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The influence of thermomechanical processing on the smooth fatigue properties of Ti–15V–3Cr–3Al–3Sn
Materials Science and Engineering A243 (1998) 97 – 102
The influence of thermomechanical processing on the smooth fatigue properties of Ti–15V–3Cr–3Al–3Sn R.R. Boyer a,*, H.J. Rack b, V. Venkatesh b a
Boeing Commercial Airplane Group, PO Box 3707, MS 6H-CJ, Seattle, WA 98124, USA b Clemson Uni6ersity, Clemson, SC 29634, USA
Abstract This study has examined the potential for enhancing the fatigue performance of Ti – 15V – 3Cr – 3Al – 3Sn sheet through variations in heat treatment and thermomechanical processing procedures. Processing variants included conventional solution treatment and aging, solution treatment with duplex- and triplex-ages, and cold and warm rolling followed by aging. Tensile and smooth fatigue performance of each processing route were evaluated and compared to ‘standard’ commercial practice. These processing avenues provided a wide range of microstructures with tensile strengths ranging from 1070 to 1610 MPa. Enhancements in fatigue performance were dependent on processing and stress level. At high stresses, i.e. within the low cycle regime, a 15% improvement in the maximum stress for a fixed 30 000 cycle lifetime was achieved. At lower stresses, i.e. within the high cycle fatigue ‘limit’ regime, a 50% improvement relative to the standard heat treatment was achieved. A good correlation of fatigue and tensile strengths was observed; generally, the smooth fatigue properties improved as the tensile strength increased. © 1998 Elsevier Science S.A. All rights reserved. Keywords: Ti – 15V – 3Cr–3Al–3Sn; Thermomechanical processing; Smooth fatigue properties
1. Introduction ‘Standard’ commercial practice for Ti – 15V–3Cr– 3Al–3Sn (Ti–15– 3) thin sheet typically involves cold rolling, solution treatment at 790°C for 5 min. and aging at 540°C for 8 h. This procedure results in a typical tensile strength of 1035 MPa and a 106 cycle run-out stress of 700 MPa [1], which is inferior to that observed in a + b titanium alloys tested at equivalent strengths. Extensive examinations have analyzed the potential for enhancing the strength and fracture toughness of Ti– 15 – 3 through modifications in the alloy thermomechanical processing procedures [2–5]. For example, Inaba et al. [2] showed that significant strength enhancement can be achieved by cold rolling followed by single step aging. They reported that, under constant aging conditions, strength increased with increasing amounts of cold work. Further, an improvement of ductility was also achieved at 80% deformation, where a microduplex equiaxed a/b struc-
* Corresponding author. Tel.: +1 206 9652463; fax: + 1 206 9651440; e-mail: [email protected] 0921-5093/98/$19.00 © 1998 Elsevier Science S.A. All rights reserved. PII S 0 9 2 1 - 5 0 9 3 ( 9 7 ) 0 0 7 8 5 - 5
ture is formed upon aging. This latter enhancement was shown to be associated with recovery of the deformation substructure prior to a precipitation. With lesser degrees of deformation, direct lamellar or acicular a precipitation occurred on dislocation arrays. With increasing deformation, aging of cold worked Ti–15–3 first involved recovery of the heavily deformed substructure resulting in formation of equiaxed b subgrains, followed by equiaxed or globular a precipitation at the b sub-grain boundary triple points. Additional refinements to the beneficial effects quoted for cold work prior to aging were examined by Niwa et al. [3]; these investigators incorporated highlow temperature duplex aging within the processing sequence. High temperature aging, between 550 and 600°C, resulted in rapid a precipitation within the b matrix and partial recovery. The 2nd aging step at 400°C resulted in a finer, higher aspect ratio a; strength levels up to 1800 MPa and tensile ductilities of 5% elongation being reported. Niwa and Takatori [4] suggested moreover that the duplex microstructure developed with this type of heat treatment (without cold work) should provide a toughness benefit. Here the
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R.R. Boyer et al. / Materials Science and Engineering A243 (1998) 97–102
coarser a precipitates produced during the high temperature aging were thought to be key to the fracture toughness gain, promoting local crack deflection by the advancing crack, thus increasing the crack driving force necessary for further propagation. Finally Okada [5] reported that over-aging between 600 and 675°C prior to cold rolling followed by low temperature aging may also provide improved mechanical properties. This author once again suggested that the combination of large a precipitates formed during the prior over-aging and the fine precipitates resulting during final low temperature aging after rolling caused this improved property combination. Yield strengths in excess of 1700 MPa were achieved, although the tensile ductility was low, less than 2% elongation. None of these prior investigators, however, have examined the resultant fatigue properties of thermomechanically processed Ti – 15 – 3, although Wagner and Gregory [6] demonstrated improved fatigue performance for Ti–3Al – 8V – 6Cr – 4Mo – 4Zr following duplex aging. The present investigation has therefore focused on assessing the utility of this approach for enhancing the whole-life smooth fatigue performance of Ti–15–3.
2. Material/experimental procedures Ti–15V–3Cr–3Al – 3Sn 2.3 mm gage sheet with composition (wt.%) 14.625 V, 2.935 Cr, 3.000 Al, 3.06 Sn and 0.09 O2 was utilized for this investigation. Single, duplex and triplex aging treatments following either solution treatment or 35% cold or warm (600°C) rolling were examined. Aging was conducted in a re-circulating air furnace, the aging temperature being controlled to within 93°C. All aging cycles were terminated with an air cool. Specimens were prepared for optical metallography by mechanical polishing and etching with Kroll’s etchant for 6 s. Initially, tensile tests were conducted at 10 − 3 s per ASTM E8 using sheet-type specimens having a 25.4mm gage length. Sinusoidal wave form fatigue tests were then conducted per ASTM E466 using a reduced section, continuous radius sample having a minimum gage width of 9.5 mm at Kt =1, R = 0.1, and f = 10 Hz. Finally, selected tensile and fatigue fracture surfaces were examined by SEM.
3. Results and discussion The microstructures considered in this study are shown in Fig. 1, with a summary of the tensile properties achieved by these treatments being given in Table 1. Duplex aging (DA) clearly enhances the homogeneity of a precipitation when contrasted to single aging
(STA). The degree of this homogeneity enhancement is moreover dependent on the extent of the high temperature (500°C) age; increasing the high temperature hold time from 8 to 24 h decreases the uniformity of a precipitation. Further, the increase in hold time at this temperature also increases the a platelet size and decreases the tensile strength achievable by duplex aging. While multiple step aging can result in higher tensile properties when compared to STA treatments, correct selection of each aging time and temperature step is required. For example, TA-1, a triplex aged condition involving a low-high-intermediate sequence, had the lowest tensile strength observed in this study. Two factors, the short aging time at 300°C and the long aging time at 600°C, appear to contribute to this result. The former appears to have been insufficient to allow adequate formation of isothermal v. The presence of this metastable phase enhances precipitation kinetics by acting as a pre-cursor to a formation [7]. Additionally the aging time at 600°C may have been excessive, thereby resulting in extensive formation of coarse a. Increasing the low temperature aging time together with a decrease in high temperature aging time should be expected to, and was successful in, enhancing tensile performance (TA-2). Enhanced precipitate uniformity was also evident in thermomechanically processed material, the difference being most apparent when comparing STA and CR-A. While the aging cycles were identical, cold rolling prior to aging increased the uniformity of a precipitation. It was also expected that warm rolling should offer additional advantages over cold rolling; the dislocation substructure developed by elevated temperature deformation having previously been shown to be more homogeneous than that produced by cold rolling [8]. Indeed, preferential a precipitation along slip bands was observed in the CR-A condition, these not being present in the warm rolled plus aged condition. Unfortunately warm rolling and aging, at least for the conditions examined in this study, resulted in a tensile strength loss of 75 MPa. This decrease appears to be a reflection of the pre-aging temperature selected, coarser a precipitates being observed in the warm rolled plus aged material. (Previous investigations [2,3,5] indicated that the best mechanical properties were achieved with deformations on the order of 80%. The effect of these higher deformation levels on the fatigue performance of this alloy is part of an ongoing study.) Finally, high temperature over-aging, cold rolling and the final lower temperature aging(A-CR-A) resulted in a duplex a size distribution as proposed by Okada [5]. The tensile properties achieved were, however, intermediate between that previously reported and the STA condition, this intermediate value being ascribed to the lower cold reductions considered in the present investigation. However, the A-CR-A condition
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Fig. 1. Ti – 15V – 3Cr –3Al–3Sn microstructures representative of the studied thermal and thermomechanical treatments. (The micron marker for the ST condition applies to the top nine micrographs. The bottom three micrographs are at 4 × the magnification of the others.)
exhibited 8% tensile elongation, a more attractive property combination than that reported by Okada. Fig. 2 summarizes the fatigue performance of Ti– 15–3. Generally, the response may be grouped by tensile strength, the two highest strength conditions, DA-2 and A-CR-A, lying at the top of the scatterband, while the fatigue data for the lowest strength conditions, TA-1, STA and WR-A, lie within a tight scatterband at the low end of the data spectra. The lower run-out stress for the STA condition observed in the present investigation, when compared to Fanning [1]
and Muneki et al. [9], is also consistent with the lower STA tensile properties observed in this study. (The material Fanning tested had a tensile strength of 1150 MPa and Muneki et al. reported 1221 MPa.) This correlation between tensile strength and fatigue performance highlights the benefits to be gained from modifying the thermal and thermomechanical processing procedures used for Ti–15–3. Increased tensile and fatigue strength with little sacrifice in tensile ductility is achievable. For example at high stresses corresponding to an expected lifetime of 30 000 cycles, smax for the
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Table 1 Ti– 15 – 3 processing conditions and mechanical properties ID
Heat treatment
UTS (Mpa)
TYS (Mpa)
Elongation (%)
STA DA-1 DA-2 TA-1 TA-2 A-CR-A CR-A WR-A
540°C 8 h 500°C/24 h+400°C/100 h 500°C/8 h+400°C/100 h 300°C/10 h+600°C/16 h+450°C/100 h 300°C/100 h+600°C/4 h+450°C/100 h 600°C/24 h+35% CR+450°C/24 h 35% CR+540°C/8 hr. 35% WR (600°C)+540°C/8 h
1103 1310 1613 1070 1317 1420 1358 1275
1000 1206 1496 979 1206 1365 1338 1220
11 10 10 10 11 8 7 11
STA, solution treat and age; DA, duplex age; TA, triplex age; A, age; CR, cold roll; WR, warm roll. All heat treatments were terminated with an air cool.
STA condition was 775 and 875 MPa for the A-CR-A and DA-2 conditions, an improvement of 15%. Furthermore the improvement in the high-cycle regime was even more significant; DA-2 offering a cyclic stress improvement of 50% over the STA condition at 5×106 cycles. Microstructural differences also appear to be more important in the low-cycle regime than in the high-cycle regime. At a 30 000-cycle lifetime, the fatigue/tensile strength ratio varies from about 0.7 to 0.55. However, at 5 ×106 cycles this spread reduces to 0.10. Although this conclusion is obviously premised on a limited data set, it does include the extremes, DA-2, TA-1 and STA. DA-2 has the highest tensile strength and highest cyclic stress at 30 000 cycles. TA-1 and STA are the lowest strength conditions, and had the lowest cyclic stress at 30 000 cycles, excluding WR-A, the latter warm rolled condition ranking poorly throughout. Finally, fractographic examination indicates that the fatigue fracture mode is almost entirely transgranular, with reasonably well-defined striations observed on all specimens examined. Of these, only the WR-A condition exhibited extensive faceting and strong microstruc-
tural dependence on the fracture path (Fig. 4). The tensile specimens examined all failed by transgranular microvoid coalescence Currently it is difficult to completely rationalize the behavior of Ti–15–3. If high stress/low cycle fatigue lifetimes are considered to be dominated by crack propagation, then microstructures containing coarse a, with a lenticular or lamellar morphology, are expected to promote local crack deflection, reduce the effective stress intensity and enhance the high stress/low-cycle fatigue performance. This does not appear to be the case. In terms of smax at a 30 000-cycle life, The best conditions for a 30 000 cycle life were A-CR-A, DA-1 and DA-2; A-CR-A does have some of the coarse a structure, but the two duplex-ages have a fine precipitate structure. Analysis of the strength-normalized lowcycle data provides similar results-two of the three best performers have a coarse precipitate microstructure. On the other hand, the high cycle life, which is dominated by initiation, would be expected to be optimized with a fine, uniform a precipitation. The two best performers are DA-2 and TA-2, and the latter exhibits coarse precipitates and substantial grain boundary a.
Fig. 2. Fatigue properties of Ti–15V–3Cr–3Al–3Sn strip processed by various techniques. The data points at 1 000 000 cycles are actually run-outs at 5 000 000 cycles.
Fig. 3. Strength normalized fatigue properties of Ti – 15V–3Cr–3Al– 3Sn strip processed by various techniques. The data points at 1 000 000 cycles are actually run-outs at 5 000 000 cycles.
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Fig. 4. Electron fractographs from WR-A illustrating a microstructure-sensitive fracture topography and A-CR-A with a flat fracture topography.
The fatigue properties of the warm rolled material are also difficult to understand; the microstructure consists of uniform, fine a precipitation, which should exhibit good high-cycle fatigue characteristics. But it performs poorly, even worse than TA-1, which has a 200 MPa lower tensile strength and coarse lamellar a and grain boundary a. Indeed, these data suggest that the presence of grain boundary a, per se, may not be totally detrimental to fatigue performance in Ti–15–3. Conditions A-CR-A and TA-2 both have extensive grain boundary precipitation, yet they exhibit good fatigue properties (and crack propagation was transgranular). They are both in the high portion of the scatterband for cycles versus maximum stress (and fall approximately in the center of the scatterband for the strength normalized plot-Figs. 3 and 4). These results reinforce the importance that Porter and Eylon [10] have placed on the relative orientation of grain boundary a with respect to the operative stress axis, i.e. only grain boundary a oriented at 45° to the applied stress axis, along the plane of maximum shear, being found to be detrimental to the fatigue performance of Ti –15–3 castings.
cycle fatigue was improved by approximately 15% using either an over-age, cold roll and age sequence or highlow duplex aging. This advantage increases to approximately 50% in the high-cycle regime. The fatigue performance of Ti–15V–3Cr–3Al–3Sn seems to be relatively insensitive to microstructure, at least on an optical scale. In the low-cycle regime, where crack propagation would be expected to dominate, features such as a coarse lamellar or lenticular a would be expected to perform well, while those conditions with the very fine structure would be expected to perform poorly. In the high-cycle area one would expect the opposite effects. These suppositions were not observed in this study.
4. Conclusions
References
Attractive strength combinations can be achieved in Ti –15V–3Cr–3Al – 3Sn sheet through innovative heat treatment procedures and modifications in thermomechanical processing procedures, tensile strengths in excess of 1600 MPa being achieved while retaining 10% elongation. A direct correlation between tensile and fatigue strength was observed, the fatigue properties being enhanced over those of the standard STA heat-treatment. The maximum stress, smooth, R = 0.1, for low-
Acknowledgements The authors would like to acknowledge the funding support provided by NASA Langley Research Center under contract NAS1-20220. We would also like to thank Luther Gammon at The Boeing Company for his assistance with the metallography.
[1] J. Fanning, Fatigue data for Timetal® 15-3, in: D. Eylon, R.R. Boyer, D.A. Koss (Eds.), Beta Titanium Alloys in the 1990’s, TMS, Warrendale, PA, 1993, pp. 439 – 461 [2] T. Inaba, K. Ameyama, M. Tokizane, ISIJ Int. 31 (8) (1991) 792 – 798. [3] N. Niwa, A. Arai, H. Takatori, K. Ito, ISIJ Int. 31 (8) (1991) 856 – 862. [4] N. Niwa, H. Takatori, Effect of step-aging on the fracture toughness of Ti – 15V – 3Cr – 3Al – 3Sn alloy, in: D. Eylon, R.R. Boyer, D.A. Koss (Eds.), Beta Titanium Alloys in the 1990’s, TMS, Warrendale, PA, 1993, pp. 237 – 247. [5] M. Okada, ISIJ Int. 31 (8) (1991) 834 – 839.
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[6] L. Wagner and J.K. Gregory@, ‘Improvement of Mechanical Behavior in Ti-3Al–8V–6Cr–4Mo–4Zr by Duplex Aging,’ ISIJ Int., 31 (8) 199 –209. [7] H.J. Rack, S. Azimzadeh, R.R. Boyer, Precipitation phenomena in metastable beta titanium alloys, presented at Aeromat 1997, Williamsburg, VA, May 12, 1997. [8] H. Ohyama, Y. Ashida, T. Nishimura, T. Maki, ISIJ Int. 31 (8) (1991) 889 – 897. [9] S. Muneki, F. Morito, J. Takahashi, T. Kainuma, Y. Kawabe,
.
High cycle fatigue properties of beta titanium alloys, in: S. Fujishiro, D. Eylon, T. Kishi (Eds.), Metallurgy and Technology of Practical Titanium Alloys, TMS, Warrendale, PA, 1994, pp. 191 – 197. [10] W.J. Porter, D. Eylon, Effect of HIP and heat treatment on fatigue initiation and tensile failure in Ti – 15V – 3Cr–3Al–3Sn castings, in: D. Eylon, R.R. Boyer, D.A. Koss (Eds.), Beta Titanium Alloys in the 1990’s, TMS, Warrendale, PA, 1993, pp. 273 – 281
The influence of thermomechanical processing on the smooth fatigue properties of Ti–15V–3Cr–3Al–3Sn R.R. Boyer a,*, H.J. Rack b, V. Venkatesh b a
Boeing Commercial Airplane Group, PO Box 3707, MS 6H-CJ, Seattle, WA 98124, USA b Clemson Uni6ersity, Clemson, SC 29634, USA
Abstract This study has examined the potential for enhancing the fatigue performance of Ti – 15V – 3Cr – 3Al – 3Sn sheet through variations in heat treatment and thermomechanical processing procedures. Processing variants included conventional solution treatment and aging, solution treatment with duplex- and triplex-ages, and cold and warm rolling followed by aging. Tensile and smooth fatigue performance of each processing route were evaluated and compared to ‘standard’ commercial practice. These processing avenues provided a wide range of microstructures with tensile strengths ranging from 1070 to 1610 MPa. Enhancements in fatigue performance were dependent on processing and stress level. At high stresses, i.e. within the low cycle regime, a 15% improvement in the maximum stress for a fixed 30 000 cycle lifetime was achieved. At lower stresses, i.e. within the high cycle fatigue ‘limit’ regime, a 50% improvement relative to the standard heat treatment was achieved. A good correlation of fatigue and tensile strengths was observed; generally, the smooth fatigue properties improved as the tensile strength increased. © 1998 Elsevier Science S.A. All rights reserved. Keywords: Ti – 15V – 3Cr–3Al–3Sn; Thermomechanical processing; Smooth fatigue properties
1. Introduction ‘Standard’ commercial practice for Ti – 15V–3Cr– 3Al–3Sn (Ti–15– 3) thin sheet typically involves cold rolling, solution treatment at 790°C for 5 min. and aging at 540°C for 8 h. This procedure results in a typical tensile strength of 1035 MPa and a 106 cycle run-out stress of 700 MPa [1], which is inferior to that observed in a + b titanium alloys tested at equivalent strengths. Extensive examinations have analyzed the potential for enhancing the strength and fracture toughness of Ti– 15 – 3 through modifications in the alloy thermomechanical processing procedures [2–5]. For example, Inaba et al. [2] showed that significant strength enhancement can be achieved by cold rolling followed by single step aging. They reported that, under constant aging conditions, strength increased with increasing amounts of cold work. Further, an improvement of ductility was also achieved at 80% deformation, where a microduplex equiaxed a/b struc-
* Corresponding author. Tel.: +1 206 9652463; fax: + 1 206 9651440; e-mail: [email protected] 0921-5093/98/$19.00 © 1998 Elsevier Science S.A. All rights reserved. PII S 0 9 2 1 - 5 0 9 3 ( 9 7 ) 0 0 7 8 5 - 5
ture is formed upon aging. This latter enhancement was shown to be associated with recovery of the deformation substructure prior to a precipitation. With lesser degrees of deformation, direct lamellar or acicular a precipitation occurred on dislocation arrays. With increasing deformation, aging of cold worked Ti–15–3 first involved recovery of the heavily deformed substructure resulting in formation of equiaxed b subgrains, followed by equiaxed or globular a precipitation at the b sub-grain boundary triple points. Additional refinements to the beneficial effects quoted for cold work prior to aging were examined by Niwa et al. [3]; these investigators incorporated highlow temperature duplex aging within the processing sequence. High temperature aging, between 550 and 600°C, resulted in rapid a precipitation within the b matrix and partial recovery. The 2nd aging step at 400°C resulted in a finer, higher aspect ratio a; strength levels up to 1800 MPa and tensile ductilities of 5% elongation being reported. Niwa and Takatori [4] suggested moreover that the duplex microstructure developed with this type of heat treatment (without cold work) should provide a toughness benefit. Here the
98
R.R. Boyer et al. / Materials Science and Engineering A243 (1998) 97–102
coarser a precipitates produced during the high temperature aging were thought to be key to the fracture toughness gain, promoting local crack deflection by the advancing crack, thus increasing the crack driving force necessary for further propagation. Finally Okada [5] reported that over-aging between 600 and 675°C prior to cold rolling followed by low temperature aging may also provide improved mechanical properties. This author once again suggested that the combination of large a precipitates formed during the prior over-aging and the fine precipitates resulting during final low temperature aging after rolling caused this improved property combination. Yield strengths in excess of 1700 MPa were achieved, although the tensile ductility was low, less than 2% elongation. None of these prior investigators, however, have examined the resultant fatigue properties of thermomechanically processed Ti – 15 – 3, although Wagner and Gregory [6] demonstrated improved fatigue performance for Ti–3Al – 8V – 6Cr – 4Mo – 4Zr following duplex aging. The present investigation has therefore focused on assessing the utility of this approach for enhancing the whole-life smooth fatigue performance of Ti–15–3.
2. Material/experimental procedures Ti–15V–3Cr–3Al – 3Sn 2.3 mm gage sheet with composition (wt.%) 14.625 V, 2.935 Cr, 3.000 Al, 3.06 Sn and 0.09 O2 was utilized for this investigation. Single, duplex and triplex aging treatments following either solution treatment or 35% cold or warm (600°C) rolling were examined. Aging was conducted in a re-circulating air furnace, the aging temperature being controlled to within 93°C. All aging cycles were terminated with an air cool. Specimens were prepared for optical metallography by mechanical polishing and etching with Kroll’s etchant for 6 s. Initially, tensile tests were conducted at 10 − 3 s per ASTM E8 using sheet-type specimens having a 25.4mm gage length. Sinusoidal wave form fatigue tests were then conducted per ASTM E466 using a reduced section, continuous radius sample having a minimum gage width of 9.5 mm at Kt =1, R = 0.1, and f = 10 Hz. Finally, selected tensile and fatigue fracture surfaces were examined by SEM.
3. Results and discussion The microstructures considered in this study are shown in Fig. 1, with a summary of the tensile properties achieved by these treatments being given in Table 1. Duplex aging (DA) clearly enhances the homogeneity of a precipitation when contrasted to single aging
(STA). The degree of this homogeneity enhancement is moreover dependent on the extent of the high temperature (500°C) age; increasing the high temperature hold time from 8 to 24 h decreases the uniformity of a precipitation. Further, the increase in hold time at this temperature also increases the a platelet size and decreases the tensile strength achievable by duplex aging. While multiple step aging can result in higher tensile properties when compared to STA treatments, correct selection of each aging time and temperature step is required. For example, TA-1, a triplex aged condition involving a low-high-intermediate sequence, had the lowest tensile strength observed in this study. Two factors, the short aging time at 300°C and the long aging time at 600°C, appear to contribute to this result. The former appears to have been insufficient to allow adequate formation of isothermal v. The presence of this metastable phase enhances precipitation kinetics by acting as a pre-cursor to a formation [7]. Additionally the aging time at 600°C may have been excessive, thereby resulting in extensive formation of coarse a. Increasing the low temperature aging time together with a decrease in high temperature aging time should be expected to, and was successful in, enhancing tensile performance (TA-2). Enhanced precipitate uniformity was also evident in thermomechanically processed material, the difference being most apparent when comparing STA and CR-A. While the aging cycles were identical, cold rolling prior to aging increased the uniformity of a precipitation. It was also expected that warm rolling should offer additional advantages over cold rolling; the dislocation substructure developed by elevated temperature deformation having previously been shown to be more homogeneous than that produced by cold rolling [8]. Indeed, preferential a precipitation along slip bands was observed in the CR-A condition, these not being present in the warm rolled plus aged condition. Unfortunately warm rolling and aging, at least for the conditions examined in this study, resulted in a tensile strength loss of 75 MPa. This decrease appears to be a reflection of the pre-aging temperature selected, coarser a precipitates being observed in the warm rolled plus aged material. (Previous investigations [2,3,5] indicated that the best mechanical properties were achieved with deformations on the order of 80%. The effect of these higher deformation levels on the fatigue performance of this alloy is part of an ongoing study.) Finally, high temperature over-aging, cold rolling and the final lower temperature aging(A-CR-A) resulted in a duplex a size distribution as proposed by Okada [5]. The tensile properties achieved were, however, intermediate between that previously reported and the STA condition, this intermediate value being ascribed to the lower cold reductions considered in the present investigation. However, the A-CR-A condition
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Fig. 1. Ti – 15V – 3Cr –3Al–3Sn microstructures representative of the studied thermal and thermomechanical treatments. (The micron marker for the ST condition applies to the top nine micrographs. The bottom three micrographs are at 4 × the magnification of the others.)
exhibited 8% tensile elongation, a more attractive property combination than that reported by Okada. Fig. 2 summarizes the fatigue performance of Ti– 15–3. Generally, the response may be grouped by tensile strength, the two highest strength conditions, DA-2 and A-CR-A, lying at the top of the scatterband, while the fatigue data for the lowest strength conditions, TA-1, STA and WR-A, lie within a tight scatterband at the low end of the data spectra. The lower run-out stress for the STA condition observed in the present investigation, when compared to Fanning [1]
and Muneki et al. [9], is also consistent with the lower STA tensile properties observed in this study. (The material Fanning tested had a tensile strength of 1150 MPa and Muneki et al. reported 1221 MPa.) This correlation between tensile strength and fatigue performance highlights the benefits to be gained from modifying the thermal and thermomechanical processing procedures used for Ti–15–3. Increased tensile and fatigue strength with little sacrifice in tensile ductility is achievable. For example at high stresses corresponding to an expected lifetime of 30 000 cycles, smax for the
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Table 1 Ti– 15 – 3 processing conditions and mechanical properties ID
Heat treatment
UTS (Mpa)
TYS (Mpa)
Elongation (%)
STA DA-1 DA-2 TA-1 TA-2 A-CR-A CR-A WR-A
540°C 8 h 500°C/24 h+400°C/100 h 500°C/8 h+400°C/100 h 300°C/10 h+600°C/16 h+450°C/100 h 300°C/100 h+600°C/4 h+450°C/100 h 600°C/24 h+35% CR+450°C/24 h 35% CR+540°C/8 hr. 35% WR (600°C)+540°C/8 h
1103 1310 1613 1070 1317 1420 1358 1275
1000 1206 1496 979 1206 1365 1338 1220
11 10 10 10 11 8 7 11
STA, solution treat and age; DA, duplex age; TA, triplex age; A, age; CR, cold roll; WR, warm roll. All heat treatments were terminated with an air cool.
STA condition was 775 and 875 MPa for the A-CR-A and DA-2 conditions, an improvement of 15%. Furthermore the improvement in the high-cycle regime was even more significant; DA-2 offering a cyclic stress improvement of 50% over the STA condition at 5×106 cycles. Microstructural differences also appear to be more important in the low-cycle regime than in the high-cycle regime. At a 30 000-cycle lifetime, the fatigue/tensile strength ratio varies from about 0.7 to 0.55. However, at 5 ×106 cycles this spread reduces to 0.10. Although this conclusion is obviously premised on a limited data set, it does include the extremes, DA-2, TA-1 and STA. DA-2 has the highest tensile strength and highest cyclic stress at 30 000 cycles. TA-1 and STA are the lowest strength conditions, and had the lowest cyclic stress at 30 000 cycles, excluding WR-A, the latter warm rolled condition ranking poorly throughout. Finally, fractographic examination indicates that the fatigue fracture mode is almost entirely transgranular, with reasonably well-defined striations observed on all specimens examined. Of these, only the WR-A condition exhibited extensive faceting and strong microstruc-
tural dependence on the fracture path (Fig. 4). The tensile specimens examined all failed by transgranular microvoid coalescence Currently it is difficult to completely rationalize the behavior of Ti–15–3. If high stress/low cycle fatigue lifetimes are considered to be dominated by crack propagation, then microstructures containing coarse a, with a lenticular or lamellar morphology, are expected to promote local crack deflection, reduce the effective stress intensity and enhance the high stress/low-cycle fatigue performance. This does not appear to be the case. In terms of smax at a 30 000-cycle life, The best conditions for a 30 000 cycle life were A-CR-A, DA-1 and DA-2; A-CR-A does have some of the coarse a structure, but the two duplex-ages have a fine precipitate structure. Analysis of the strength-normalized lowcycle data provides similar results-two of the three best performers have a coarse precipitate microstructure. On the other hand, the high cycle life, which is dominated by initiation, would be expected to be optimized with a fine, uniform a precipitation. The two best performers are DA-2 and TA-2, and the latter exhibits coarse precipitates and substantial grain boundary a.
Fig. 2. Fatigue properties of Ti–15V–3Cr–3Al–3Sn strip processed by various techniques. The data points at 1 000 000 cycles are actually run-outs at 5 000 000 cycles.
Fig. 3. Strength normalized fatigue properties of Ti – 15V–3Cr–3Al– 3Sn strip processed by various techniques. The data points at 1 000 000 cycles are actually run-outs at 5 000 000 cycles.
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Fig. 4. Electron fractographs from WR-A illustrating a microstructure-sensitive fracture topography and A-CR-A with a flat fracture topography.
The fatigue properties of the warm rolled material are also difficult to understand; the microstructure consists of uniform, fine a precipitation, which should exhibit good high-cycle fatigue characteristics. But it performs poorly, even worse than TA-1, which has a 200 MPa lower tensile strength and coarse lamellar a and grain boundary a. Indeed, these data suggest that the presence of grain boundary a, per se, may not be totally detrimental to fatigue performance in Ti–15–3. Conditions A-CR-A and TA-2 both have extensive grain boundary precipitation, yet they exhibit good fatigue properties (and crack propagation was transgranular). They are both in the high portion of the scatterband for cycles versus maximum stress (and fall approximately in the center of the scatterband for the strength normalized plot-Figs. 3 and 4). These results reinforce the importance that Porter and Eylon [10] have placed on the relative orientation of grain boundary a with respect to the operative stress axis, i.e. only grain boundary a oriented at 45° to the applied stress axis, along the plane of maximum shear, being found to be detrimental to the fatigue performance of Ti –15–3 castings.
cycle fatigue was improved by approximately 15% using either an over-age, cold roll and age sequence or highlow duplex aging. This advantage increases to approximately 50% in the high-cycle regime. The fatigue performance of Ti–15V–3Cr–3Al–3Sn seems to be relatively insensitive to microstructure, at least on an optical scale. In the low-cycle regime, where crack propagation would be expected to dominate, features such as a coarse lamellar or lenticular a would be expected to perform well, while those conditions with the very fine structure would be expected to perform poorly. In the high-cycle area one would expect the opposite effects. These suppositions were not observed in this study.
4. Conclusions
References
Attractive strength combinations can be achieved in Ti –15V–3Cr–3Al – 3Sn sheet through innovative heat treatment procedures and modifications in thermomechanical processing procedures, tensile strengths in excess of 1600 MPa being achieved while retaining 10% elongation. A direct correlation between tensile and fatigue strength was observed, the fatigue properties being enhanced over those of the standard STA heat-treatment. The maximum stress, smooth, R = 0.1, for low-
Acknowledgements The authors would like to acknowledge the funding support provided by NASA Langley Research Center under contract NAS1-20220. We would also like to thank Luther Gammon at The Boeing Company for his assistance with the metallography.
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