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Effect of lamellar spacing on fatigue crack growth behaviour of a TiAl-based aluminide with lamellar microstructure
Materials Science and Engineering A 532 (2012) 13–20
Contents lists available at SciVerse ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Effect of lamellar spacing on fatigue crack growth behaviour of a TiAl-based aluminide with lamellar microstructure Yoji Mine a,∗ , Kazuki Takashima b , Paul Bowen c a
Department of Mechanical Engineering, Kyushu University, Moto-oka, Nishi-ku, Fukuoka 819-0395, Japan Department of Materials Science and Engineering, Kumamoto University, Kurokami, Kumamoto 860-8555, Japan Interdisciplinary Research Centre in Materials for High Performance Applications/School of Metallurgy and Materials, The University of Birmingham, Edgbaston, Birmingham B15 2TT, UK b c
a r t i c l e
i n f o
Article history: Received 21 March 2011 Received in revised form 11 October 2011 Accepted 15 October 2011 Available online 30 October 2011 Keywords: Mechanical characterization Intermetallics Titanium alloys Fatigue Interfaces
a b s t r a c t The fatigue crack growth resistance of a Ti–48Al–2Mn–2Nb (at.%) alloy with different lamellar spacings was investigated at room temperature in laboratory air. The crack growth resistance in the air-cooled specimen with a lamellar spacing of ∼0.9 m was essentially higher than that in the furnace-cooled specimen with a lamellar spacing of ∼3.8 m. The increased resistance is attributed to the zigzag crack growth across the fine lamellar plates in the air-cooled specimen. However, some furnace-cooled specimens actually exhibited superior fatigue crack growth resistance, comparable to that obtained by air cooling. In these specimens, evidence of crack branching and crack deflection, caused by interlamellar cracking away from the mode I crack growth direction, was frequently observed. Hence, the fatigue crack growth resistance can vary widely, especially for the furnace-cooled specimens of nominally similar fully lamellar microstructures. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Gamma (␥)-TiAl single phase alloys are essentially very brittle. However, two-phase alloys containing a small amount of ␣2 -Ti3 Al possess improved mechanical properties. Their mechanical properties depend strongly on their microstructure, which varies widely with heat-treatments [1]. The microstructures of the (␥ + ␣2 ) alloys can be roughly divided into lamellar and duplex microstructures. The lamellar microstructure has a good balance of elevated-temperature properties and fracture toughness. Microstructural features, such as lamellar orientation [2–4], colony size [5], colony boundary character [6], and plate thickness [3,7,8] affect the fracture toughness of alloys with lamellar microstructures. It is important to establish fatigue crack growth behaviour as well as fracture toughness. There have been a few studies on the effect of microstructure [9], lamellar orientation [10–12], and colony size [13,14] on fatigue crack growth behaviour. Meanwhile, fine lamellar structures of copper [15–17] have recently attracted great interest not only to simultaneously enhance strength and ductility but also to improve fatigue resistance. Similarly, refinement of lamellar structures of TiAl alloys is expected to provide improved fatigue characteristics.
∗ Corresponding author. Tel.: +81 92 802 3286; fax: +81 92 802 3065. E-mail address: [email protected] (Y. Mine). 0921-5093/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2011.10.055
In a previous investigation [13], the fatigue crack growth behaviour showed a dependency on the lamellar orientation rather than the lamellar colony size, and the fatigue crack growth resistance in air-cooled specimens was slightly higher than that in furnace-cooled specimens. In the current study, the fatigue crack growth resistance was measured using air- and furnace-cooled specimens with similar lamellar colony size, and the microstructures of the air- and furnace-cooled specimens were characterized by transmission electron microscopy (TEM). We discuss the influences of lamellar plate thickness and lamellar interface character on fatigue crack growth behaviour.
2. Materials and experimental methods The material used in this study was a Ti–48Al–2Mn–2Nb (at.%) alloy, produced by plasma arc remelting and subsequently casting at the Interdisciplinary Research Centre, University of Birmingham. Specimens with dimensions of 10 mm × 10 mm × 70 mm were cut by electrospark discharge machining from the as-cast ingot. Each specimen was wrapped in a tantalum sheet and encapsulated in a silica tube under a vacuum of 5 × 10−4 Pa. The specimens were heated to a temperature of 1473 K within the (␥ + ␣) phase region and maintained at this temperature for 129.6 ks to homogenize the microstructure after casting. When the specimens were further heated to a temperature of 1663 K within the ␣ phase region, the average lamellar colony size increased with an increase in the
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Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20 Table 1 Average lamellar plate thickness and spacing. Lamellar thickness (m)
Air-cooled specimen (A1) Furnace-cooled specimen (F1)
Fig. 1. Relation between average lamellar colony size and holding time at ␣ phase region during heat treatments.
holding time, as shown in Fig. 1. In the current study, the specimens held at this temperature for 1.8 ks were followed by air cooling and furnace cooling to room temperature. These heat treatments provided the fully lamellar microstructures with an average lamellar colony size of 0.90, 0.91, 0.97 and 1.00 mm for the air-cooled specimens, denoted as specimens A1, A2, A3 and A4, respectively, and 0.70, 0.77, 0.80 and 0.93 mm for the furnace-cooled specimens, denoted as specimens F1, F2, F3 and F4, respectively. Fig. 2a and b shows typical microstructures of the air- and furnace-cooled specimens, respectively. The side surfaces of the single-edge-notched (SEN) specimens were polished with emery paper up to #1000 and with alumina particles followed by etching with Krolls II reagent
␥
␣2
Mixed
0.15 0.51
0.03 0.11
0.13 0.47
Lamellar spacing (m)
0.91 3.84
(2% HF, 1% HNO3 and 97% H2 O). A through thickness notch of 2.5 mm depth was introduced in the centre of the specimens using a diamond saw of 300 m in thickness so that the initial notch length to width ratio, a0 /W, can be equal to 0.25. Here, these SEN specimens contained ten colonies or so through the thickness. Unlike usual polycrystalline specimens, therefore, the total fatigue crack growth resistance may reflect crack interfering effects with interlamellar cracking appeared over a few colonies. The thicknesses of the lamellar plates were measured with a JEOL JEM-2000FX transmission electron microscope operated at an accelerating voltage of 200 kV. The characters of the lamellar interfaces were identified from selected-area electron diffraction patterns and TEM dark-field images. For TEM, 3-mm diameter discs were thinned by grinding with emery paper followed by electropolishing for perforation at a temperature of 253 K using a twin jet electropolisher with a solution of 60% methanol, 36% n-butyl alcohol and 4% hydrogen peroxide. Fatigue crack growth tests were carried out at room temperature in laboratory air using a servo-hydraulic testing machine. Precracking was introduced at a frequency of 2 Hz and a stress ratio, R (the ratio of the minimum to maximum stress intensity factor applied over the fatigue cycle) of 0.1 under cyclic uniaxial compression, until the pre-crack extended approximately 250 m from the notch root. Fatigue crack growth tests were then performed at a frequency of 10 Hz and an R of 0.1 under three-point bending. The crack length was monitored by use of the direct current potential difference technique. Crack growth rates were measured under conditions of increasing applied stress intensity factor range, K, with crack extension, i.e., under a constant cyclic loading range. The tests were interrupted at a predetermined number of cycles to observe the crack profiles by optical microscopy. The fatigue surfaces of all the specimens after failure were examined by scanning electron microscopy (SEM). 3. Results 3.1. Lamellar microstructures of air- and furnace-cooled specimens
Fig. 2. Microstructures of (a) air- and (b) furnace-cooled specimens.
Fig. 3a and b shows the TEM bright-field images of typical microstructures of the air- and furnace-cooled specimens, respectively. Table 1 contains the average lamellar plate thicknesses of ␥ and ␣2 , and the average distance between the ␣2 plates, i.e., the average lamellar spacing, of the air- and furnace-cooled specimens. A decrease in the cooling rate during the heat treatments, i.e., from air cooling to furnace cooling, increased the lamellar plate thickness four-fold. From the measurement of the thickness and length of all the lamellae in Fig. 3, the volume fraction of the ␣2 phase was determined to be 3.9% for the air-cooled specimen and 2.3% for the furnace-cooled specimen. The specimens also exhibited different lamellar interface character distributions. Fig. 4 shows the orientations of the lamellar plates in the air- and furnace-cooled specimens, based on the six orientation variants of the ␥ plates with respect to the basal plane of the ␣2 plate, i.e., {1 1 1}␥ //(0 0 0 1)␣2 , reported in Ref. [18]. The lamellar interfaces can be classified into five boundary types; the ␥/␥ antiphase (0◦ -rotated) boundary, the ␥/␥ pseudo twin (60◦ -rotated) boundary, the ␥/␥ order fault
Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
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Fig. 3. Bright-field transmission electron micrographs of (a) air- and (b) furnace-cooled specimens.
Fig. 4. Schematic illustrations of lamellar plate orientations in (a) air- and (b) furnace-cooled specimens, corresponding to Fig. 3a and b, respectively, based on (c) six orientation variants of ␥ phase with respect to basal plane of ␣2 phase {1 1 1}␥ //(0 0 0 1)␣2 (after Zghal et al. [18]).
(120◦ -rotated) boundary, the ␥/␥ true twin (180◦ -rotated) boundary, and the ␥/␣2 interphase boundary. The fractions of these boundaries are summarized in Table 2. In the air-cooled specimen, the fraction of the ␥/␥ true twin boundaries was almost 50% and that of the ␥/␣2 interphase boundaries was 30%. A decrease in the cooling rate during the heat treatments decreased the ␥/␥ true twin and ␥/␣2 interphase boundary fractions but increased the ␥/␥ pseudo twin boundary fraction. The ratio of ␥/␥ true twin to ␥/␣2
interphase boundary was 1.64 for the air-cooled specimen and 1.33 for the furnace-cooled specimen. 3.2. Fatigue crack growth resistance Fig. 5 plots the fatigue crack growth rate, da/dN, against the applied stress intensity factor range, K. In Fig. 5a, the airand furnace-cooled specimens are denoted by open and closed
Table 2 Number and fraction of five possible lamellar boundaries.
Air-cooled specimen (A1)
Number Fraction (%)
␥/␥ antiphase (0◦ or 360◦ -rotated) 5 6.8
Furnace-cooled specimen (F1)
Number Fraction (%)
4 7.8
Type of boundary
␥/␥ pseudo twin (60◦ -rotated) 5 6.8
␥/␥ order fault (120◦ -rotated) 6 8.1
␥/␥ true twin (180◦ -rotated) 36 48.6
␥/␣2 interphase 22 29.7
12 23.5
7 13.8
16 31.4
12 23.5
16
Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
Fig. 5. Fatigue crack growth resistance curves: (a) air- and furnace-cooled specimens and (b) furnace-cooled specimens with different lamellar colony size.
symbols, respectively. The crack growth resistance in the air-cooled specimens was, in general, higher than that in the furnace-cooled specimens. However, the scattering of the crack growth resistances was larger in the furnace-cooled specimens than in the air-cooled specimens. The crack growth resistance was improved for specimens F3 and F4 compared to the other furnace-cooled specimens, and comparable to that of the air-cooled specimens. In particular, the crack growth rates of specimens F3 and F4 appear to decelerate √ at K values above 10 MPa m. Fig. 5b compares the fatigue crack growth resistances of the furnace-cooled specimens with different colony sizes in the range from 0.70 to 1.47 mm. Except for specimens F3 and F4, the fatigue crack growth resistances were nearly identical despite increasing colony size. Therefore, it is unlikely that the different colony sizes in the furnace-cooled specimens were responsible for the scattering of the fatigue crack growth resistance.
growth direction (Fig. 7a). Fig. 8a–d shows the crack growth process of specimen F3, which also exhibits superior resistance to crack growth. The crack deviated from the mode I crack growth direction caused by the interlamellar cracking (Fig. 8a) and then branched (Fig. 8b). This process was repeated until the crack went through the colony (Fig. 8c and d). Fig. 9 shows the fatigue surfaces of specimens F1, F3, and F4. The fractographic observation confirms the high tortuosity in specimens F3 and F4 (Fig. 9b and c) compared to the F1 (Fig. 9a). For example, flat regions significantly inclined to the mode I crack growth plane and secondary cracks were observed on the fatigue surfaces of specimens F3 and F4. Thus, the fatigue crack growth resistance can be improved by extrinsic toughening effects, depending on the misorientation between neighbouring colonies at the crack front. It must also be noted that, unlike the air-cooled specimens, the flat regions, in which few steps were observed, were rather smooth and featureless (Fig. 9d).
3.3. Fatigue crack profiles and fracture surfaces 4. Discussion Fig. 6 shows the fatigue crack profile and fracture surface of specimen A4, representative of the air-cooled specimens. It is found that the crack propagated both parallel to and perpendicular to the lamellar interfaces at the specimen surface (Fig. 6a). The fatigue crack extended roughly perpendicular to the first principal stress direction, i.e., parallel to the mode I crack growth direction. The fatigue surface was composed of pillar-shaped feature regions and flat regions, of sizes which approach the average lamellar colony size (Fig. 6b), corresponding to translamellar cracking and interlamellar cracking, respectively. On the flat regions marked in Fig. 6b, step-like markings in three directions with an angle of 120◦ between them were observed at high magnification (Fig. 6c). No significant evidence of extrinsic toughening, such as crack branching and crack deflection, was observed in the crack profiles and on the fatigue surfaces of the air-cooled specimens. The crack growth resistance curves in the furnace-cooled specimens were separated into two groups (Fig. 5). These could also be easily recognized in their crack profiles. Fig. 7a and b shows the crack profiles of specimens F1 and F4, respectively. The specimens with a tortuous crack path including crack deflection, branching and/or bridging (Fig. 7b) exhibited high resistance compared to those having a roughly straight crack path in the mode I crack
The lamellar plate thicknesses of both the ␥ and ␣2 and the lamellar spacing were increased with decreasing cooling rate during the heat treatments (Table 1). The fatigue crack growth resistance in the air-cooled specimens with thinner lamellar plates was often higher than that in the furnace-cooled specimens with thicker lamellar plates (Fig. 5). This observation can clearly be rationalized if the translamellar crack growth dominates the crack growth resistance of these microstructures. The fractographic observations indicate definite differences in the interlamellar fracture features between the air- and furnacecooled specimens (Figs. 6c and 9d). The interlamellar facets in the furnace-cooled specimens were composed of very smooth planes with few steps, whereas in the air-cooled specimens, step-like markings in three directions with a mutual angle of 120◦ were observed on the interlamellar fracture regions. This suggests that the interlamellar crack growth in the air-cooled specimens was associated with that on the {1 1 1} planes across the ␥ plates. Nakano et al. [3] carried out the fracture tests using polysynthetically twinned (PST) crystal compact-tension specimens with different lamellar spacings. According to this work, when a crack was introduced parallel to the lamellar plates, flat regions on
Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
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Fig. 6. Fatigue crack profile (a) and fracture surface of air-cooled specimen (specimen A4): (b) macroscopic view and (c) step-like markings in flat region. Nominal crack growth direction is from left to right.
(0 0 0 1)␣2 and steps with {1 1 1}␥ planes appeared on the fracture surface. In this case, a refinement of the lamellar spacing led to the increased number of steps. This tendency is consistent with the finding in the current study despite the loading mode: monotonic loading and cyclic loading.
The lamellar interfaces in the air-cooled specimen were characterized by ∼50% ␥/␥ twin boundaries and ∼30% ␥/␣2 interphase boundaries (Table 2). In the furnace-cooled specimen, the fractions of the ␥/␥ true twin and ␥/␣2 interphase boundaries decreased to ∼30% and ∼25%, respectively, while the ␥/␥ pseudo-twin
Fig. 7. Crack profile of furnace-cooled specimen: (a) F1 and (b) F4. Nominal crack growth direction is from left to right.
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Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
Fig. 8. Crack extension concurrent with crack deflection and branching in specimen F3. Nominal crack growth direction is from left to right.
Fig. 9. Scanning electron micrographs of fatigue surfaces of furnace-cooled specimens: macroscopic views of specimens (a) F1, (b) F3, and (c) F4, and (d) enlarged flat region in (a).
Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
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Fig. 10. Schematic illustrations of interlamellar crack growth mechanisms in (a) air- and (b) furnace-cooled specimens.
boundary fraction increased to ∼25%. The calculated cleavage energies of the (1 0 0) and (1 1 1) planes for the ␥ phase and the (0 0 0 1) plane for the ␣2 phase, and the interfacial fracture energies for each lamellar boundary were reported to be similar in the range of 4.5–4.8 J m−2 [19,20]. We could not identify the boundary that was most susceptible to interlamellar cracking. However, according to in situ TEM studies [20–23], the deformation transfer behaviour was closely dependent on the lamellar interface character. At the ␥/␥ lamellar boundaries, the stress concentration was most likely accommodated by deformation twinning or dislocation gliding. On the other hand, even if the primary slip system 1 1¯ 0 0 1 1 2¯ 0 can be activated, not enough slip systems in the 0 0 0 1 direction occur in the ␣2 plates with an ordered hexagonal structure. Therefore, it is more difficult for the deformation to progress through the ␥/␣2 interphase boundary. Moreover, a study on ␣2 single crystals indicated [24] that cleavage-like fractures can occur on the (00 0 1) and 1 1¯ 0 2 planes. It was also suggested that the 1 1¯ 0 2 fracture participants in the crack growth across the ␥ plates, resulting in the formation of multiple steps in specimens with thin lamellar plates. If the interlamellar cracking can occur preferentially in the ␣2 plates to accommodate the stress concentration at the ␥/␣2
boundaries, the distance between the ␣2 plates, i.e., the lamellar spacing, should be a crucial factor. Fig. 10 illustrates schematically the mechanism for the interlamellar crack growth in the air- and furnace-cooled specimens when the orientation of the lamellae is nearly parallel to the mode I crack growth direction. It is considered that, in the air-cooled specimen with a lamellar spacing of ∼0.9 m, the crack traverses the ␥ plates by linking of the microcracks formed in a short range ahead of the crack tip rather than propagating straight along the ␣2 plate (Fig. 10a). Meanwhile, the furnace-cooled specimen had a wider lamellar spacing of ∼3.8 m, which is in the range of the critical lamellar thickness of 3–5 m for forming microcracks [21]. Thus, the crack propagates mainly on planes parallel to the lamellar interfaces and occasionally across the lamellar plates to form only a few steps in the furnace-cooled specimens with greater lamellar spacings (Fig. 10b). The observation of more steps when the lamellar plate spacing is refined may provide a mechanism for the increased crack growth resistance of the air-cooled specimens. The fatigue crack growth resistances of specimens F3 and F4 were higher than those of the other furnace-cooled specimens (Fig. 5). The crack in specimens F1 and F2 propagated roughly parallel to the mode I crack growth direction (Fig. 7a). In contrast, the
Fig. 11. Relation between crack growth and lamellar direction in (a) air- and (b) furnace-cooled specimens.
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Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20
crack growth behaviour of specimens F3 and F4 gives evidence of crack deflection, branching and/or bridging, caused by interlamellar cracking away from the mode I crack growth direction (Figs. 7b and 8). Fig. 11 illustrates the relation between crack growth and lamellar colony misorientation in the air- and furnace-cooled specimens. In the furnace-cooled specimens, which are susceptible to interlamellar cracking, the fatigue crack growth rates are accelerated when the orientation of the lamellae is aligned parallel to the mode I crack growth direction. Interlamellar cracking away from the mode I crack growth direction, which depends on the misorientation between neighbouring colonies at the crack front, results in crack deflection, branching, and ligament bridging, and an improvement in the fatigue crack growth resistance. Therefore, the fatigue crack growth resistance can vary widely, particularly in the furnace-cooled specimens. 5. Conclusions The effect of lamellar microstructural features on the fatigue crack growth resistance of Ti–48Al–2Mn–2Nb (at.%), composed of the ␥ and ␣2 phases, was investigated at room temperature in laboratory air. The following conclusions have been made: (1) The lamellar thickness and lamellar spacing were increased but the ␣2 phase fraction decreased with the decrease in cooling rate from air cooling to furnace cooling. (2) The fatigue crack growth resistance in the air-cooled specimen with smaller lamellar spacings was essentially higher than that in the furnace-cooled specimens with greater lamellar spacings. In the air-cooled specimens, the crack zigzagged across the fine lamellar plates caused by the linkage of smaller-spaced microcracking, even when the crack grew nominally parallel to the lamellar plates. (3) The fatigue crack growth resistance in some of the furnacecooled specimen was improved by extrinsic toughening effect because of the favourable misorientation between neighbouring colonies at the crack front.
Acknowledgments The authors would like to extend their gratitude to Prof. M. Nishida of Kyushu University for insightful discussions and valuable comments on the TEM observations. The authors thank Mr. Y. Morofuji, graduate student of Kumamoto University, for his assistance with the experimental program. References [1] Y.-W. Kim, Acta Metall. Mater. 40 (1992) 1121–1134. [2] S. Mitao, T. Isawa, S. Tsuyama, Scr. Metall. Mater. 26 (1992) 1405–1410. [3] T. Nakano, T. Kawanaka, H.Y. Yasuda, Y. Umakoshi, Mater. Sci. Eng. A 194 (1995) 43–51. [4] S. Yokoshima, M. Yamaguchi, Acta Mater. 44 (1996) 873–883. [5] K.S. Chan, Y.-W. Kim, Metall. Trans. A 23 (1992) 1663–1677. [6] P. Wang, N. Bhate, K.S. Chan, K.S. Kumar, Acta Mater. 51 (2003) 1573–1591. [7] K.S. Chan, Y-W. Kim, Acta Metall. Mater. 43 (1995) 439–451. [8] N.J. Rogers, P.D. Crofts, I.P. Jones, P. Bowen, Mater. Sci. Eng. A 192–193 (1995) 379–386. [9] K.T. Venkateswara Rao, Y.W. Kim, C.L. Muhlstein, R.O. Ritchie, Mater. Sci. Eng. A 192–193 (1995) 474–482. [10] P. Bowen, R.A. Chave, A.W. James, Mater. Sci. Eng. A 192–193 (1995) 443–456. [11] R. Gnanamoorthy, Y. Mutoh, K. Hayashi, Y. Mizuhara, Scr. Metall. Mater. 33 (1995) 907–912. [12] H. Shiota, K. Tokaji, Y. Ohta, Mater. Sci. Eng. A 243 (1998) 169–175. [13] Y. Mine, Y. Morofuji, S. Ando, K. Takashima, H. Tonda, J. Soc. Mater. Sci. Jpn. 46 (1997) 1167–1172 (in Japanese). [14] C. Choi, H.J. Kim, Y.-T. Lee, Y.-W. Kim, C.S. Lee, Mater. Sci. Eng. A 329–331 (2002) 545–556. [15] L. Lu, Y.F. Shen, X.H. Chen, L.H. Qian, K. Lu, Science 304 (2004) 422–426. [16] Y.F. Shen, L. Lu, Q.H. Lu, Z.H. Jin, K. Lu, Scr. Mater. 52 (2005) 989–994. [17] A. Singh, L. Tang, M. Dao, L. Lu, S. Suresh, Acta Mater. 59 (2011) 2437–2446. [18] S. Zghal, S. Naka, A. Couret, Acta Mater. 45 (1997) 3005–3015. [19] M.H. Yoo, C.L. Fu, Metall. Trans. A 29 (1998) 49–63. [20] M. Chen, D. Lin, D. Chen, Scr. Mater. 36 (1997) 497–501. [21] D. Lin, M. Chen, Interaction between deformation process and lamellar interfaces in polysynthetically twinned (PST) crystals of TiAl, in: Y.-W. Kim, R. Wagner, M. Yamaguchi (Eds.), Gamma Titanium Aluminides, TMS, 1995, pp. 323–330. [22] D. Hu, M.H. Loretto, Intermetallics 7 (1999) 1299–1306. [23] Y.H. Lu, Y.G. Zhang, L.J. Qiao, Y.-B. Wang, C.Q. Chen, W.Y. Chu, Mater. Sci. Eng. A 289 (2000) 91–98. [24] Y. Umakoshi, T. Nakano, B. Ogawa, Scr. Mater. 34 (1996) 1161–1169.
Contents lists available at SciVerse ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Effect of lamellar spacing on fatigue crack growth behaviour of a TiAl-based aluminide with lamellar microstructure Yoji Mine a,∗ , Kazuki Takashima b , Paul Bowen c a
Department of Mechanical Engineering, Kyushu University, Moto-oka, Nishi-ku, Fukuoka 819-0395, Japan Department of Materials Science and Engineering, Kumamoto University, Kurokami, Kumamoto 860-8555, Japan Interdisciplinary Research Centre in Materials for High Performance Applications/School of Metallurgy and Materials, The University of Birmingham, Edgbaston, Birmingham B15 2TT, UK b c
a r t i c l e
i n f o
Article history: Received 21 March 2011 Received in revised form 11 October 2011 Accepted 15 October 2011 Available online 30 October 2011 Keywords: Mechanical characterization Intermetallics Titanium alloys Fatigue Interfaces
a b s t r a c t The fatigue crack growth resistance of a Ti–48Al–2Mn–2Nb (at.%) alloy with different lamellar spacings was investigated at room temperature in laboratory air. The crack growth resistance in the air-cooled specimen with a lamellar spacing of ∼0.9 m was essentially higher than that in the furnace-cooled specimen with a lamellar spacing of ∼3.8 m. The increased resistance is attributed to the zigzag crack growth across the fine lamellar plates in the air-cooled specimen. However, some furnace-cooled specimens actually exhibited superior fatigue crack growth resistance, comparable to that obtained by air cooling. In these specimens, evidence of crack branching and crack deflection, caused by interlamellar cracking away from the mode I crack growth direction, was frequently observed. Hence, the fatigue crack growth resistance can vary widely, especially for the furnace-cooled specimens of nominally similar fully lamellar microstructures. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Gamma (␥)-TiAl single phase alloys are essentially very brittle. However, two-phase alloys containing a small amount of ␣2 -Ti3 Al possess improved mechanical properties. Their mechanical properties depend strongly on their microstructure, which varies widely with heat-treatments [1]. The microstructures of the (␥ + ␣2 ) alloys can be roughly divided into lamellar and duplex microstructures. The lamellar microstructure has a good balance of elevated-temperature properties and fracture toughness. Microstructural features, such as lamellar orientation [2–4], colony size [5], colony boundary character [6], and plate thickness [3,7,8] affect the fracture toughness of alloys with lamellar microstructures. It is important to establish fatigue crack growth behaviour as well as fracture toughness. There have been a few studies on the effect of microstructure [9], lamellar orientation [10–12], and colony size [13,14] on fatigue crack growth behaviour. Meanwhile, fine lamellar structures of copper [15–17] have recently attracted great interest not only to simultaneously enhance strength and ductility but also to improve fatigue resistance. Similarly, refinement of lamellar structures of TiAl alloys is expected to provide improved fatigue characteristics.
∗ Corresponding author. Tel.: +81 92 802 3286; fax: +81 92 802 3065. E-mail address: [email protected] (Y. Mine). 0921-5093/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2011.10.055
In a previous investigation [13], the fatigue crack growth behaviour showed a dependency on the lamellar orientation rather than the lamellar colony size, and the fatigue crack growth resistance in air-cooled specimens was slightly higher than that in furnace-cooled specimens. In the current study, the fatigue crack growth resistance was measured using air- and furnace-cooled specimens with similar lamellar colony size, and the microstructures of the air- and furnace-cooled specimens were characterized by transmission electron microscopy (TEM). We discuss the influences of lamellar plate thickness and lamellar interface character on fatigue crack growth behaviour.
2. Materials and experimental methods The material used in this study was a Ti–48Al–2Mn–2Nb (at.%) alloy, produced by plasma arc remelting and subsequently casting at the Interdisciplinary Research Centre, University of Birmingham. Specimens with dimensions of 10 mm × 10 mm × 70 mm were cut by electrospark discharge machining from the as-cast ingot. Each specimen was wrapped in a tantalum sheet and encapsulated in a silica tube under a vacuum of 5 × 10−4 Pa. The specimens were heated to a temperature of 1473 K within the (␥ + ␣) phase region and maintained at this temperature for 129.6 ks to homogenize the microstructure after casting. When the specimens were further heated to a temperature of 1663 K within the ␣ phase region, the average lamellar colony size increased with an increase in the
14
Y. Mine et al. / Materials Science and Engineering A 532 (2012) 13–20 Table 1 Average lamellar plate thickness and spacing. Lamellar thickness (m)
Air-cooled specimen (A1) Furnace-cooled specimen (F1)
Fig. 1. Relation between average lamellar colony size and holding time at ␣ phase region during heat treatments.
holding time, as shown in Fig. 1. In the current study, the specimens held at this temperature for 1.8 ks were followed by air cooling and furnace cooling to room temperature. These heat treatments provided the fully lamellar microstructures with an average lamellar colony size of 0.90, 0.91, 0.97 and 1.00 mm for the air-cooled specimens, denoted as specimens A1, A2, A3 and A4, respectively, and 0.70, 0.77, 0.80 and 0.93 mm for the furnace-cooled specimens, denoted as specimens F1, F2, F3 and F4, respectively. Fig. 2a and b shows typical microstructures of the air- and furnace-cooled specimens, respectively. The side surfaces of the single-edge-notched (SEN) specimens were polished with emery paper up to #1000 and with alumina particles followed by etching with Krolls II reagent
␥
␣2
Mixed
0.15 0.51
0.03 0.11
0.13 0.47
Lamellar spacing (m)
0.91 3.84
(2% HF, 1% HNO3 and 97% H2 O). A through thickness notch of 2.5 mm depth was introduced in the centre of the specimens using a diamond saw of 300 m in thickness so that the initial notch length to width ratio, a0 /W, can be equal to 0.25. Here, these SEN specimens contained ten colonies or so through the thickness. Unlike usual polycrystalline specimens, therefore, the total fatigue crack growth resistance may reflect crack interfering effects with interlamellar cracking appeared over a few colonies. The thicknesses of the lamellar plates were measured with a JEOL JEM-2000FX transmission electron microscope operated at an accelerating voltage of 200 kV. The characters of the lamellar interfaces were identified from selected-area electron diffraction patterns and TEM dark-field images. For TEM, 3-mm diameter discs were thinned by grinding with emery paper followed by electropolishing for perforation at a temperature of 253 K using a twin jet electropolisher with a solution of 60% methanol, 36% n-butyl alcohol and 4% hydrogen peroxide. Fatigue crack growth tests were carried out at room temperature in laboratory air using a servo-hydraulic testing machine. Precracking was introduced at a frequency of 2 Hz and a stress ratio, R (the ratio of the minimum to maximum stress intensity factor applied over the fatigue cycle) of 0.1 under cyclic uniaxial compression, until the pre-crack extended approximately 250 m from the notch root. Fatigue crack growth tests were then performed at a frequency of 10 Hz and an R of 0.1 under three-point bending. The crack length was monitored by use of the direct current potential difference technique. Crack growth rates were measured under conditions of increasing applied stress intensity factor range, K, with crack extension, i.e., under a constant cyclic loading range. The tests were interrupted at a predetermined number of cycles to observe the crack profiles by optical microscopy. The fatigue surfaces of all the specimens after failure were examined by scanning electron microscopy (SEM). 3. Results 3.1. Lamellar microstructures of air- and furnace-cooled specimens
Fig. 2. Microstructures of (a) air- and (b) furnace-cooled specimens.
Fig. 3a and b shows the TEM bright-field images of typical microstructures of the air- and furnace-cooled specimens, respectively. Table 1 contains the average lamellar plate thicknesses of ␥ and ␣2 , and the average distance between the ␣2 plates, i.e., the average lamellar spacing, of the air- and furnace-cooled specimens. A decrease in the cooling rate during the heat treatments, i.e., from air cooling to furnace cooling, increased the lamellar plate thickness four-fold. From the measurement of the thickness and length of all the lamellae in Fig. 3, the volume fraction of the ␣2 phase was determined to be 3.9% for the air-cooled specimen and 2.3% for the furnace-cooled specimen. The specimens also exhibited different lamellar interface character distributions. Fig. 4 shows the orientations of the lamellar plates in the air- and furnace-cooled specimens, based on the six orientation variants of the ␥ plates with respect to the basal plane of the ␣2 plate, i.e., {1 1 1}␥ //(0 0 0 1)␣2 , reported in Ref. [18]. The lamellar interfaces can be classified into five boundary types; the ␥/␥ antiphase (0◦ -rotated) boundary, the ␥/␥ pseudo twin (60◦ -rotated) boundary, the ␥/␥ order fault
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Fig. 3. Bright-field transmission electron micrographs of (a) air- and (b) furnace-cooled specimens.
Fig. 4. Schematic illustrations of lamellar plate orientations in (a) air- and (b) furnace-cooled specimens, corresponding to Fig. 3a and b, respectively, based on (c) six orientation variants of ␥ phase with respect to basal plane of ␣2 phase {1 1 1}␥ //(0 0 0 1)␣2 (after Zghal et al. [18]).
(120◦ -rotated) boundary, the ␥/␥ true twin (180◦ -rotated) boundary, and the ␥/␣2 interphase boundary. The fractions of these boundaries are summarized in Table 2. In the air-cooled specimen, the fraction of the ␥/␥ true twin boundaries was almost 50% and that of the ␥/␣2 interphase boundaries was 30%. A decrease in the cooling rate during the heat treatments decreased the ␥/␥ true twin and ␥/␣2 interphase boundary fractions but increased the ␥/␥ pseudo twin boundary fraction. The ratio of ␥/␥ true twin to ␥/␣2
interphase boundary was 1.64 for the air-cooled specimen and 1.33 for the furnace-cooled specimen. 3.2. Fatigue crack growth resistance Fig. 5 plots the fatigue crack growth rate, da/dN, against the applied stress intensity factor range, K. In Fig. 5a, the airand furnace-cooled specimens are denoted by open and closed
Table 2 Number and fraction of five possible lamellar boundaries.
Air-cooled specimen (A1)
Number Fraction (%)
␥/␥ antiphase (0◦ or 360◦ -rotated) 5 6.8
Furnace-cooled specimen (F1)
Number Fraction (%)
4 7.8
Type of boundary
␥/␥ pseudo twin (60◦ -rotated) 5 6.8
␥/␥ order fault (120◦ -rotated) 6 8.1
␥/␥ true twin (180◦ -rotated) 36 48.6
␥/␣2 interphase 22 29.7
12 23.5
7 13.8
16 31.4
12 23.5
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Fig. 5. Fatigue crack growth resistance curves: (a) air- and furnace-cooled specimens and (b) furnace-cooled specimens with different lamellar colony size.
symbols, respectively. The crack growth resistance in the air-cooled specimens was, in general, higher than that in the furnace-cooled specimens. However, the scattering of the crack growth resistances was larger in the furnace-cooled specimens than in the air-cooled specimens. The crack growth resistance was improved for specimens F3 and F4 compared to the other furnace-cooled specimens, and comparable to that of the air-cooled specimens. In particular, the crack growth rates of specimens F3 and F4 appear to decelerate √ at K values above 10 MPa m. Fig. 5b compares the fatigue crack growth resistances of the furnace-cooled specimens with different colony sizes in the range from 0.70 to 1.47 mm. Except for specimens F3 and F4, the fatigue crack growth resistances were nearly identical despite increasing colony size. Therefore, it is unlikely that the different colony sizes in the furnace-cooled specimens were responsible for the scattering of the fatigue crack growth resistance.
growth direction (Fig. 7a). Fig. 8a–d shows the crack growth process of specimen F3, which also exhibits superior resistance to crack growth. The crack deviated from the mode I crack growth direction caused by the interlamellar cracking (Fig. 8a) and then branched (Fig. 8b). This process was repeated until the crack went through the colony (Fig. 8c and d). Fig. 9 shows the fatigue surfaces of specimens F1, F3, and F4. The fractographic observation confirms the high tortuosity in specimens F3 and F4 (Fig. 9b and c) compared to the F1 (Fig. 9a). For example, flat regions significantly inclined to the mode I crack growth plane and secondary cracks were observed on the fatigue surfaces of specimens F3 and F4. Thus, the fatigue crack growth resistance can be improved by extrinsic toughening effects, depending on the misorientation between neighbouring colonies at the crack front. It must also be noted that, unlike the air-cooled specimens, the flat regions, in which few steps were observed, were rather smooth and featureless (Fig. 9d).
3.3. Fatigue crack profiles and fracture surfaces 4. Discussion Fig. 6 shows the fatigue crack profile and fracture surface of specimen A4, representative of the air-cooled specimens. It is found that the crack propagated both parallel to and perpendicular to the lamellar interfaces at the specimen surface (Fig. 6a). The fatigue crack extended roughly perpendicular to the first principal stress direction, i.e., parallel to the mode I crack growth direction. The fatigue surface was composed of pillar-shaped feature regions and flat regions, of sizes which approach the average lamellar colony size (Fig. 6b), corresponding to translamellar cracking and interlamellar cracking, respectively. On the flat regions marked in Fig. 6b, step-like markings in three directions with an angle of 120◦ between them were observed at high magnification (Fig. 6c). No significant evidence of extrinsic toughening, such as crack branching and crack deflection, was observed in the crack profiles and on the fatigue surfaces of the air-cooled specimens. The crack growth resistance curves in the furnace-cooled specimens were separated into two groups (Fig. 5). These could also be easily recognized in their crack profiles. Fig. 7a and b shows the crack profiles of specimens F1 and F4, respectively. The specimens with a tortuous crack path including crack deflection, branching and/or bridging (Fig. 7b) exhibited high resistance compared to those having a roughly straight crack path in the mode I crack
The lamellar plate thicknesses of both the ␥ and ␣2 and the lamellar spacing were increased with decreasing cooling rate during the heat treatments (Table 1). The fatigue crack growth resistance in the air-cooled specimens with thinner lamellar plates was often higher than that in the furnace-cooled specimens with thicker lamellar plates (Fig. 5). This observation can clearly be rationalized if the translamellar crack growth dominates the crack growth resistance of these microstructures. The fractographic observations indicate definite differences in the interlamellar fracture features between the air- and furnacecooled specimens (Figs. 6c and 9d). The interlamellar facets in the furnace-cooled specimens were composed of very smooth planes with few steps, whereas in the air-cooled specimens, step-like markings in three directions with a mutual angle of 120◦ were observed on the interlamellar fracture regions. This suggests that the interlamellar crack growth in the air-cooled specimens was associated with that on the {1 1 1} planes across the ␥ plates. Nakano et al. [3] carried out the fracture tests using polysynthetically twinned (PST) crystal compact-tension specimens with different lamellar spacings. According to this work, when a crack was introduced parallel to the lamellar plates, flat regions on
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Fig. 6. Fatigue crack profile (a) and fracture surface of air-cooled specimen (specimen A4): (b) macroscopic view and (c) step-like markings in flat region. Nominal crack growth direction is from left to right.
(0 0 0 1)␣2 and steps with {1 1 1}␥ planes appeared on the fracture surface. In this case, a refinement of the lamellar spacing led to the increased number of steps. This tendency is consistent with the finding in the current study despite the loading mode: monotonic loading and cyclic loading.
The lamellar interfaces in the air-cooled specimen were characterized by ∼50% ␥/␥ twin boundaries and ∼30% ␥/␣2 interphase boundaries (Table 2). In the furnace-cooled specimen, the fractions of the ␥/␥ true twin and ␥/␣2 interphase boundaries decreased to ∼30% and ∼25%, respectively, while the ␥/␥ pseudo-twin
Fig. 7. Crack profile of furnace-cooled specimen: (a) F1 and (b) F4. Nominal crack growth direction is from left to right.
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Fig. 8. Crack extension concurrent with crack deflection and branching in specimen F3. Nominal crack growth direction is from left to right.
Fig. 9. Scanning electron micrographs of fatigue surfaces of furnace-cooled specimens: macroscopic views of specimens (a) F1, (b) F3, and (c) F4, and (d) enlarged flat region in (a).
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Fig. 10. Schematic illustrations of interlamellar crack growth mechanisms in (a) air- and (b) furnace-cooled specimens.
boundary fraction increased to ∼25%. The calculated cleavage energies of the (1 0 0) and (1 1 1) planes for the ␥ phase and the (0 0 0 1) plane for the ␣2 phase, and the interfacial fracture energies for each lamellar boundary were reported to be similar in the range of 4.5–4.8 J m−2 [19,20]. We could not identify the boundary that was most susceptible to interlamellar cracking. However, according to in situ TEM studies [20–23], the deformation transfer behaviour was closely dependent on the lamellar interface character. At the ␥/␥ lamellar boundaries, the stress concentration was most likely accommodated by deformation twinning or dislocation gliding. On the other hand, even if the primary slip system 1 1¯ 0 0 1 1 2¯ 0 can be activated, not enough slip systems in the 0 0 0 1 direction occur in the ␣2 plates with an ordered hexagonal structure. Therefore, it is more difficult for the deformation to progress through the ␥/␣2 interphase boundary. Moreover, a study on ␣2 single crystals indicated [24] that cleavage-like fractures can occur on the (00 0 1) and 1 1¯ 0 2 planes. It was also suggested that the 1 1¯ 0 2 fracture participants in the crack growth across the ␥ plates, resulting in the formation of multiple steps in specimens with thin lamellar plates. If the interlamellar cracking can occur preferentially in the ␣2 plates to accommodate the stress concentration at the ␥/␣2
boundaries, the distance between the ␣2 plates, i.e., the lamellar spacing, should be a crucial factor. Fig. 10 illustrates schematically the mechanism for the interlamellar crack growth in the air- and furnace-cooled specimens when the orientation of the lamellae is nearly parallel to the mode I crack growth direction. It is considered that, in the air-cooled specimen with a lamellar spacing of ∼0.9 m, the crack traverses the ␥ plates by linking of the microcracks formed in a short range ahead of the crack tip rather than propagating straight along the ␣2 plate (Fig. 10a). Meanwhile, the furnace-cooled specimen had a wider lamellar spacing of ∼3.8 m, which is in the range of the critical lamellar thickness of 3–5 m for forming microcracks [21]. Thus, the crack propagates mainly on planes parallel to the lamellar interfaces and occasionally across the lamellar plates to form only a few steps in the furnace-cooled specimens with greater lamellar spacings (Fig. 10b). The observation of more steps when the lamellar plate spacing is refined may provide a mechanism for the increased crack growth resistance of the air-cooled specimens. The fatigue crack growth resistances of specimens F3 and F4 were higher than those of the other furnace-cooled specimens (Fig. 5). The crack in specimens F1 and F2 propagated roughly parallel to the mode I crack growth direction (Fig. 7a). In contrast, the
Fig. 11. Relation between crack growth and lamellar direction in (a) air- and (b) furnace-cooled specimens.
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crack growth behaviour of specimens F3 and F4 gives evidence of crack deflection, branching and/or bridging, caused by interlamellar cracking away from the mode I crack growth direction (Figs. 7b and 8). Fig. 11 illustrates the relation between crack growth and lamellar colony misorientation in the air- and furnace-cooled specimens. In the furnace-cooled specimens, which are susceptible to interlamellar cracking, the fatigue crack growth rates are accelerated when the orientation of the lamellae is aligned parallel to the mode I crack growth direction. Interlamellar cracking away from the mode I crack growth direction, which depends on the misorientation between neighbouring colonies at the crack front, results in crack deflection, branching, and ligament bridging, and an improvement in the fatigue crack growth resistance. Therefore, the fatigue crack growth resistance can vary widely, particularly in the furnace-cooled specimens. 5. Conclusions The effect of lamellar microstructural features on the fatigue crack growth resistance of Ti–48Al–2Mn–2Nb (at.%), composed of the ␥ and ␣2 phases, was investigated at room temperature in laboratory air. The following conclusions have been made: (1) The lamellar thickness and lamellar spacing were increased but the ␣2 phase fraction decreased with the decrease in cooling rate from air cooling to furnace cooling. (2) The fatigue crack growth resistance in the air-cooled specimen with smaller lamellar spacings was essentially higher than that in the furnace-cooled specimens with greater lamellar spacings. In the air-cooled specimens, the crack zigzagged across the fine lamellar plates caused by the linkage of smaller-spaced microcracking, even when the crack grew nominally parallel to the lamellar plates. (3) The fatigue crack growth resistance in some of the furnacecooled specimen was improved by extrinsic toughening effect because of the favourable misorientation between neighbouring colonies at the crack front.
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