an interrupted fracture toughness test. Coalescence of the crack tip with an array of grown voids and a parallel array of voids ahead of it are evident.

of pearlite, were maintained nearly the same in all the three Fe-C-based alloys under consideration. A 5% cobalt addition does not influence significantly the strength behaviour of Fe-O.2C alloy. On the other hand, a 5% nickel addition increases the strength appreciably and reduces the ductility and work-hardening exponent. The addition of 5% cobalt significantly enhances (80%) & of Fe-O.2C alloy, while the addition of 5% nickel increases the same marginally (18%). Nickel, generally considered to be a beneficial element in the presence of carbon, is seen to influence the impact toughness dramatically with only a marginal effect on &. On the other hand, cobalt addition, with no observable effect on impact toughness, causes significant improvement in &. These results further suggest that there is no one-to-one correlation between impact toughness and static fracture toughness for this class of alloys. Similar findings have been reported on steels in recent years(22,23,62) although there is a fairly widespread use of correlations between impact toughness and fracture toughness.

UHS Low-alloy

Fig. 11. Variation

Steels and Enhanced

Fracture

Toughness

of characteristic distance 1, with grain size. lC is microstructure-dependent decreases with coarsening of grain size.

The reasons for the dramatic effect of nickel on impact toughness observed study could be traced to alloy softening at the high strain rates prevailing conditions,(63) as also suggested by Johey’+” and Petch.‘6”

3. DEVELOPMENT

OF NiSiCrCoMo

227

and

in the present under impact

STEEL

It is evident from the foregoing that cobalt toughens iron, and even more positively in the presence of carbon. In our efforts to arrive at a new alloy chemistry that possesses a strength/fracture toughness combination matching that of high-alloy maraging 250 steel, cobalt addition has been resorted to. To obtain ultrahigh strength levels, carbon content in the range 0.3-0.4% is needed’@ even when tempering is confined to stage I. Such proportions also ensure lath martensitic structure. Nickel, silicon and chromium were chosen as the other alloying additions, keeping in view the beneficial effects of these elements. Chromium has a powerful effect on hardenability with no

Progress in Materials Science

228

C = 0.34 ( z )'“(anin

80

pm)

280

180 J,, (model),

380

kJ/m’

Fig. 12. A comparison of experimentally measured Jr< values with those predtcted by the model. The predicted Jr< value is within f 15% of the experimental value for a number of single-phase materials covering a wide range of measured J,, values.

adverse effect on strength and toughness of iron (see Section 2.1). To achieve required hardenability, as also suggested by Pickering,‘““’ its content is chosen at 1.0 wt%. As discussed earlier, silicon causes pronounced solid-solution strengthening of ferrite. Silicon is also known to strongly inhibit the growth of transition carbides and the formation of Fe,C during tempering, allowing the steel to be tempered at relatively higher temperatures (up to 588 K) while avoiding embrittlement. In other words, silicon shifts the tempered martensite embrittlement (TME) to higher temperatures. These beneficial effects of high silicon content have profitably been made use of in the development of 300M steel. Nickel is known to increase the resistance to cleavage fracture of iron and, as also revealed in the present study, causes a significant reduction in ductile-to-brittle transition temperature (DBTT). Further, Garrison’s study revealed that combined additions of nickel and silicon influence the size and distribution of sulfide inclusions and thereby the fracture toughness. Keeping in view these observations, the

Table

Maternal Fe-0.2C Fe-O.ZCSCo Fe-O.ZC-5Ni

3. Tensile,

Impact

and fracture

toughness

properties

Yield stress (MPa)

UTS (MPa)

Elongation (25 mm gauge length)

244 242 297

370 409 467

44 47 37

of Fe--C--X alloys Charpy V-notch impact energy

(J)

JIG (kJ mm’)

I6 15 120

130 232 153

UHS Low-alloy Steels and Enhanced Fracture Toughness

229

contents of nickel and silicon were chosen as per Garrison’s chemistry at 3 and 2 wt%, respectively. 3.1, Cobalt-modljied

NiSiCr

Steels

To arrive at an optimum treatment, the tempering behaviour of the base NiSiCr steel was established.‘b7,68’The variation of strength with tempering temperature (Fig. 13) shows that the NiSiCr steel retains high strength up to 523 K. Beyond about 573 K, there occurs a steep fall in strength, which is also accompanied by a loss in ductility. A trough in ductility variation with tempering temperature occurs in the temperature range 673 to 873 K. The impact energy at room temperature increases with tempering temperature up to around 573 K, beyond which a steep fall occurs, with a minimum at 773 K (Fig. 14). It is evident from the data presented in Figs 13 and 14 that the steel develops an optimum combination of strength and impact toughness in the 523 K tempered condition. Accordingly, the fracture toughness and ductile-to-brittle transition temperature were evaluated for the steel in the 523 K tempered condition. Fracture toughness tests conducted according to ASTM Standard E399 yielded K,, of 103 MPa m”2. The impact toughness in the 523 K tempered condition is plotted against test temperature in Fig. 15. The 20 J criterion yields a DBTT of 173 K.

UTS

RA Et

I I lo 300

I I 700 500 Tempering temperature (K)

I 900

Fig. 13. Tensile properties of NiSiCr steel as a function of tempering temperature. To retain ultrahigh strength levels it is essential to temper at temperatures <573 K.

230

Progress in Materials Science o NiSiCr steel l NiSiCrCo steel

011

300

I

I

500 700 Tempering temperature (K)

I

900

Fig. 14. Charpy V-notch impact energy as a function of tempering temperature for NiSiCr and NiSiCrCo steels. The impact energy in both the steels increases with tempering temperature up to around 573 K, beyond which a steep fall occurs.

Mechanical properties of the base NiSiCr steel, conforming to Garrison steel in chemistry, subjected to an optimum treatment of 523 K tempering are summarised in Table 4. The data presented in Table 4 indicate that the strength/toughness properties reported by Garrison could not be fully reproduced. The NiSiCr steel studied here possesses K,,of 103 KPa m”* at a yield strength of 1550 MPa in the optimum heat-treated condition as against K,, of 115 MPa ml’* at 1682 MPa yield strength reported by Garrison.@@ Transmission electron microscopy (TEM) and Auger electron spectroscopy (AES) suggest’67.6sJ that the loss in ductility and impact toughness of NiSiCr steel in the tempering temperature range 623 to 873 K is due to the simultaneous occurrence of tempered martensite embrittlement (TME) and tempered embrittlement (TE). It is now well established that TME occurs in low-alloy steels when tempered in the temperature range 523 to 673 K. A higher temperature range of 623 to 873 K for NiSiCr steel is understandable as silicon is known to shift TME to higher temperatures. TEM carried out on specimens tempered at 523, 773 and 873 K revealed the presence of interlath-retained austenite in the 523 K tempered condition, transformation of retained austenite to carbides in the 773 K tempered condition and coarse interlath carbides in the 873 K tempered condition (Fig. 16). Further, AES carried out on 773 K tempered condition showed segregation of phosphorus, which was absent in the as-quenched condition, pointing to the occurrence of temper embrittlement. Cobalt addition was made to the base NiSiCr steel to achieve further improvements in fracture toughness. The tempering behaviour of NiSiCrCo steel (Fig. 17) suggests that the steel retains higher strength up to 573 K, as against 523 K observed with NiSiCr steel. Ductility variation follows a trough in the tempering temperature range from 673

UHS Low-alloy Steels and Enhanced Fracture Toughness

o NiSiCr steel l NiSiCrCo steel 523 K tempered

231

. .

123 K

1 I

il73K I 200 Temperature (K)

0

Fig. 15. Variation of Charpy V-notch impact energy with test temperature for NiSiCr and NiSiCrCo steels in 523 K tempered condition. Cobalt-bearing steel shows higher impact energy at all test temperatures and the 20 J criterion yields a DBTT of 123 K for NiSiCrCo steel as against 173 K for NiSiCr steel.

to 873 K. The impact energy data included in Fig. 14 indicate that the impact toughness increases with tempering temperature up to 523 K, beyond which a steep fall occurs. It is evident from Figs 14,17 that NiSiCrCo steel, in a manner similar to that observed with the base NiSiCr steel, develops an optimum combination of strength and impact toughness in the 523 K tempered condition. Fracture toughness tests carried out in this optimum condition yielded a K,, of 140 MPa ml’*. At the same time, the impact properties are improved with cobalt addition to the base steel (Table 4 and Fig. 15). The 20 J criterion yields a DBTT of 123 K for NiSiCrCo steel as against 173 K for NiSiCr steel (Fig. 15). Mechanical properties of NiSiCrCo steel in the 523 K tempered condition are included in Table 4. Table 4. Mechanical properties of NiSiCr-based steels in the optimum heat-treated condrtron

Material NiSiCr NiSiCrCo NiSiCrMo NiSiCrCoMo

Yield stress (MPa) 1550 1360 1648 1530

UTS (MPa) 1825 1670 1904 1890

Reduction in area (%) 47 50 48 48

Charpy V-notch Impact energy at RT

KC

(J)

(MPa m’ 2,

36 52 35 34

103 140 102 120

232

Progress in Materials Science

Fig. 16. TEM micrographs of NiSiCr steel tempered at (a) 523 K and (b) 773 K showing in terlath retained austcmite and carbide stringers, respectively. (c) Coarse interlath carbicIes are seen in the 873 K tempered condition.

UHS Low-alloy Steels and Enhanced Fracture Toughness

Ni Si Cr Co steel

233

o UTS l TS ARA

60 .RA 50 Et 40

30 20 I

5 300

Fig

17 Tensile properties

of NiSiCrCo retains higher

I 500

I 700

Temperature

(K)

I 900

steel as a function of tempering strength up to about 573 K.

temperature.

The steel

The increase in fracture toughness with an accompanying decrement in strength with cobalt addition to base steel is in line with the observations made(55’on cobalt additions to Armco iron. Microsegregation studies carried out on NiSiCrCo steel using a scanning Auger microscope point to segregation of carbon to the grain boundaries (Fig. 18). The increase in fracture toughness with cobalt addition is attributable to the segregation of carbon to the grain boundaries, thus increasing the cohesive strength of the grain boundary and rendering the crack initiation process much more difficult.(‘,69’ To build up strength in the presence of cobalt in the base NiSiCr steel, molybdenum addition was resorted to. The studies on iron-based solid solutions have shown that, while molybdenum is a solid-solution strengthener, the fracture toughness of Armco iron is only marginally lowered as a result of molybdenum addition, below a certain level of concentration. The influence of molybdenum on the base NiSiCr steel was therefore studied. The addition of molybdenum to the base NiSiCr steel leads to an increase in yield and ultimate tensile strength with negligible reduction in fracture toughness (Table 4). The increase in yield strength is as much as 7% while the increase in UTS is around 5%. As a follow-up, molybdenum addition was made to the cobalt-containing steel. To optimise the composition further, melts with varying contents of cobalt and molybdenum were taken and processed. The fracture toughness of the optimised NiSiCrCoMo steel is nearly 15% higher than that of the base NiSiCr steel at a matching strength level (Table 4). SIMS studies carried out on NiSiCrCoMo steel in the optimised condition pointed to

Progress in Materials Science

234

Ni-Si-Cr-Co steel

Fe

0

100

200

300

400 500 600 Electron energy (eV)

700

800

900

I 00

Fig. 18. Scanning Auger spectrum showing carbon segregation at grain boundaries in NiSiCrCo steel.

carbon segregation to grain boundaries (Fig. 19), reinforcing the suggestion made from the results on binary Fe-Co alloys. The microstructure in the heat-treated conditions of cobalt- and/or molybdenummodified NiSiCr steels comprised lightly tempered lath martensite, with interlath films of retained austenite. The retained austenite content varied between 2 and 4%. Fine intralath transition carbides are seen at higher magnifications in all the cases. A typical high-magnification transmission electron micrograph of NiSiCrCo steel in the 573 K tempered condition is shown in Fig. 20. The carbides appear to grow as rods or wavy needles. Cobalt- and/or molybdenum-modified NiSiCr steels are characterised by MS temperatures”O) of 627-630 K and the lath martensitic structure is in line with the microstructural predictions (‘I) based on Ms. Likewise, fracture in all four cases occurred by microvoid coalescence. The fracture zone just ahead of the stretch zone of fracture toughness specimens tested in heat-treated conditions was examined under an ISI IOOA scanning electron microscope. Microvoids are seen to form in association with fractured particles and/or interfaces (Fig. 21). These observations with regard to microstructure and fractography of NiSiCr alloys with cobalt, molybdenum or Co + MO additions are similar to those reported by Garrison”‘) for his NiSiCr base steel. 3.2. Comparison with other Uitmhigh-strength

Steels

The strength/toughness properties of NiSiCrCoMo steel are compared in Fig. 22 with those of the low-alloy steels AISI 4340, 300 M and D6ac, and highly alloyed maraging steels. Figure 22 shows the range of properties attainable with low-alloy quenched and tempered (standard heat treatment) steels and 18Ni maraging steels as

UHS Low-alloy Steels and Enhanced Fracture Toughness

235

Fig. 19. SIMS image suggesting carbon segregation to grain boundaries in NiSiCrCoMo steel.

two separate bands. For comparison, the best reported strength/toughness data for AISI 4340, 300 M and D6ac are included. In the case of AISI 4340 steel, the higher value of K,, (92 MPa m”*) has not been considered since the beneficial effect of HTA, as described earlier, is accompanied by a decrease in impact energy. The K,, values attainable through short-term isothermal treatment (78 MPa rn’!‘) and through calcium treatment (75 MPa m’:2) being nearly the same, the latter value is chosen keeping in view the practical limitations of the former treatment. For the same reason, with regard to 300M steel, the strength (yield strength = 1600 MPa) and toughness (K,, = 68 MPa ml’*) data corresponding to the conventional treatment are selected although retained austenite-bainite-martensite mixed microstructure yields K,, of 85 MPa m’12. The K,c value of 102 MPa ml’*chosen for D6ac corresponds to the high fracture toughness aus-bay quench treatment. The data on NiSiCrCoMo steel fall over a range as they cover values obtained on laboratory- as well as industrial-scale melts. It is evident from Fig. 22 that the modified treatments have shifted the strength/toughness combination of AISI 4340, 300M and D6ac steels to the regime of maraging steels. On the other hand, the strength/toughness combination of NiSiCrCoMo steel (designated as DMR-S 250), attainable in the commonly used quenched and tempered condition, is better than the

236

Progress in Materials Science

Fig. 20. TEM micrograph of NiSiCrCo steel showing intralath transition carbides.

best reported values for low-alloy steels AISI 4340, 300M and D6ac and falls in the upper-bound range for high-alloy 18Ni(250) maraging steel. NiSiCrCoMo steel possesses fatigue crack growth resistance comparable to that of 250-grade maraging steel. The fatigue crack growth data on NiSiCrCoMo steel are presented as a function of stress intensity range (AK) and compared with those for 250-grade maraging steel in Fig. 23. The data in either case could be described by Paris power-law of the form da/dN = C(AQm, where C and m are material constants that are determined experimentally: da/dN = 4.8 x 10m’3(AK)4,0 NiSiCrCoMo

steel

da/dN = 4.1 x 10-‘2(AK)3.25 maraging steel

(7) (8)

where da/dN is in m cycle-’ and AK is in MPa ml/‘. 3.3. Reproducibility of Properties of NiSiCrCoMo

Steel on Tonnage Scale

Three 5 ton capacity melts of NiSiCrCoMo steel (DMR-S 250) were taken and processed by Mishra Dhatu Nigam Limited (MIDHANI), a special alloy undertaking at Hyderabad. The mechanical properties of DMR-S 250 processed on this scale are compared in Table 5 with those derived from a laboratory-scale melt. The attractive

UHS Low-alloy Steels and Enhanced Fracture Toughness

10pm Fig. 21. Typical SEM fractographs of fracture toughness specimens showing ductile dimpled fracture. Microvoids form in association with fractured particles (a) and/or interfaces (b).

237

Progress in Materials Science

238 200

* .

DMR-S2SO AlSI4340

MARAGWC

0

1000

2000

3&U

YIELD STRENGTH, MPa Fig. 22. A comparison of fracture toughness/yield strength data for NiSiCrCoMo steel with those for UHS low-alloy steels and high-alloy maraging steels. The strength/toughness combination of NiSiCrCoMo steel, in the commonly used quenched and tempered condition, is better than the best reported values for low-alloy steels AISI 4340, 300M and DBac, and falls in the upper-bound range for 18Ni(250) maraging steel.

I

AK

x Maraging steel

(MPa mrn)

Fig. 23. Fatigue crack growth data for NiSiCiCoMo steel and 18Ni(250) maraging steel. NiSiCrCoMo steel possesses fatigue crack growth resistance comparable to that of 18Ni(250) maraging steel.

UHS Low-alloy Table

Steels and Enhanced

5. Mechamcal

properties

Property

Fracture

of NlSlCrCoMo Lab-scale

Yield strength (MPa) Ultimate tenslle strength (MPd) Elongation (25 mm gauge length) (%) Charpy V-notched impact energy at RT (J) K,, (MPd m”) ,‘The range is a result of properties hot-formed products

combination reproduced

obtained

Toughness steel Industrial-scale

melt

of strength and toughness observed in this industrial-scale melt.

3.4. Machinuhilit_v

melt

1610 I970 136 2x--3?” 90-l lo”

1530 I x90 14 34 120 at various

239

stages of processmg

in the laboratory

of the Ingot to

melt is seen to have been

and Formahilit~

Spheroidise-annealing softens the NiSiCrCoMo steel considerably. The properties of the steel in the spheroidise-annealed condition are, again, quite comparable to those of the 250-grade maraging steel in the solutionised condition (Table 6). NiSiCrCoMo steel in the softened condition can easily be rough-machined before final subjection to hardening and tempering treatment in order to develop an optimum combination of strength and fracture toughness. The steel in the softened condition possesses good formability. The Erichsen cupping test yielded a cup depth of 10.8 mm for 2.3 mm sheet. In conformity with these observations, a 7 mm thick blank of 570 mm diameter could be formed. without intermediate annealing, to a shell of 350 mm inner radius.

3.5.

Weldubilit?

The gas tungsten-arc (GTA) process, generally recommended for low-alloy steels, has been chosen. Detailed studies were undertaken to arrive at a suitable filler material. Tension test results of butt-welded 2.3 mm thick sheet specimens (Table 7) with W2 filler (trade name of MIDHANI, Hyderabad for their filler developed for 250-grade maraging steel) and heat-treated yields 96% weld efficiency in terms of yield strength (Table 7). Further, the base metal strength could be fully realised in the butt-welded and heat-treated condition with base steel as filler. A weld efficiency of 90% (in terms of yield strength) could be achieved (Table 7), in experiments on 4.5 mm thick sheets, with the base steel as filler, even in the heat-treated and welded condition. Attempts are currently underway to achieve further improvements in weld efficiency in the heat-treated and welded condition. The weldments also possess good fracture toughness. A 4.5 mm thick weldment, with base steel as filler, has a fracture toughness (plane stress, Kc) of

Table

6. A comparison

of mechanical

Steel NiSlCrCoMo (spheroldlse-annealed) 18N1(250) maraging (solution annealed)

properties

m the softened

Yield strength (Mt’d) 652 725

UTS (MPa) x53 IO00

condition Elongation (%) 19.4 18.0

Progress in Materials Science

240

Table 7. Tensile properties of weldments of NiSiCrCoMo steel

Prior condition Welded + heat-treated W2 filler Base filler Heat-treated + welded Base filler NiSiCrCoMo base steel

UTS (MPa)

Elongation (50 mm gauge length) (%)

1555 1644

1847 2002

3.2 4.4

1458 1622

1576 2013

1.3 9.5

Yield strength (MPa)

110 MPa rn’12in the welded and heat-treated condition compared with the base value of 130 MPa m’12. To summarise, the NiSiCrCoMo steel (DMR-S 250) discussed in this presentation promises to be an inexpensive, and therefore attractive, substitute for the more expensive, high-alloy steel like the well-known 18Ni(250) maraging steel for lightweight, highperformance structural applications. ACKNOWLEDGEMENTS The authors are very grateful to Mr S. L. N. Acharyulu, Director, DMRL for encouragement and permission to publish this work. The authors express sincere gratitude to Mr J. Marthanda Murthy for his dedication and unflinching involvement in the development work of the NiSiCrCoMo steel reported here. Thanks are due to Dr R. D. K. Misra and Mr T. V. Balasubramanian for their help in conducting Auger studies and Mrs Vydehi Arun Joshi, Dr K. Muraleedharan and Dr Tapan Roy for TEM studies. Dr R. D. K. Misra helped in obtaining SIMS pictures. We are deeply grateful to Mr K. K. Sinha and Mr R. K. Mahapatra, former and present CMDs of MIDHANI, respectively, and their colleagues for their leadership in processing the tonnage-scale melts of the steel. We thank Larsen and Toubro Limited, Mumbai for weldability and formability studies. Thanks are also due to Dr T. Mohandas for supply of weldments. Dr Vikas Kumar helped with measurements of fatigue and fracture toughness properties of weldments. Grateful appreciation is accorded to Professor W. M. Garrison Jr; our meeting and subsequent correspondence with him has provided the predominant impetus to our work. REFERENCES 1. G. B. Olson, 2. 3.

4. 5. 6.

7. 8.

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Toughness

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