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Effects of lamellar spacing on mechanical properties of fully lamellar Ti–39.4mol%Al alloy
Materials Science and Engineering A319– 321 (2001) 360– 363 www.elsevier.com/locate/msea
Effects of lamellar spacing on mechanical properties of fully lamellar Ti–39.4mol%Al alloy K. Maruyama a,*, N. Yamada a,1, H. Sato b a
b
Department of Materials Science, Graduate School of Engineering, Tohoku Uni6ersity, Sendai 980 -8579, Japan Department of Intelligent Machine and Systems Engineering, Faculty of Science and Technology, Hirosaki Uni6ersity, Hirosaki 036 -8561, Japan
Abstract Effects of lamellar spacing on mechanical properties of a fully lamellar TiAl alloy were investigated at room temperature and at 950K by compression tests. Lamellar spacing was varied from 20 to 590 nm, keeping the same grain size of 90 mm. The Hall–Petch relation holds between lamellar spacing and yield stress of the alloy. However, the yield stress saturates at a value of about 1 GPa below a critical lamellar spacing of about 100 nm. The Hall– Petch slope, critical lamellar spacing and saturation stress can be explained consistently by the pile-up model of dislocations at lamellar interfaces. Contrary to the beneficial effect on yield stress, the refinement of lamellar spacing deteriorates ductility of the alloy probably due to the suppression of the hard deformation modes. © 2001 Elsevier Science B.V. All rights reserved. Keywords: Titanium aluminide; Lamellar microstructure; Lamellar spacing; Yield stress; Fracture; Hall– Petch relation
1. Introduction High strength at elevated temperatures and good ductility at room temperature are essential requirements for high temperature structural applications of TiAl alloys. The alloys consist of gTiAl and a2Ti3Al phases, and take various microstructures depending on their heat treatment. Among the microstructures, a finegrained fully lamellar structure can provide a good combination of room temperature ductility and creep strength at elevated temperatures [1]. For this reason, mechanical properties of fully lamellar TiAl alloys have been studied extensively in the last decade. As for elevated temperature strength, a fine and thermally stable lamellar microstructure is essential for good creep resistance [2 – 5]. Maziasz and Liu [6,7] and Dimiduk et al. [8] have pointed out that the Hall –Petch relation holds between yield stress sy and lamellar spacing l at elevated temperatures as well as room temperature, * Corresponding author. Tel./Fax: + 81-22-2177324. E-mail address: [email protected] (K. Maruyama). 1 Now with Toshiba Corporation.
(1) sy = s0 + k/ u where s0 and k are material constants. These findings recommend lamellar refinement for improving mechanical properties of TiAl alloys. Most of experiments on fully lamellar TiAl alloys have been carried out in ranges of lamellar spacing thicker than 100 nm. It is of engineering as well as scientific importance to know whether further refinement of lamellar spacing can improve mechanical properties of the alloys. Eq. (1) has been proved over the ranges of lamellar spacing from 35 to 150 nm in Ref. [8], and from 100 to 440 nm in Refs. [6,7]. In these experiments, however, not only lamellar spacing but also grain size and a2 volume fraction were changed by the heat treatment to control the lamellar microstructures. On the other hand, Sun [9] has expected on the basis of the pile-up model of dislocations that yield stress saturates at a value and the Hall –Petch relation does not hold below a critical lamellar spacing. This expectation has not been proved for TiAl alloys. It has not been well understood how lamellar spacing affects ductility of lamellar materials. In the present paper, yield stress and ductility of a fully lamellar TiAl alloy are studied at room temperature and 950 K paying special attention to the effects of lamellar spacing.
0921-5093/01/$ - see front matter © 2001 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 1 - 5 0 9 3 ( 0 1 ) 0 0 9 2 6 - 1
K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
2. Experimental procedure The material used was a binary Ti-39.4mol%Al alloy (0.043mol%O) close to the eutectoid composition. The alloy contains 70vol% of a2 phase. The alloy was induction-skull melted, and then hot extruded at 1473 K to 55% reduction of area. Compression specimens, 2×2 mm in cross section and 3 mm in height, were machined so that their stress axis was parallel to the extrusion direction. All the specimens were annealed in a vacuum of 3× 10 − 4Pa at 1450 K (in a single phase regime) for 600 s to have the same grain size of 90 mm. Furnace cooling after the solution treatment (initial cooling rate of 0.7 Ks − 1) was enough to prevent the precipitation of g plates in a2 matrix. This is a benefit of the low aluminum composition selected in the present study. After the solution treatment, specimens were furnace cooled to a temperature below the eutectoid point, and then isothermally aged. The time– temperature –precipitation diagram of the alloy [10] gave the aging time at which specimens were fully converted to the lamellar microstructure. The shortest aging time
was 3 ks at the nose temperature of 1175 K, and the longest, 288 ks at 1300 K. Lamellar spacing was measured by TEM in the direction perpendicular to the lamellar interfaces. Since most of lamellar interfaces were a2/g boundaries in the alloy of low g volume fraction, an average of lamellar spacings was taken without regard to the phase type of the adjacent lamella. Compression tests were carried out under an initial strain rate of 4.9× 10 − 3s − 1. The tests at 950K were performed in a vacuum of 3× 10 − 3Pa. 0.5% proof stress was taken as yield stress. Some tests were interrupted after 2.5% plastic deformation to study the deformation substructure.
3. Results and discussion
3.1. Lamellar microstructure Fully lamellar microstructures were formed by isothermal aging of the Ti-39.4mol%Al alloy at temperatures ranging from 950 to 1300 K. The average values of lamellar spacing l are given in Fig. 1 as a function of the extent of undercooling DT from the eutectoid temperature of 1400 K. The lamellar spacing decreases with increasing DT, and can be described by the following relation: l8 DT − n
Fig. 1. Lamellar spacing as a function of the extent of undercooling from the eutectoid temperature.
361
(2)
The exponent n is expected to be unity in the case of pearlite structures of eutectoid steels, but n=2 in the present alloy as indicated by the dashed line in Fig. 1. If different u− DT relations hold above and below the nose temperature of g precipitation, the experimental result may be represented by the solid line: n=1.4 above the nose, and n= 3.8 below the nose. In the present study, lamellar spacing could be varied from 20 to 590 nm without introducing acicular or massive phase. This is another benefit of the present alloy having low aluminum concentration. The specimens subjected to compression tests are indicated by the solid symbols in Fig. 1. Blackburn’s orientation relationship [11] between the a2 and g phases was confirmed over the whole range of lamellar spacing. Grain boundaries are straight in fine lamellar microstructures, whereas interlocked grain boundaries are formed in coarse lamellar microstructures.
3.2. Yield stress
Fig. 2. Stress –strain curves at room temperature.
Fig. 2 shows representative stress versus strain curves at room temperature (R.T.). The yield stress increases with decreasing lamellar spacing l. Some specimens failed in a brittle manner even in compression tests, and the ductility and apparent strain hardening rate decrease with decreasing l. These changes of stress–strain
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K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
has predicted the saturation based on multiple pile-ups of dislocations at each lamellar interface. According to the traditional pile-up model of edge dislocations at a single barrier, the Hall–Petch slope k, the critical lamellar spacing l*, and the saturation stress s* are given by,
Fig. 3. Hall – Petch plot of yield stress, together with the theoretical lines of Eqs. (3) and (5).
k= M Gb~*/(1− w)
(3)
u*= Gb/(1− w)~*
(4)
|*= M~*
(5)
where M is the Taylor factor (= 3), G the shear modulus (61 GPa), b the length of Burgers vector (0.28 nm), n the Poisson’s ratio (0.28), and t* the shear strength of the barrier to dislocation motion. The simulation result of Sun [9] agrees well with Eqs. (3)–(5). The basic idea of Eq. (4) is that yield stress does not increase anymore when the average number of dislocations piled up at each interface is less than unity. The dashed lines in Fig. 3 are the prediction from Eqs. (3)– (5). t* was taken to be 300 MPa as a reasonable example. The predicted values are k=0.25MPa m, l*= 80 nm, and s* =900 GPa. s0 in Eq. (1) is not zero but neglected when drawing the dashed line. The predicted values of k, l* and s* agree fairly well with the experimental results, suggesting that the pile-up model can consistently explain k, l* and s* of the TiAl alloy. The aforementioned findings suggest that the maximum yield stress attainable by lamellar refinement is s*. The barrier strength t* should be increased to improve the saturation stress s*.
Fig. 4. Fracture strain as a function of lamellar spacing.
3.3. Ductility curve shape at R.T. are essentially the same at 950 K. The values of yield stress are plotted against l in Fig. 3. The Hall– Petch relation (Eq. (1)) holds between the yield stress and l when l \ 100 nm. The Hall–Petch slope k is 0.29MPa m at R.T. and 0.26MPa m at 950 K. The same relation has been reported on fully lamellar TiAl alloys: k =0.22MPa m at R.T. and 0.15MPa m at 1073 K in Ref.[7], k= 0.11MPa m in Ref. [8], and k = 0.50 and 0.41MPa m on hard oriented PST crystals [12]. The present values of k are similar to these values reported in the literature. It should be noted in Fig. 3 that the yield stress saturates at a value s* below a critical lamellar spacing l* of about 100 nm. The experimental values of s* are around 1 GPa, and decrease with increasing temperature. They are substantially lower than 3.6 GPa expected in Ref. [8]. The saturation of yield stress is common to R.T. and 950 K, and should be attributed not to interface sliding but to some intrinsic mechanism. However, the decrease of yield stress from l =70 to 20 nm at 950 K suggests the contribution of interface sliding. Sun [9]
The TiAl alloy shows apparent strain softening due to microcracking at the later stage of stress– strain curve in Fig. 2. Fracture occurred in compression tests, and specimens were crashed to pieces in some cases. Fracture strain was defined as the strain at which flow stress was reduced to 90% of the maximum value, and is indicated by the short vertical line in Fig. 2. The values of fracture strain are plotted again lamellar spacing in Fig. 4. The fracture strain is reduced at the lower temperature and decreases with decreasing lamellar spacing. This result demonstrates a detrimental consequence of lamellar refinement on ductility of lamellar TiAl alloys. Fig. 5 shows deformation substructure after 2.5% plastic deformation of a coarse lamellar microstructure. The lamellar interfaces are parallel to the beam direction in the micrographs. a= 112( 0 type dislocations in a2 lamellae and cross twins in g lamellae are seen in Fig. 5(a) and (b), respectively. The a dislocations cause the soft modes of deformation, and the cross twins, the hard modes. Therefore, the plastic deformation in the coarse lamellar specimen is isotropic in a macroscopic scale. On the other hand, cross twins
K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
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with decreasing lamellar spacing, and the Hall–Petch relation holds between the yield stress and lamellar spacing when lamellar spacing is thicker than a critical value l*. 2. The yield stress saturates at a value s* below l*. s* is close to 1 GPa and l* is about 100 nm in the Ti-39.4mol%Al alloy. The pile-up model of dislocations can explain consistently the saturation stress, critical lamellar spacing and Hall–Petch slope. 3. Refinement of lamellar microstructure reduces fracture strain, since the hard deformation modes are suppressed in fine lamellar microstructures.
Acknowledgements This research was supported by JSPS (No.JSPSRFTF96R12301) and the Ministry of Education, Science, Sports and Culture, Japan (No.11450259). The material used was kindly provided by Kobe Steel Ltd. The authors would like to thank Mr G. Suzuki of Tohoku University for his help with mechanical testing.
References
Fig. 5. Deformation substructures formed in a thick lamellar microstructure (l= 590 nm) after 2.5% plastic deformation at 950 K. (a) g=202( 1a2, (b) g = 11( 1g.
were hardly seen in fine lamellar microstructures, suggesting that the plastic deformation takes place primarily by the soft deformation modes in the fine lamellar specimens. This anisotropic deformation builds up incompatibility stresses and eventually promotes the premature cracking of the fine lamellar microstructures.
4. Conclusions 1. Yield stress of fully lamellar TiAl alloys increases
[1] Y.-W. Kim, JOM 46 (7) (1994) 30. [2] K. Maruyama, R. Yamamoto, H. Nakakuki, N. Fujitsuna, Mater. Sci. Eng. A239-240 (1997) 419. [3] J. Beddoes, J. Triantafillou, L. Zhao, MRS Symp. Proc. 460 (1997) 293. [4] R. Yamamoto, K. Mizoguchi, G. Wegmann, K. Maruyama, Intermetallics 6 (1998) 699. [5] R. Yamamoto, K. Mizoguchi, G. Wegmann, K. Maruyama, in: Y.-W. Kim, et al. (Eds.), Gamma Titanium Aluminides 1999, TMS, Warrendale, 1999, p. 461. [6] P.J. Maziasz, C.T. Liu, Metall. Mater. Trans. 29A (1998) 105. [7] C.T. Liu, P.J. Maziasz, Intermetallics 6 (1998) 653. [8] D.M. Dimiduk, P.M. Hazzledine, T.A. Parthasarathy, S. Seshagiri, M.G. Mendiratta, Metall. Mater. Trans. 29A (1998) 37. [9] Y.Q. Sun, Mater. Sci. Eng. A239-240 (1997) 131. [10] N. Yamada, H. Sato, K. Maruyama, Report of the 123rd Committee on Heat-Resisting Metals and Alloys, JSPS 40 (1999) 273. [11] M.J. Blackburn, in: R.T. Jaffee, N.E. Promisel (Eds.), Science, Technology and Application of Titanium, Pergamon Press, London, 1970, p. 633. [12] Y. Umakoshi, T. Nakano, Acta Metall. Mater. 41 (1993) 1155.
Effects of lamellar spacing on mechanical properties of fully lamellar Ti–39.4mol%Al alloy K. Maruyama a,*, N. Yamada a,1, H. Sato b a
b
Department of Materials Science, Graduate School of Engineering, Tohoku Uni6ersity, Sendai 980 -8579, Japan Department of Intelligent Machine and Systems Engineering, Faculty of Science and Technology, Hirosaki Uni6ersity, Hirosaki 036 -8561, Japan
Abstract Effects of lamellar spacing on mechanical properties of a fully lamellar TiAl alloy were investigated at room temperature and at 950K by compression tests. Lamellar spacing was varied from 20 to 590 nm, keeping the same grain size of 90 mm. The Hall–Petch relation holds between lamellar spacing and yield stress of the alloy. However, the yield stress saturates at a value of about 1 GPa below a critical lamellar spacing of about 100 nm. The Hall– Petch slope, critical lamellar spacing and saturation stress can be explained consistently by the pile-up model of dislocations at lamellar interfaces. Contrary to the beneficial effect on yield stress, the refinement of lamellar spacing deteriorates ductility of the alloy probably due to the suppression of the hard deformation modes. © 2001 Elsevier Science B.V. All rights reserved. Keywords: Titanium aluminide; Lamellar microstructure; Lamellar spacing; Yield stress; Fracture; Hall– Petch relation
1. Introduction High strength at elevated temperatures and good ductility at room temperature are essential requirements for high temperature structural applications of TiAl alloys. The alloys consist of gTiAl and a2Ti3Al phases, and take various microstructures depending on their heat treatment. Among the microstructures, a finegrained fully lamellar structure can provide a good combination of room temperature ductility and creep strength at elevated temperatures [1]. For this reason, mechanical properties of fully lamellar TiAl alloys have been studied extensively in the last decade. As for elevated temperature strength, a fine and thermally stable lamellar microstructure is essential for good creep resistance [2 – 5]. Maziasz and Liu [6,7] and Dimiduk et al. [8] have pointed out that the Hall –Petch relation holds between yield stress sy and lamellar spacing l at elevated temperatures as well as room temperature, * Corresponding author. Tel./Fax: + 81-22-2177324. E-mail address: [email protected] (K. Maruyama). 1 Now with Toshiba Corporation.
(1) sy = s0 + k/ u where s0 and k are material constants. These findings recommend lamellar refinement for improving mechanical properties of TiAl alloys. Most of experiments on fully lamellar TiAl alloys have been carried out in ranges of lamellar spacing thicker than 100 nm. It is of engineering as well as scientific importance to know whether further refinement of lamellar spacing can improve mechanical properties of the alloys. Eq. (1) has been proved over the ranges of lamellar spacing from 35 to 150 nm in Ref. [8], and from 100 to 440 nm in Refs. [6,7]. In these experiments, however, not only lamellar spacing but also grain size and a2 volume fraction were changed by the heat treatment to control the lamellar microstructures. On the other hand, Sun [9] has expected on the basis of the pile-up model of dislocations that yield stress saturates at a value and the Hall –Petch relation does not hold below a critical lamellar spacing. This expectation has not been proved for TiAl alloys. It has not been well understood how lamellar spacing affects ductility of lamellar materials. In the present paper, yield stress and ductility of a fully lamellar TiAl alloy are studied at room temperature and 950 K paying special attention to the effects of lamellar spacing.
0921-5093/01/$ - see front matter © 2001 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 1 - 5 0 9 3 ( 0 1 ) 0 0 9 2 6 - 1
K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
2. Experimental procedure The material used was a binary Ti-39.4mol%Al alloy (0.043mol%O) close to the eutectoid composition. The alloy contains 70vol% of a2 phase. The alloy was induction-skull melted, and then hot extruded at 1473 K to 55% reduction of area. Compression specimens, 2×2 mm in cross section and 3 mm in height, were machined so that their stress axis was parallel to the extrusion direction. All the specimens were annealed in a vacuum of 3× 10 − 4Pa at 1450 K (in a single phase regime) for 600 s to have the same grain size of 90 mm. Furnace cooling after the solution treatment (initial cooling rate of 0.7 Ks − 1) was enough to prevent the precipitation of g plates in a2 matrix. This is a benefit of the low aluminum composition selected in the present study. After the solution treatment, specimens were furnace cooled to a temperature below the eutectoid point, and then isothermally aged. The time– temperature –precipitation diagram of the alloy [10] gave the aging time at which specimens were fully converted to the lamellar microstructure. The shortest aging time
was 3 ks at the nose temperature of 1175 K, and the longest, 288 ks at 1300 K. Lamellar spacing was measured by TEM in the direction perpendicular to the lamellar interfaces. Since most of lamellar interfaces were a2/g boundaries in the alloy of low g volume fraction, an average of lamellar spacings was taken without regard to the phase type of the adjacent lamella. Compression tests were carried out under an initial strain rate of 4.9× 10 − 3s − 1. The tests at 950K were performed in a vacuum of 3× 10 − 3Pa. 0.5% proof stress was taken as yield stress. Some tests were interrupted after 2.5% plastic deformation to study the deformation substructure.
3. Results and discussion
3.1. Lamellar microstructure Fully lamellar microstructures were formed by isothermal aging of the Ti-39.4mol%Al alloy at temperatures ranging from 950 to 1300 K. The average values of lamellar spacing l are given in Fig. 1 as a function of the extent of undercooling DT from the eutectoid temperature of 1400 K. The lamellar spacing decreases with increasing DT, and can be described by the following relation: l8 DT − n
Fig. 1. Lamellar spacing as a function of the extent of undercooling from the eutectoid temperature.
361
(2)
The exponent n is expected to be unity in the case of pearlite structures of eutectoid steels, but n=2 in the present alloy as indicated by the dashed line in Fig. 1. If different u− DT relations hold above and below the nose temperature of g precipitation, the experimental result may be represented by the solid line: n=1.4 above the nose, and n= 3.8 below the nose. In the present study, lamellar spacing could be varied from 20 to 590 nm without introducing acicular or massive phase. This is another benefit of the present alloy having low aluminum concentration. The specimens subjected to compression tests are indicated by the solid symbols in Fig. 1. Blackburn’s orientation relationship [11] between the a2 and g phases was confirmed over the whole range of lamellar spacing. Grain boundaries are straight in fine lamellar microstructures, whereas interlocked grain boundaries are formed in coarse lamellar microstructures.
3.2. Yield stress
Fig. 2. Stress –strain curves at room temperature.
Fig. 2 shows representative stress versus strain curves at room temperature (R.T.). The yield stress increases with decreasing lamellar spacing l. Some specimens failed in a brittle manner even in compression tests, and the ductility and apparent strain hardening rate decrease with decreasing l. These changes of stress–strain
362
K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
has predicted the saturation based on multiple pile-ups of dislocations at each lamellar interface. According to the traditional pile-up model of edge dislocations at a single barrier, the Hall–Petch slope k, the critical lamellar spacing l*, and the saturation stress s* are given by,
Fig. 3. Hall – Petch plot of yield stress, together with the theoretical lines of Eqs. (3) and (5).
k= M Gb~*/(1− w)
(3)
u*= Gb/(1− w)~*
(4)
|*= M~*
(5)
where M is the Taylor factor (= 3), G the shear modulus (61 GPa), b the length of Burgers vector (0.28 nm), n the Poisson’s ratio (0.28), and t* the shear strength of the barrier to dislocation motion. The simulation result of Sun [9] agrees well with Eqs. (3)–(5). The basic idea of Eq. (4) is that yield stress does not increase anymore when the average number of dislocations piled up at each interface is less than unity. The dashed lines in Fig. 3 are the prediction from Eqs. (3)– (5). t* was taken to be 300 MPa as a reasonable example. The predicted values are k=0.25MPa m, l*= 80 nm, and s* =900 GPa. s0 in Eq. (1) is not zero but neglected when drawing the dashed line. The predicted values of k, l* and s* agree fairly well with the experimental results, suggesting that the pile-up model can consistently explain k, l* and s* of the TiAl alloy. The aforementioned findings suggest that the maximum yield stress attainable by lamellar refinement is s*. The barrier strength t* should be increased to improve the saturation stress s*.
Fig. 4. Fracture strain as a function of lamellar spacing.
3.3. Ductility curve shape at R.T. are essentially the same at 950 K. The values of yield stress are plotted against l in Fig. 3. The Hall– Petch relation (Eq. (1)) holds between the yield stress and l when l \ 100 nm. The Hall–Petch slope k is 0.29MPa m at R.T. and 0.26MPa m at 950 K. The same relation has been reported on fully lamellar TiAl alloys: k =0.22MPa m at R.T. and 0.15MPa m at 1073 K in Ref.[7], k= 0.11MPa m in Ref. [8], and k = 0.50 and 0.41MPa m on hard oriented PST crystals [12]. The present values of k are similar to these values reported in the literature. It should be noted in Fig. 3 that the yield stress saturates at a value s* below a critical lamellar spacing l* of about 100 nm. The experimental values of s* are around 1 GPa, and decrease with increasing temperature. They are substantially lower than 3.6 GPa expected in Ref. [8]. The saturation of yield stress is common to R.T. and 950 K, and should be attributed not to interface sliding but to some intrinsic mechanism. However, the decrease of yield stress from l =70 to 20 nm at 950 K suggests the contribution of interface sliding. Sun [9]
The TiAl alloy shows apparent strain softening due to microcracking at the later stage of stress– strain curve in Fig. 2. Fracture occurred in compression tests, and specimens were crashed to pieces in some cases. Fracture strain was defined as the strain at which flow stress was reduced to 90% of the maximum value, and is indicated by the short vertical line in Fig. 2. The values of fracture strain are plotted again lamellar spacing in Fig. 4. The fracture strain is reduced at the lower temperature and decreases with decreasing lamellar spacing. This result demonstrates a detrimental consequence of lamellar refinement on ductility of lamellar TiAl alloys. Fig. 5 shows deformation substructure after 2.5% plastic deformation of a coarse lamellar microstructure. The lamellar interfaces are parallel to the beam direction in the micrographs. a= 112( 0 type dislocations in a2 lamellae and cross twins in g lamellae are seen in Fig. 5(a) and (b), respectively. The a dislocations cause the soft modes of deformation, and the cross twins, the hard modes. Therefore, the plastic deformation in the coarse lamellar specimen is isotropic in a macroscopic scale. On the other hand, cross twins
K. Maruyama et al. / Materials Science and Engineering A319–321 (2001) 360–363
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with decreasing lamellar spacing, and the Hall–Petch relation holds between the yield stress and lamellar spacing when lamellar spacing is thicker than a critical value l*. 2. The yield stress saturates at a value s* below l*. s* is close to 1 GPa and l* is about 100 nm in the Ti-39.4mol%Al alloy. The pile-up model of dislocations can explain consistently the saturation stress, critical lamellar spacing and Hall–Petch slope. 3. Refinement of lamellar microstructure reduces fracture strain, since the hard deformation modes are suppressed in fine lamellar microstructures.
Acknowledgements This research was supported by JSPS (No.JSPSRFTF96R12301) and the Ministry of Education, Science, Sports and Culture, Japan (No.11450259). The material used was kindly provided by Kobe Steel Ltd. The authors would like to thank Mr G. Suzuki of Tohoku University for his help with mechanical testing.
References
Fig. 5. Deformation substructures formed in a thick lamellar microstructure (l= 590 nm) after 2.5% plastic deformation at 950 K. (a) g=202( 1a2, (b) g = 11( 1g.
were hardly seen in fine lamellar microstructures, suggesting that the plastic deformation takes place primarily by the soft deformation modes in the fine lamellar specimens. This anisotropic deformation builds up incompatibility stresses and eventually promotes the premature cracking of the fine lamellar microstructures.
4. Conclusions 1. Yield stress of fully lamellar TiAl alloys increases
[1] Y.-W. Kim, JOM 46 (7) (1994) 30. [2] K. Maruyama, R. Yamamoto, H. Nakakuki, N. Fujitsuna, Mater. Sci. Eng. A239-240 (1997) 419. [3] J. Beddoes, J. Triantafillou, L. Zhao, MRS Symp. Proc. 460 (1997) 293. [4] R. Yamamoto, K. Mizoguchi, G. Wegmann, K. Maruyama, Intermetallics 6 (1998) 699. [5] R. Yamamoto, K. Mizoguchi, G. Wegmann, K. Maruyama, in: Y.-W. Kim, et al. (Eds.), Gamma Titanium Aluminides 1999, TMS, Warrendale, 1999, p. 461. [6] P.J. Maziasz, C.T. Liu, Metall. Mater. Trans. 29A (1998) 105. [7] C.T. Liu, P.J. Maziasz, Intermetallics 6 (1998) 653. [8] D.M. Dimiduk, P.M. Hazzledine, T.A. Parthasarathy, S. Seshagiri, M.G. Mendiratta, Metall. Mater. Trans. 29A (1998) 37. [9] Y.Q. Sun, Mater. Sci. Eng. A239-240 (1997) 131. [10] N. Yamada, H. Sato, K. Maruyama, Report of the 123rd Committee on Heat-Resisting Metals and Alloys, JSPS 40 (1999) 273. [11] M.J. Blackburn, in: R.T. Jaffee, N.E. Promisel (Eds.), Science, Technology and Application of Titanium, Pergamon Press, London, 1970, p. 633. [12] Y. Umakoshi, T. Nakano, Acta Metall. Mater. 41 (1993) 1155.