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Study on compression deformation, damage and fracture behavior of TiAl alloys
Materials Science and Engineering A 527 (2010) 2468–2477
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Study on compression deformation, damage and fracture behavior of TiAl alloys Part II. Fracture behavior R. Cao a,b,∗ , L. Li a,b , J.H. Chen a,b , J. Zhang c a b c
State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials, Lanzhou University of Technology, Lanzhou 730050, China Key Laboratory of Non-ferrous Metal Alloys, The Ministry of Education, Lanzhou University of Technology, Lanzhou 730050, China Central Iron and Steel Research Institute, Beijing 100081, China
a r t i c l e
i n f o
Article history: Received 24 April 2008 Received in revised form 27 November 2008 Accepted 4 December 2008
Keywords: TiAl alloys Compression tests Fracture behavior Sensitivity to strain rate
a b s t r a c t Compression deformation, damage and fracture behaviors of TiAl alloys are studied systematically. The deformation and damage processes in compression tests of TiAl specimens are detailed in Part I and the fracture behaviors are described and discussed here the Part II of this work. Results indicate that: the compression properties are superior to the tensile properties and specially show a much higher fracture strength and plasticity. The macroscopic features of the fracture surfaces indicate that the fracture mechanisms are different between the full lamellar (FL) alloy and the duplex (DP) alloy. In FL alloy, the shear cracking dominates the fracture, and the shear crack extends throughout the whole specimen. At both ends of the specimen, the shear cracks initiate and propagate in directions inclined to the compression axis with an angle near 45◦ . In the middle of the specimen, the shear cracks propagate parallel to the compression axis. While in the DP alloy after two short shear cracks initiate and extend at both ends, the longitudinal normal cracking is triggered and propagates parallel to the compression axis and dominates the fracture. In the microscopic features the regions of shear cracking show the smooth interlamellar shear fracture facets with some translamellar shear fracture steps. The longitudinal normal cracking in DP specimens shows mixed interlamellar and translamellar cleavage fracture facets. Effect of loading rate manifests itself in increasing the shear yield strength of FL and the fracture strength of DP alloys. © 2009 Published by Elsevier B.V.
1. Introduction In the past two decades, ␥-TiAl-based alloys have received tremendous attention for high-temperature and high-performance structural applications, primarily due to their low density, high strength to weight ratio as well as their good creep and oxidation resistance [1,2]. However, as intermetallic compounds, some of their properties such as the toughness, formability at room temperature are much poorer than those of conventional structural materials therefore their applications are limited. Until now, a great number of studies on the fracture mechanism of TiAl alloys in tensile have been done and a series of systematic results have been achieved [3–5]. However, the reports on compression deformation, damage and fracture behavior are relatively rare. Ref. [6] found that fracture was caused by the growth and connection of grain boundary cracks. The ␣2 phase itself acted as an obstacle to prevent the crack propagation at the cross-points with
∗ Corresponding author at: State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials, Lanzhou University of Technology, Lanzhou 730050, China. Tel.: +86 931 2973529; fax: +86 931 2755806. E-mail addresses: [email protected], [email protected] (R. Cao). 0921-5093/$ – see front matter © 2009 Published by Elsevier B.V. doi:10.1016/j.msea.2008.12.012
the cracking by causing kinking and necking. Ref. [7] shows that there were three modes of crack initiation and propagation in the compression loading tests. As shown in Fig. 1(a) cracks initiate and propagate parallel to the compression axis. In Fig. 1(b) three cracks initiate and propagate at the intersection of three grain boundaries and one crack propagates along the compression axis, the other two cracks extend with an angle inclined to the compression axis. In Fig. 1(c) two cracks along compression axis direction initiate and propagate at both ends of an inclined precrack, which is produced at an inclined grain boundary. At the compression stress regions, microcracks nucleate at the precracks, cavities, inclusions and other defects. With applied load increasing, some cracks can propagate to the main crack or propagate and connect with each other and finally induce the specimen fracture. Ref. [8] shows that the fracture strength, the fracture strain and the cracking behavior were very sensitive to strain rate. At a strain rate of 5 × 10−3 s−1 (quasistatic condition), the fully lamellar samples showed large cracks in the direction of the compression cone, whereas in the dynamical conditions ((2–4) × 103 s−1 ), compression specimens exhibited a great number of short interlamellar microcracks. The high strain rate limited the length of the crack propagation. Therefore, in the quasi-static conditions, cracks grew to macroscopic crack, however at the dynamic deformation, few numbers of microcracks were
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Fig. 1. Three types of crack initiation and propagation in the compression loading mode [7].
Table 1 Compositions of TiAl alloy (at%). Ti Al V Cr
Balance 47.5 2.5 1.0
barely to grow to the macroscopic cracks. In the engineering applications, compression effects appear on blades of the combustion gas turbine by continuous start–shut down of the compressible gas. In case the mechanical connection is made by inserting the shaft of the TiAl wheel into the steel sleeve with the interference fit, the extra high compression stress maybe causes the cracking of the TiAl shaft. Therefore, the compression deformation and fracture behavior of TiAl alloys deserve to be studied for correct design and production of many engineering components with perfect performance. This is just the aim, which this work focuses on. 2. Experimental
Fig. 3. Dimensions of compression test specimen.
The average colony (grain) size and the average lath width of FL microstructure (Fig. 2a) were measured as 190–300 m and 2–3 m, respectively. The average grain size of the DP microstructure (Fig. 2b) was 20–30 m. The cubic compression specimens (12 mm × 5 mm × 5 mm) are used. The dimensions are shown in Fig. 3. 2.2. Compression tests
2.1. Material and specimens A TiAl alloy with compositions shown in Table 1 was used. All samples were taken from a forged pancake that has been deformed at 1100 ◦ C for a 70% height reduction. Two types of microstructures, duplex (DP) and near fully lamellar (FL) as shown in Fig. 2, were obtained. The samples were first wrapped in quartz tubes, treated by hot iso-static pressing at 120 MPa and 950 ◦ C in argon for 3 h and then put into the furnace at pre-determined temperatures. DP microstructure samples were obtained by annealing at 1250 ◦ C for 18 h and the FL microstructure samples by annealing at 1370 ◦ C for 1 h.
At room temperature compression tests were conducted in air by SHIMADSU AG-10TA universal test machine at strain rate of 1 × 10−3 s−1 . Engineering stress–strain curves were plotted in accord with the test data for all fractured specimens. 2.3. Compression tests at various strain rates Compression tests were done at various strain rates in range of 1 × 10−5 s−1 to 4 × 10−2 s−1 . The yield stress ( 0.2 ), ultimate compression stress ( uc ), fracture work of unit area (wf ) were measured with the engineering stress–strain curves of compression tests.
Fig. 2. Microstructures of (a) fully lamellar and (b) duplex TiAl-based alloy.
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Table 2 Mechanical properties measured in compression and tensile tests at strain rate of 1 × 10−3 s−1 for FL specimens.
Table 3 Mechanical properties measured in compression and tensile tests at strain rate of 1 × 10−3 s−1 for DP specimens.
Testing type
0.2 (MPa)
ut (MPa)
εmax (%)
f (MPa)
εf (%)
Testing type
0.2 (MPa)
ut (MPa)
εmax (%)
f (MPa)
εf (%)
Tensile Compression
484 484
524 2263
0.686 28.3
524 2144
0.686 31.4
Tensile Compression
325 379
409 2227
1.351 30.9
409 1868
1.351 31.2
Note: 0.2 : yield stress, ut : ultimate tensile stress, εmax : strain at ultimate tensile stress, f : fracture stress, εf : fracture strain.
Parameters of macroscopic mechanical properties measured at various strain rates were compared. 2.4. Observation on the macro- and micro-features of fracture surfaces The macro-features of the fracture surfaces of specimens fractured at different strain rates were observed by magnifying glass and compared in details for both FL and DP alloys. The crack configurations on the side surfaces of the specimens were also observed to identify the propagation process and the macro-features of the cracks. The micro-features of the fracture surfaces were observed by SEM and were related to the macro-features of the fracture surfaces. 2.5. FEM calculation A three-dimensional FEM model having eight-node threedimensional stress reduced integration elements C3D8R was used with the ABAQUS code. A total 35,175 elements and 39,938 nodes were taken in to account. The distributions of normal stress 11 and shear stress along a 45◦ -inclined cross-section in the specimens were calculated at various compression loads. For the FEM calculation, the following reference frame was adopted: the Y-axis was along the direction of specimen’s length, i.e. the loading direction; the X-axis was oriented to the direction of specimen’s width and the Z-axis was oriented to the direction of specimen’s thickness. 3. Experimental results 3.1. Results of compression tests at static strain rate Tables 2 and 3 show the mechanical properties measured in compression and tensile at strain rate of 1 × 10−3 s−1 for FL and DP specimens, respectively. Fig. 4(a) and (b) shows the engineering stress–strain curves measured. For comparison the data and curves measured in tensile tests [10] are also listed in the tables and plotted in the figures. From Tables 2 and 3 and Fig. 4, it is found that for
both FL and DP alloys the ultimate compression stress uc , and the fracture strain εf measured in compression tests are much higher than those measured in tensile tests, while the yield strength 0.2 is compatible. It is worthy to indicate that both the tensile and compression strength of DP alloy are lower than those of FL alloy, while the tensile fracture strain of DP is appreciably higher than that of FL. 3.2. Results of compression tests at different strain rates Tables 4 and 5 show the mechanical properties measured in compression tests at various strain rates for FL and DP specimens, respectively. Corresponding stress–strain curves measured at various compression strain rates for FL and DP are shown in Fig. 5. Fig. 6 shows the relationships between the strain rate and the 0.2 , ultimate compression stress uc and fracture strain εf of FL alloys. Fig. 7 shows those relationships of DP alloys. From Figs. 6 and 7, it can be seen that the trends of the variation of mechanical properties with increasing strain rates are different for two types of microstructures. For FL alloy, compression yield strength ( 0.2 ) appreciably increases with increasing the strain rate, however only minor increase is found for the uc . For DP TiAl alloy, while uc appreciably increases with increasing the strain rate, only monor change is measured in the 0.2 . 3.3. Results of observation on the macro-and micro-features of the fracture surfaces of specimens fractured in compression tests Fig. 8(a and b) shows the drawings of the configurations of cracks produced in FL specimens fractured at strain rates of 4 × 10−2 s−1 and 1 × 10−4 s−1 , respectively. Fig. 8(a) shows that the crack is composed of two inclined cracks initiated from both top and bottom surfaces and a short longitudinal crack in the middle. On the fracture surface all three parts of the crack show the bright smooth shear cracking pattern. The bright smooth shear cracking regions almost cover all of the fracture surfaces of specimens. Fig. 8(b) shows two inclined cracks, which start from both top and bottom surfaces and stop at the center part of the specimen. A short longitudinal ligament is left in the middle part of the specimen. Fig. 8(b) apparently shows the process of cracking that the crack shown in Fig. 8(a) starts
Fig. 4. Engineering stress–strain curves measured in compression and tensile tests for FL(a) and DP(b).
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Table 4 Mechanical properties measured in compression tests at various strain rates for FL specimens.
ε´ (s−1 ) −2
4 × 10 2 × 10−2 1 × 10−2 1 × 10−3 1 × 10−4
No.
0.2 (MPa)
uc (MPa)
εmax (%)
Wf (J/mm2 )
FL-16, FL-53 FL-28, FL-11, FL-52, FL-13 FL-03, FL-24 FL-21, FL-30 FL-31, FL-33
559.7 540.68 494.74 483.36 440.19
2275.29 2275.08 2262.84 2260.79 2166.69
29.56 28.23 29.08 28.3 29.0
5.434 5.552 5.418 5.507 5.505
Note: 0.2 : yield stress, uc : ultimate compression stress, εmax : strain at ultimate compression stress, Wf : fracture energy of unit area. Table 5 Mechanical properties measured in compression tests at various strain rates for DP specimens.
ε´ (s−1 )
No.
0.2 (MPa)
uc (MPa)
εmax (%)
Wf (J/mm2 )
2 × 10−2 1 × 10−2 1 × 10−3 1 × 10−4 1 × 10−5
DP-09 DP-32 DP-40 DP-06 DP-11 DP-33 DP-15 DP-17 DP-38 DP-25 DP-50
399.64 383.32 379.53 419.03 428.76
2466.57 2283.68 2226.75 2223.19 2258.70
33.45 30.58 30.91 31.4 33.74
6.446 5.370 5.269 5.784 5.700
Fig. 5. Engineering stress–strain curves measured at various compression strain rates for FL(a) and DP(b) TiAl alloys.
Fig. 6. Relationships between the strain rate and mechanical properties of FL alloys.
Fig. 7. Relationships between the strain rate and mechanical properties of DP alloys.
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Fig. 8. Drawings of crack configuration of FL specimens fractured at strain rate of 4 × 10−2 (a) and 1 × 10−4 (b).
Fig. 9. Fracture surface of fully lamellar ␥-TiAl alloys fractured at strain rate of 2 × 10−2 s−1 .
from both top and bottom surface as inclined shear cracks. With increasing the applied load the two inclined shear cracks propagate to the center line where they are connected by shearing the longitudinal ligament. It is further revealed that the shear cracking inclined with an angle near 45◦ at both ends and parallel to the compression axis in the middle dominates the fracture processes of FL specimens in all range of applied strain rates. Fig. 9 shows the micro-features of the fracture surfaces of specimen FL-53, which fractured at strain rates of 4 × 10−2 s−1 . To the macroscopic bright smooth shear cracking regions on the fracture surfaces, the corresponding micro-
features (Fig. 9(a)) shows that they are composed of a lot of smooth shear cracking facets which locate on various layers with steps connecting them. Fig. 9(b) shows the magnified figure of the shear cracking facets in Fig. 9(a), which are composed of a number of interlamellar shear cracking facets. Fig. 9(c) shows the magnified figure of the steps in Fig. 9(a), which are the translamellar shear cracking facets. Fig. 10 shows the micro-feature of the fracture surface of specimen FL-31, which fractured at strain rates of 1 × 10−4 s−1 . Similar to the specimen fractured at high strain rates, large interlamellar shear cracking facets (Fig. 10(b)) with translamellar shear cracking
Fig. 10. Fracture surface of fully lamellar ␥-TiAl alloys fractured at strain rate of 1 × 10−4 s−1 .
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Fig. 11. Drawings of crack configuration of DP specimens fractured at strain rate of 2 × 10−2 s−1 (DP-09) and (b) 1 × 10−5 s−1 (DP-50).
Fig. 12. Fracture surface of duplex ␥-TiAl alloys fractured at strain rate of 2 × 10−2 s−1 .
steps (Fig. 10(c)) dominate the fracture surface (Fig. 10(a)). However, less steps which are composed of the translamellar shear cracking facets are observed in specimens fractured at the lower stain rate. Fig. 11(a and b) shows the drawings of the configurations of cracks produced in DP specimens which fractured at strain rates of 2 × 10−2 s−1 (DP-09), and 1 × 10−5 s−1 (DP-50), respectively. Contrary to the FL specimens, in DP specimens the longitudinal (parallel to the compression axis) normal cracking, instead of the inclined shear cracking, dominates the fracture. In general an inclined shear crack is initiated from the top surface of the specimen and propagates a short distance then transfers to a longitudinal normal crack (or two). The longitudinal normal crack propagates through the major part of the fracture surface and then transfers again to an inclined shear crack which extends to the bottom surface. In specimen fractured at strain rate of 1 × 10−5 s−1 , the longitudinal normal crack also dominates the fracture surface with only two short inclined shear cracks at both ends. However, with increasing the strain rate to 2 × 10−2 s−1 , the length of the bottom shear crack
increases to a value accounting for 40% of total length of the fracture surface. Figs. 12 and 13 show the fracture surface of DP specimens which fractured at the strain rate of 2 × 10−2 s−1 and 1 × 10−5 s−1 , respectively. Fig. 12(a) and 13(a) are the low magnification figures which show bright smooth shear cracking regions (A) and gray rough longitudinal normal cracking regions (B). Corresponding to the bright smooth shear cracking regions in Fig. 12(a) and 13(a), the high magnification figures Fig. 12(b) and 13(b) show the interlamellar shear cracking facets with translamellar shear cracking steps, which look like what observed in FL specimens. Corresponding to the gray rough longitudinal normal cracking regions the high magnification figures Fig 12(c) and 13(c) show the cleavage fracture facets composed of interlamellar and translamellar cleavage cracking facets. With increasing the strain rate more translamellar cracking facets appear in this region. In summary for DP specimens the inclined cracks close to both upper and bottom surfaces show the bright smooth region in the macroscopic scale and shear cracking facets in microscopic scale
Fig. 13. Fracture surface of duplex ␥-TiAl alloys fractured at strain rate of 1 × 10−5 s−1 .
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(Fig. 12(b) and Fig. 13(b)). The longitudinal normal cracks, which dominate the fracture surface, show the pattern of cleavage cracking (Fig. 12(c) and Fig. 13(c)). 4. Discussion 4.1. Fracture processes Based on the observations of macro- and micro-feature of the fracture surfaces, the fracture mechanisms for FL and DP alloys are analyzed as follows: For both FL and DP specimen the shear cracking first starts from the top or/and the bottom surface. It means that in the compression test the shear stresses at the top or/and the bottom surfaces are the controlling parameters. In part I of this work, the shear cracks at the top and side surface of the specimen were observed at the applied load of 900 MPa. The corresponding shear stress is in the range of 478–620 MPa. It means that for FL alloy the shear strength is higher than the tensile strength (400–500 MPa). As shown in Fig. 4 at the loading rate of 1 × 10−3 s−1 the ultimate compression stress uc shown on the engineering stress–strain curve of FL alloy is around 2200 MPa. Fig. 14 shows the FEM-calculated distribution of the shear stress along a 45◦ inclined cross-section at the maximum applied compression stress of 2200 MPa. At this maximum applied compression stress, the peak shear stress of 1400 MPa is produced at points close to the top and bottom surfaces. It is considered that at this moment the shear cracks have extended to a length that is sufficient to make the applied load drops. As shown in Fig. 15 at this moment the peak normal stress induced by the ballooning of the specimen reaches a value of only 200 MPa at the center line about 2 mm from the top surface. By comparing it with values listed in Table 3 this peak normal stress does not reach the tensile strength of both FL and DP alloys (524 MPa for FL and 409 MPa for DP). This is the reason why the normal crack could not be directly initiated at this moment by the normal stress only. It is triggered by the shear cracks which were produced at the top and bottom surfaces and extend to the center of specimen. The formation of the longitudinal cracking are different for DP and FL alloys. For DP specimen when the shear crack reaches the center part (around 2 mm in depth) the normal stress induced by the ballooning of the specimen incorporates the normal stress that produced by the shear crack [9] (schematically shown in Fig. 16). The composed normal stress reaches the tensile strength of 400 MPa for DP alloy and causes the longitudinal normal cleavage cracking along the compression axis. As shown in Fig. 11 of the macro-fracture surface of DP specimen, the normal cracking dominates the fracture (longer than the half length of the fracture surface). However, for
Fig. 15. Distribution of the normal stress along the center line of the specimen at the applied load of 2200 MPa. (a) Normal stress distribution along longitudinal compression axis and (b) normal stress distribution along horizontal center line.
FL specimen the combined action of the normal stress produced by the ballooning of specimen and that produced by the shear cracking is insufficient to cause the longitudinal normal cracking. When the shear cracks initiated from both top and bottom surfaces extend further and decrease the area of the longitudinal ligament (Fig. 8(b)), the ligament between two shear cracks is broken by the shear stress acting parallel to the axis (schematically shown in Fig. 17). The normal stress induced by the ballooning of the specimens assists the broken process and makes the crack widening. Because the FEM cannot simulate the propagation of the shear cracking and the measured compression stress–strain data cannot characterize the behavior in tension, the exact normal stress at the shear crack tip cannot be calculated. By comparing the maximum tensile strengths listed in Table 3 for FL (524 MPa) and DP (409 MPa) alloys, the normal stress, which is jointly induced by the ballooning of the specimen and by the shear crack is estimated to be a value between 409 MPa and 524 MPa. This normal stress affords to cause the longitudinal normal crack in DP specimen but is insufficient to cause it in FL specimen. In the middle part of the FL specimen, the longitudinal crack connecting two inclined shear cracks is also a shear crack, which is certificated by the bright smooth pattern in the macro-feature of the fracture surface and by the shear fracture facets in the micro-feature of the fracture surface (Figs. 9 and 10). 4.2. Comparison between the FL and DP alloys’ fracture behaviors
Fig. 14. Distribution of the shear stress along a 45◦ inclined cross-section at applied load of 2200 MPa.
Based on the data listed in Table 3 and the observation of the fracture surfaces, the fracture behaviors of FL and DP alloys are compared as follows: In both tensile and compression tests, the yield strength y of FL is higher than that of DP, while the grain size of
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Fig. 16. Schematic of the normal stress produced in DP specimen.
Fig. 17. Schematic of the shear stress produced in FL specimen.
the former is larger than that of the latter. According to the wellknown Hall–Petch formula ( y = i + Kd−1/2 ), for the FL alloy with a larger grain size the higher yield strength must result from a high constant K. The constant K is determined by the process by which the mobile dislocations are produced in the neighboring un-yield
grain. The higher constant K of the larger grain FL is considered to be caused by the large inclined angles between lamellar layers in neighboring grains, which makes it more difficult to produce mobile dislocations in neighboring grains. Comparing with the fine grain of the two-phase DP alloy it is much more difficult for the dislocation pile to deliver the yield from one grain to the other in the FL alloy. Therefore, even though the FL alloy has a large grain it still shows the higher yield strength than that of DP alloy. It is also found that in Table 3, the tensile strength of FL is higher than that of the DP, however the maximum compression strength ( uc ) of both alloys is compatible. In [10,13] the higher tensile strength of FL has been attributed to the much rougher fracture surface, which was also to be caused by the very large lamellar layers deviating with a large angle from the tensile stress. Fig. 18 shows the path of the crack propagation in the FL alloy. As shown in Fig. 18(a) a crack propagates in direction with an angle inclined to the normal stress that acts in the horizontal direction. Fig. 18(b) shows that the crack propagates further as the interlamellar crack through several horizontally oriented boundaries between the large lamellae. Apparently the extension direction of these interlamellar cracks is parallel to the normal stress. Fig. 19 comparing the paths of crack propagation in FL and DP alloys. It is seen that the crack path deviates apparently from the maximum normal stress in FL alloy, while the crack path almost along the direction perpendicular to the maximum normal stress in the DP alloy. These figures (Figs. 18 and 19) show the effects of the microstructures. In the FL alloy, the large weakest lamellae interfaces make the crack deviating from the higher tensile stress. The deviation from the maximum stress decreases the normal stress acting on the crack surfaces and increases the length
Fig. 18. The configuration (path) of crack propagation in the FL alloy.
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Fig. 19. Comparison of the crack propagation path in FL(a) and DP(b) alloys.
of cracking. The higher tensile strength of FL is attributed to these effects. However, in the case of compression test, for both FL and DP alloys the mechanism controlling the ultimate compression stress is the same that is the shear strength. It is thought that the difference between the interlamellar and the translamellar shear strength is much less than the difference between the interlamellar and translamellar tensile strength, the effect of the inclined angle between the large lamellar layers is diminished. Therefore, the difference between the ultimate compression strengths of the large full lamellar microstructure (FL) and the fine duplex phase (DP) microstructure is not distinct. Therefore, the ultimate compression strength uc is compatible for FL and DP alloys. It is worthy to note again that the difference between the tensile strength of FL and DP results in the different fracture mechanism in the compression tests. Due to its lower tensile strength the longitudinal normal cracking dominates the fracture in compression test of DP. Instead for FL specimen, the higher tensile strength prohibits the transition from shear cracking to normal cracking in the middle of the specimen, and the shear cracking propagate through whole specimen. 4.3. Comparison between tensile and compression fracture behaviors From Table 3 the glaring difference presents between the tensile strength ut (around 409–524 MPa) and the ultimate compression uc (around 2200 MPa). The much lower values of the ut measured in tensile tests are attributed to two factors. The first factor is the effect of the micro-cracking damage produced during the tensile process. In Ref. [10] when the tensile preloading was carried out in the plastic regime the micro-cracking damage produced by the tensile preloading–unloading process appreciably decreased the final tensile fracture stress. These microcrack damages could also be induced in a directly loading-fracturing process before the applied stress reaches the ultimate tensile stress ut . The accumulated micro-cracking damage might seriously reduce the cross-section area of the specimen and decreased the tensile strength. However, as shown in the first part of this work [11], no effect of the compression preloading appears until the applied load reached the ultimate compression stress uc . It means that in compression test before the applied load reaches the ultimate compression stress, the produced damage has much less significant effects on the compression fracture behavior.
The second factor is the difference of the fracture mechanisms between the tensile test and the compression tests of the TiAl alloy. For the tensile test the fracture mechanism is the cleavage fracture caused by the tensile normal stress, while for the compression test it is controlled by the shear cracking. The weakest interlamellar interface has the lowest microscopic tensile strength of less than 100 MPa [12]. The global tensile strength was measured as lower than 524 MPa for FL and 409 MPa for DP. These are the strength against the normal cleavage cracking, which is controlled by the tensile stress. However, it has a higher strength of 620 MPa for initiation and 1400 MPa for fracturing against the shear cracking, which is the controlling factor in the compression test. Furthermore the maximum shear stress which acts in an inclined section, is lower than the applied normal compression stress. The same reasons can be used to explain why a much higher fracture strains are measured in compression tests (around 0.3) than that measured in tensile tests (0.007–0.014). The pre-requested strain for the cleavage fracture by tensile stress is much lower than that for the shear cracking. From Table 3 it is also found that the yield strength y measured in tensile tests and compression tests are similar. It means that no effective damage is produced in the elastic regime for both tensile and compression tests. This idea is identified by previous work and this work. In Ref. [10], in tensile test when the preloading was carried out in the elastic regime the loci of the preloading, unloading and reloading overlapped each other. In the first part of this work the same phenomena were observed in compression tests. It identified that in the elastic regime the preloading and unloading processes do not produce damage that will affect the subsequent loading process. 4.4. Effects of loading rate Appreciable effect of the loading rate was observed on the yield strength of FL alloy at the highest loading rates. Because for y of FL the controlling factor is the shear strength, the effects of loading rate can be explained by conventional concepts based on the movement of dislocations. Nevertheless an appreciable effect of the loading rate on the yield strength appeared in the FL alloy, only minor effect on the ultimate compression stress was observed. It is considered to be caused by difference of the locations where the slip and the fracture occurred in the FL alloy. From the observation of the specimen surface during the compression test the slip lines were produced in all volume of grains, while the micro-shear cracks
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prefer to initiated and propagated along large interlamellar interfaces. For the slipping in the grain due to the related movement of dislocations the effect of loading rate is significant, yet for the shear cracking along the lamellae boundaries this effect is minor. For the DP alloy, due to its fine lamellar grains with ␥-phase particles located in grain boundaries the shear cracks propagate through a way mixed by the interlamellar interfaces within the lamellae matrix and the ␥ grain matrix. In this case the effect of loading rate on the dislocation movement will appreciably affect the fracture behavior. Thus, an appreciable effect of the loading rate on the ultimate compression stress appears in DP alloy. The reason why no effect of the loading rate is observed on the yield strength of DP alloy is not yet clear. 5. Conclusion Based on the results of macroscopic compression tests, observations of the macro- and micro-fracture surfaces and the FEM calculations following ideas can be drawn for the present TiAl alloys: (1) In the compression test the fracture starts from the top or/and the bottom surfaces by the shear cracking with an angle around 45◦ inclined to the compression axis. For FL specimen the shear cracking propagates through whole specimen and dominates the fracture. While for DP specimen the inclined shear cracking transforms to longitudinal normal cracking which dominates the fracture. (2) The different fracture behavior results from the higher tensile strength of the FL alloy, which is composed of large lamellae with large angle inclined to both the normal stress and the lamellae in neighboring grains.
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(3) The superior strength and plasticity measured in compression tests relative to those measured in the tensile tests are attributed to the less deteriorate effects of the micro-cracking damage and the higher resistance to the shear cracking than the resistance to the tensile cracking of TiAl alloys. (4) The effect of the loading rate on the fracture behavior can be explained by conventional concepts based on the movement of dislocations. Acknowledgements This work was financially supported by the National Nature Science Foundation of China (No. 50471109), Nature Science Foundation of Gansu Province (No. 3ZS061-A25-037) and Opening Foundation of State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials (No. SKL04003). References [1] Y.-W. Kim, JOM 41 (7) (1989) 24–30. [2] D.M. Dimiduk, P.M. Hazzledine, T.A. Parthasarathy, Metall. Mater. Trans. A A29 (1) (1998) 37–47. [3] K.S. Chan, Y.-W. Kim, Metall. Trans. A 23A (6) (1992) 1663–1677. [4] K.S. Chan, Metall. Mater. Trans. A 31A (2) (2000) 71–80. [5] C.T. Liu, J.H. Schneibel, P.J. Maziasz, Intermetallics 4 (1996) 429–440. [6] K. Nonaka, K. Tanosaki, Mater. Trans. JIM 33 (9) (1992) 802–810. [7] W.F. Brace, B.W. Paulding, C. Scholz, J. Geophys. Res. 71 (16) (1966) 3939– 3953. [8] A. Bartels, H. Clemens, G. Dehm, Z. Metallkd 92 (3) (2002) 180–185. [9] A.N. Stroh, Proc. Roy. Soc. 64B (1951) 747. [10] R. Cao, H. Zhu, J.H. Chen, J. Zhang, J. Mater. Sci. 43 (2008) 299–311. [11] R. Cao, L. Li, J.H. Chen, J. Zhang. (submitted for publication). [12] J.H. Chen, R. Cao, G.Z. Wang, J. Zhang, Metall. Mater. Trans. A 35A (2004) 439–456. [13] R. Cao, J.H. Chen, J. Zhang, H. Zhu, Eng. Fract. Mech. 25 (2008) 4019–4035.
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Study on compression deformation, damage and fracture behavior of TiAl alloys Part II. Fracture behavior R. Cao a,b,∗ , L. Li a,b , J.H. Chen a,b , J. Zhang c a b c
State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials, Lanzhou University of Technology, Lanzhou 730050, China Key Laboratory of Non-ferrous Metal Alloys, The Ministry of Education, Lanzhou University of Technology, Lanzhou 730050, China Central Iron and Steel Research Institute, Beijing 100081, China
a r t i c l e
i n f o
Article history: Received 24 April 2008 Received in revised form 27 November 2008 Accepted 4 December 2008
Keywords: TiAl alloys Compression tests Fracture behavior Sensitivity to strain rate
a b s t r a c t Compression deformation, damage and fracture behaviors of TiAl alloys are studied systematically. The deformation and damage processes in compression tests of TiAl specimens are detailed in Part I and the fracture behaviors are described and discussed here the Part II of this work. Results indicate that: the compression properties are superior to the tensile properties and specially show a much higher fracture strength and plasticity. The macroscopic features of the fracture surfaces indicate that the fracture mechanisms are different between the full lamellar (FL) alloy and the duplex (DP) alloy. In FL alloy, the shear cracking dominates the fracture, and the shear crack extends throughout the whole specimen. At both ends of the specimen, the shear cracks initiate and propagate in directions inclined to the compression axis with an angle near 45◦ . In the middle of the specimen, the shear cracks propagate parallel to the compression axis. While in the DP alloy after two short shear cracks initiate and extend at both ends, the longitudinal normal cracking is triggered and propagates parallel to the compression axis and dominates the fracture. In the microscopic features the regions of shear cracking show the smooth interlamellar shear fracture facets with some translamellar shear fracture steps. The longitudinal normal cracking in DP specimens shows mixed interlamellar and translamellar cleavage fracture facets. Effect of loading rate manifests itself in increasing the shear yield strength of FL and the fracture strength of DP alloys. © 2009 Published by Elsevier B.V.
1. Introduction In the past two decades, ␥-TiAl-based alloys have received tremendous attention for high-temperature and high-performance structural applications, primarily due to their low density, high strength to weight ratio as well as their good creep and oxidation resistance [1,2]. However, as intermetallic compounds, some of their properties such as the toughness, formability at room temperature are much poorer than those of conventional structural materials therefore their applications are limited. Until now, a great number of studies on the fracture mechanism of TiAl alloys in tensile have been done and a series of systematic results have been achieved [3–5]. However, the reports on compression deformation, damage and fracture behavior are relatively rare. Ref. [6] found that fracture was caused by the growth and connection of grain boundary cracks. The ␣2 phase itself acted as an obstacle to prevent the crack propagation at the cross-points with
∗ Corresponding author at: State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials, Lanzhou University of Technology, Lanzhou 730050, China. Tel.: +86 931 2973529; fax: +86 931 2755806. E-mail addresses: [email protected], [email protected] (R. Cao). 0921-5093/$ – see front matter © 2009 Published by Elsevier B.V. doi:10.1016/j.msea.2008.12.012
the cracking by causing kinking and necking. Ref. [7] shows that there were three modes of crack initiation and propagation in the compression loading tests. As shown in Fig. 1(a) cracks initiate and propagate parallel to the compression axis. In Fig. 1(b) three cracks initiate and propagate at the intersection of three grain boundaries and one crack propagates along the compression axis, the other two cracks extend with an angle inclined to the compression axis. In Fig. 1(c) two cracks along compression axis direction initiate and propagate at both ends of an inclined precrack, which is produced at an inclined grain boundary. At the compression stress regions, microcracks nucleate at the precracks, cavities, inclusions and other defects. With applied load increasing, some cracks can propagate to the main crack or propagate and connect with each other and finally induce the specimen fracture. Ref. [8] shows that the fracture strength, the fracture strain and the cracking behavior were very sensitive to strain rate. At a strain rate of 5 × 10−3 s−1 (quasistatic condition), the fully lamellar samples showed large cracks in the direction of the compression cone, whereas in the dynamical conditions ((2–4) × 103 s−1 ), compression specimens exhibited a great number of short interlamellar microcracks. The high strain rate limited the length of the crack propagation. Therefore, in the quasi-static conditions, cracks grew to macroscopic crack, however at the dynamic deformation, few numbers of microcracks were
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Fig. 1. Three types of crack initiation and propagation in the compression loading mode [7].
Table 1 Compositions of TiAl alloy (at%). Ti Al V Cr
Balance 47.5 2.5 1.0
barely to grow to the macroscopic cracks. In the engineering applications, compression effects appear on blades of the combustion gas turbine by continuous start–shut down of the compressible gas. In case the mechanical connection is made by inserting the shaft of the TiAl wheel into the steel sleeve with the interference fit, the extra high compression stress maybe causes the cracking of the TiAl shaft. Therefore, the compression deformation and fracture behavior of TiAl alloys deserve to be studied for correct design and production of many engineering components with perfect performance. This is just the aim, which this work focuses on. 2. Experimental
Fig. 3. Dimensions of compression test specimen.
The average colony (grain) size and the average lath width of FL microstructure (Fig. 2a) were measured as 190–300 m and 2–3 m, respectively. The average grain size of the DP microstructure (Fig. 2b) was 20–30 m. The cubic compression specimens (12 mm × 5 mm × 5 mm) are used. The dimensions are shown in Fig. 3. 2.2. Compression tests
2.1. Material and specimens A TiAl alloy with compositions shown in Table 1 was used. All samples were taken from a forged pancake that has been deformed at 1100 ◦ C for a 70% height reduction. Two types of microstructures, duplex (DP) and near fully lamellar (FL) as shown in Fig. 2, were obtained. The samples were first wrapped in quartz tubes, treated by hot iso-static pressing at 120 MPa and 950 ◦ C in argon for 3 h and then put into the furnace at pre-determined temperatures. DP microstructure samples were obtained by annealing at 1250 ◦ C for 18 h and the FL microstructure samples by annealing at 1370 ◦ C for 1 h.
At room temperature compression tests were conducted in air by SHIMADSU AG-10TA universal test machine at strain rate of 1 × 10−3 s−1 . Engineering stress–strain curves were plotted in accord with the test data for all fractured specimens. 2.3. Compression tests at various strain rates Compression tests were done at various strain rates in range of 1 × 10−5 s−1 to 4 × 10−2 s−1 . The yield stress ( 0.2 ), ultimate compression stress ( uc ), fracture work of unit area (wf ) were measured with the engineering stress–strain curves of compression tests.
Fig. 2. Microstructures of (a) fully lamellar and (b) duplex TiAl-based alloy.
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Table 2 Mechanical properties measured in compression and tensile tests at strain rate of 1 × 10−3 s−1 for FL specimens.
Table 3 Mechanical properties measured in compression and tensile tests at strain rate of 1 × 10−3 s−1 for DP specimens.
Testing type
0.2 (MPa)
ut (MPa)
εmax (%)
f (MPa)
εf (%)
Testing type
0.2 (MPa)
ut (MPa)
εmax (%)
f (MPa)
εf (%)
Tensile Compression
484 484
524 2263
0.686 28.3
524 2144
0.686 31.4
Tensile Compression
325 379
409 2227
1.351 30.9
409 1868
1.351 31.2
Note: 0.2 : yield stress, ut : ultimate tensile stress, εmax : strain at ultimate tensile stress, f : fracture stress, εf : fracture strain.
Parameters of macroscopic mechanical properties measured at various strain rates were compared. 2.4. Observation on the macro- and micro-features of fracture surfaces The macro-features of the fracture surfaces of specimens fractured at different strain rates were observed by magnifying glass and compared in details for both FL and DP alloys. The crack configurations on the side surfaces of the specimens were also observed to identify the propagation process and the macro-features of the cracks. The micro-features of the fracture surfaces were observed by SEM and were related to the macro-features of the fracture surfaces. 2.5. FEM calculation A three-dimensional FEM model having eight-node threedimensional stress reduced integration elements C3D8R was used with the ABAQUS code. A total 35,175 elements and 39,938 nodes were taken in to account. The distributions of normal stress 11 and shear stress along a 45◦ -inclined cross-section in the specimens were calculated at various compression loads. For the FEM calculation, the following reference frame was adopted: the Y-axis was along the direction of specimen’s length, i.e. the loading direction; the X-axis was oriented to the direction of specimen’s width and the Z-axis was oriented to the direction of specimen’s thickness. 3. Experimental results 3.1. Results of compression tests at static strain rate Tables 2 and 3 show the mechanical properties measured in compression and tensile at strain rate of 1 × 10−3 s−1 for FL and DP specimens, respectively. Fig. 4(a) and (b) shows the engineering stress–strain curves measured. For comparison the data and curves measured in tensile tests [10] are also listed in the tables and plotted in the figures. From Tables 2 and 3 and Fig. 4, it is found that for
both FL and DP alloys the ultimate compression stress uc , and the fracture strain εf measured in compression tests are much higher than those measured in tensile tests, while the yield strength 0.2 is compatible. It is worthy to indicate that both the tensile and compression strength of DP alloy are lower than those of FL alloy, while the tensile fracture strain of DP is appreciably higher than that of FL. 3.2. Results of compression tests at different strain rates Tables 4 and 5 show the mechanical properties measured in compression tests at various strain rates for FL and DP specimens, respectively. Corresponding stress–strain curves measured at various compression strain rates for FL and DP are shown in Fig. 5. Fig. 6 shows the relationships between the strain rate and the 0.2 , ultimate compression stress uc and fracture strain εf of FL alloys. Fig. 7 shows those relationships of DP alloys. From Figs. 6 and 7, it can be seen that the trends of the variation of mechanical properties with increasing strain rates are different for two types of microstructures. For FL alloy, compression yield strength ( 0.2 ) appreciably increases with increasing the strain rate, however only minor increase is found for the uc . For DP TiAl alloy, while uc appreciably increases with increasing the strain rate, only monor change is measured in the 0.2 . 3.3. Results of observation on the macro-and micro-features of the fracture surfaces of specimens fractured in compression tests Fig. 8(a and b) shows the drawings of the configurations of cracks produced in FL specimens fractured at strain rates of 4 × 10−2 s−1 and 1 × 10−4 s−1 , respectively. Fig. 8(a) shows that the crack is composed of two inclined cracks initiated from both top and bottom surfaces and a short longitudinal crack in the middle. On the fracture surface all three parts of the crack show the bright smooth shear cracking pattern. The bright smooth shear cracking regions almost cover all of the fracture surfaces of specimens. Fig. 8(b) shows two inclined cracks, which start from both top and bottom surfaces and stop at the center part of the specimen. A short longitudinal ligament is left in the middle part of the specimen. Fig. 8(b) apparently shows the process of cracking that the crack shown in Fig. 8(a) starts
Fig. 4. Engineering stress–strain curves measured in compression and tensile tests for FL(a) and DP(b).
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Table 4 Mechanical properties measured in compression tests at various strain rates for FL specimens.
ε´ (s−1 ) −2
4 × 10 2 × 10−2 1 × 10−2 1 × 10−3 1 × 10−4
No.
0.2 (MPa)
uc (MPa)
εmax (%)
Wf (J/mm2 )
FL-16, FL-53 FL-28, FL-11, FL-52, FL-13 FL-03, FL-24 FL-21, FL-30 FL-31, FL-33
559.7 540.68 494.74 483.36 440.19
2275.29 2275.08 2262.84 2260.79 2166.69
29.56 28.23 29.08 28.3 29.0
5.434 5.552 5.418 5.507 5.505
Note: 0.2 : yield stress, uc : ultimate compression stress, εmax : strain at ultimate compression stress, Wf : fracture energy of unit area. Table 5 Mechanical properties measured in compression tests at various strain rates for DP specimens.
ε´ (s−1 )
No.
0.2 (MPa)
uc (MPa)
εmax (%)
Wf (J/mm2 )
2 × 10−2 1 × 10−2 1 × 10−3 1 × 10−4 1 × 10−5
DP-09 DP-32 DP-40 DP-06 DP-11 DP-33 DP-15 DP-17 DP-38 DP-25 DP-50
399.64 383.32 379.53 419.03 428.76
2466.57 2283.68 2226.75 2223.19 2258.70
33.45 30.58 30.91 31.4 33.74
6.446 5.370 5.269 5.784 5.700
Fig. 5. Engineering stress–strain curves measured at various compression strain rates for FL(a) and DP(b) TiAl alloys.
Fig. 6. Relationships between the strain rate and mechanical properties of FL alloys.
Fig. 7. Relationships between the strain rate and mechanical properties of DP alloys.
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Fig. 8. Drawings of crack configuration of FL specimens fractured at strain rate of 4 × 10−2 (a) and 1 × 10−4 (b).
Fig. 9. Fracture surface of fully lamellar ␥-TiAl alloys fractured at strain rate of 2 × 10−2 s−1 .
from both top and bottom surface as inclined shear cracks. With increasing the applied load the two inclined shear cracks propagate to the center line where they are connected by shearing the longitudinal ligament. It is further revealed that the shear cracking inclined with an angle near 45◦ at both ends and parallel to the compression axis in the middle dominates the fracture processes of FL specimens in all range of applied strain rates. Fig. 9 shows the micro-features of the fracture surfaces of specimen FL-53, which fractured at strain rates of 4 × 10−2 s−1 . To the macroscopic bright smooth shear cracking regions on the fracture surfaces, the corresponding micro-
features (Fig. 9(a)) shows that they are composed of a lot of smooth shear cracking facets which locate on various layers with steps connecting them. Fig. 9(b) shows the magnified figure of the shear cracking facets in Fig. 9(a), which are composed of a number of interlamellar shear cracking facets. Fig. 9(c) shows the magnified figure of the steps in Fig. 9(a), which are the translamellar shear cracking facets. Fig. 10 shows the micro-feature of the fracture surface of specimen FL-31, which fractured at strain rates of 1 × 10−4 s−1 . Similar to the specimen fractured at high strain rates, large interlamellar shear cracking facets (Fig. 10(b)) with translamellar shear cracking
Fig. 10. Fracture surface of fully lamellar ␥-TiAl alloys fractured at strain rate of 1 × 10−4 s−1 .
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Fig. 11. Drawings of crack configuration of DP specimens fractured at strain rate of 2 × 10−2 s−1 (DP-09) and (b) 1 × 10−5 s−1 (DP-50).
Fig. 12. Fracture surface of duplex ␥-TiAl alloys fractured at strain rate of 2 × 10−2 s−1 .
steps (Fig. 10(c)) dominate the fracture surface (Fig. 10(a)). However, less steps which are composed of the translamellar shear cracking facets are observed in specimens fractured at the lower stain rate. Fig. 11(a and b) shows the drawings of the configurations of cracks produced in DP specimens which fractured at strain rates of 2 × 10−2 s−1 (DP-09), and 1 × 10−5 s−1 (DP-50), respectively. Contrary to the FL specimens, in DP specimens the longitudinal (parallel to the compression axis) normal cracking, instead of the inclined shear cracking, dominates the fracture. In general an inclined shear crack is initiated from the top surface of the specimen and propagates a short distance then transfers to a longitudinal normal crack (or two). The longitudinal normal crack propagates through the major part of the fracture surface and then transfers again to an inclined shear crack which extends to the bottom surface. In specimen fractured at strain rate of 1 × 10−5 s−1 , the longitudinal normal crack also dominates the fracture surface with only two short inclined shear cracks at both ends. However, with increasing the strain rate to 2 × 10−2 s−1 , the length of the bottom shear crack
increases to a value accounting for 40% of total length of the fracture surface. Figs. 12 and 13 show the fracture surface of DP specimens which fractured at the strain rate of 2 × 10−2 s−1 and 1 × 10−5 s−1 , respectively. Fig. 12(a) and 13(a) are the low magnification figures which show bright smooth shear cracking regions (A) and gray rough longitudinal normal cracking regions (B). Corresponding to the bright smooth shear cracking regions in Fig. 12(a) and 13(a), the high magnification figures Fig. 12(b) and 13(b) show the interlamellar shear cracking facets with translamellar shear cracking steps, which look like what observed in FL specimens. Corresponding to the gray rough longitudinal normal cracking regions the high magnification figures Fig 12(c) and 13(c) show the cleavage fracture facets composed of interlamellar and translamellar cleavage cracking facets. With increasing the strain rate more translamellar cracking facets appear in this region. In summary for DP specimens the inclined cracks close to both upper and bottom surfaces show the bright smooth region in the macroscopic scale and shear cracking facets in microscopic scale
Fig. 13. Fracture surface of duplex ␥-TiAl alloys fractured at strain rate of 1 × 10−5 s−1 .
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(Fig. 12(b) and Fig. 13(b)). The longitudinal normal cracks, which dominate the fracture surface, show the pattern of cleavage cracking (Fig. 12(c) and Fig. 13(c)). 4. Discussion 4.1. Fracture processes Based on the observations of macro- and micro-feature of the fracture surfaces, the fracture mechanisms for FL and DP alloys are analyzed as follows: For both FL and DP specimen the shear cracking first starts from the top or/and the bottom surface. It means that in the compression test the shear stresses at the top or/and the bottom surfaces are the controlling parameters. In part I of this work, the shear cracks at the top and side surface of the specimen were observed at the applied load of 900 MPa. The corresponding shear stress is in the range of 478–620 MPa. It means that for FL alloy the shear strength is higher than the tensile strength (400–500 MPa). As shown in Fig. 4 at the loading rate of 1 × 10−3 s−1 the ultimate compression stress uc shown on the engineering stress–strain curve of FL alloy is around 2200 MPa. Fig. 14 shows the FEM-calculated distribution of the shear stress along a 45◦ inclined cross-section at the maximum applied compression stress of 2200 MPa. At this maximum applied compression stress, the peak shear stress of 1400 MPa is produced at points close to the top and bottom surfaces. It is considered that at this moment the shear cracks have extended to a length that is sufficient to make the applied load drops. As shown in Fig. 15 at this moment the peak normal stress induced by the ballooning of the specimen reaches a value of only 200 MPa at the center line about 2 mm from the top surface. By comparing it with values listed in Table 3 this peak normal stress does not reach the tensile strength of both FL and DP alloys (524 MPa for FL and 409 MPa for DP). This is the reason why the normal crack could not be directly initiated at this moment by the normal stress only. It is triggered by the shear cracks which were produced at the top and bottom surfaces and extend to the center of specimen. The formation of the longitudinal cracking are different for DP and FL alloys. For DP specimen when the shear crack reaches the center part (around 2 mm in depth) the normal stress induced by the ballooning of the specimen incorporates the normal stress that produced by the shear crack [9] (schematically shown in Fig. 16). The composed normal stress reaches the tensile strength of 400 MPa for DP alloy and causes the longitudinal normal cleavage cracking along the compression axis. As shown in Fig. 11 of the macro-fracture surface of DP specimen, the normal cracking dominates the fracture (longer than the half length of the fracture surface). However, for
Fig. 15. Distribution of the normal stress along the center line of the specimen at the applied load of 2200 MPa. (a) Normal stress distribution along longitudinal compression axis and (b) normal stress distribution along horizontal center line.
FL specimen the combined action of the normal stress produced by the ballooning of specimen and that produced by the shear cracking is insufficient to cause the longitudinal normal cracking. When the shear cracks initiated from both top and bottom surfaces extend further and decrease the area of the longitudinal ligament (Fig. 8(b)), the ligament between two shear cracks is broken by the shear stress acting parallel to the axis (schematically shown in Fig. 17). The normal stress induced by the ballooning of the specimens assists the broken process and makes the crack widening. Because the FEM cannot simulate the propagation of the shear cracking and the measured compression stress–strain data cannot characterize the behavior in tension, the exact normal stress at the shear crack tip cannot be calculated. By comparing the maximum tensile strengths listed in Table 3 for FL (524 MPa) and DP (409 MPa) alloys, the normal stress, which is jointly induced by the ballooning of the specimen and by the shear crack is estimated to be a value between 409 MPa and 524 MPa. This normal stress affords to cause the longitudinal normal crack in DP specimen but is insufficient to cause it in FL specimen. In the middle part of the FL specimen, the longitudinal crack connecting two inclined shear cracks is also a shear crack, which is certificated by the bright smooth pattern in the macro-feature of the fracture surface and by the shear fracture facets in the micro-feature of the fracture surface (Figs. 9 and 10). 4.2. Comparison between the FL and DP alloys’ fracture behaviors
Fig. 14. Distribution of the shear stress along a 45◦ inclined cross-section at applied load of 2200 MPa.
Based on the data listed in Table 3 and the observation of the fracture surfaces, the fracture behaviors of FL and DP alloys are compared as follows: In both tensile and compression tests, the yield strength y of FL is higher than that of DP, while the grain size of
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Fig. 16. Schematic of the normal stress produced in DP specimen.
Fig. 17. Schematic of the shear stress produced in FL specimen.
the former is larger than that of the latter. According to the wellknown Hall–Petch formula ( y = i + Kd−1/2 ), for the FL alloy with a larger grain size the higher yield strength must result from a high constant K. The constant K is determined by the process by which the mobile dislocations are produced in the neighboring un-yield
grain. The higher constant K of the larger grain FL is considered to be caused by the large inclined angles between lamellar layers in neighboring grains, which makes it more difficult to produce mobile dislocations in neighboring grains. Comparing with the fine grain of the two-phase DP alloy it is much more difficult for the dislocation pile to deliver the yield from one grain to the other in the FL alloy. Therefore, even though the FL alloy has a large grain it still shows the higher yield strength than that of DP alloy. It is also found that in Table 3, the tensile strength of FL is higher than that of the DP, however the maximum compression strength ( uc ) of both alloys is compatible. In [10,13] the higher tensile strength of FL has been attributed to the much rougher fracture surface, which was also to be caused by the very large lamellar layers deviating with a large angle from the tensile stress. Fig. 18 shows the path of the crack propagation in the FL alloy. As shown in Fig. 18(a) a crack propagates in direction with an angle inclined to the normal stress that acts in the horizontal direction. Fig. 18(b) shows that the crack propagates further as the interlamellar crack through several horizontally oriented boundaries between the large lamellae. Apparently the extension direction of these interlamellar cracks is parallel to the normal stress. Fig. 19 comparing the paths of crack propagation in FL and DP alloys. It is seen that the crack path deviates apparently from the maximum normal stress in FL alloy, while the crack path almost along the direction perpendicular to the maximum normal stress in the DP alloy. These figures (Figs. 18 and 19) show the effects of the microstructures. In the FL alloy, the large weakest lamellae interfaces make the crack deviating from the higher tensile stress. The deviation from the maximum stress decreases the normal stress acting on the crack surfaces and increases the length
Fig. 18. The configuration (path) of crack propagation in the FL alloy.
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Fig. 19. Comparison of the crack propagation path in FL(a) and DP(b) alloys.
of cracking. The higher tensile strength of FL is attributed to these effects. However, in the case of compression test, for both FL and DP alloys the mechanism controlling the ultimate compression stress is the same that is the shear strength. It is thought that the difference between the interlamellar and the translamellar shear strength is much less than the difference between the interlamellar and translamellar tensile strength, the effect of the inclined angle between the large lamellar layers is diminished. Therefore, the difference between the ultimate compression strengths of the large full lamellar microstructure (FL) and the fine duplex phase (DP) microstructure is not distinct. Therefore, the ultimate compression strength uc is compatible for FL and DP alloys. It is worthy to note again that the difference between the tensile strength of FL and DP results in the different fracture mechanism in the compression tests. Due to its lower tensile strength the longitudinal normal cracking dominates the fracture in compression test of DP. Instead for FL specimen, the higher tensile strength prohibits the transition from shear cracking to normal cracking in the middle of the specimen, and the shear cracking propagate through whole specimen. 4.3. Comparison between tensile and compression fracture behaviors From Table 3 the glaring difference presents between the tensile strength ut (around 409–524 MPa) and the ultimate compression uc (around 2200 MPa). The much lower values of the ut measured in tensile tests are attributed to two factors. The first factor is the effect of the micro-cracking damage produced during the tensile process. In Ref. [10] when the tensile preloading was carried out in the plastic regime the micro-cracking damage produced by the tensile preloading–unloading process appreciably decreased the final tensile fracture stress. These microcrack damages could also be induced in a directly loading-fracturing process before the applied stress reaches the ultimate tensile stress ut . The accumulated micro-cracking damage might seriously reduce the cross-section area of the specimen and decreased the tensile strength. However, as shown in the first part of this work [11], no effect of the compression preloading appears until the applied load reached the ultimate compression stress uc . It means that in compression test before the applied load reaches the ultimate compression stress, the produced damage has much less significant effects on the compression fracture behavior.
The second factor is the difference of the fracture mechanisms between the tensile test and the compression tests of the TiAl alloy. For the tensile test the fracture mechanism is the cleavage fracture caused by the tensile normal stress, while for the compression test it is controlled by the shear cracking. The weakest interlamellar interface has the lowest microscopic tensile strength of less than 100 MPa [12]. The global tensile strength was measured as lower than 524 MPa for FL and 409 MPa for DP. These are the strength against the normal cleavage cracking, which is controlled by the tensile stress. However, it has a higher strength of 620 MPa for initiation and 1400 MPa for fracturing against the shear cracking, which is the controlling factor in the compression test. Furthermore the maximum shear stress which acts in an inclined section, is lower than the applied normal compression stress. The same reasons can be used to explain why a much higher fracture strains are measured in compression tests (around 0.3) than that measured in tensile tests (0.007–0.014). The pre-requested strain for the cleavage fracture by tensile stress is much lower than that for the shear cracking. From Table 3 it is also found that the yield strength y measured in tensile tests and compression tests are similar. It means that no effective damage is produced in the elastic regime for both tensile and compression tests. This idea is identified by previous work and this work. In Ref. [10], in tensile test when the preloading was carried out in the elastic regime the loci of the preloading, unloading and reloading overlapped each other. In the first part of this work the same phenomena were observed in compression tests. It identified that in the elastic regime the preloading and unloading processes do not produce damage that will affect the subsequent loading process. 4.4. Effects of loading rate Appreciable effect of the loading rate was observed on the yield strength of FL alloy at the highest loading rates. Because for y of FL the controlling factor is the shear strength, the effects of loading rate can be explained by conventional concepts based on the movement of dislocations. Nevertheless an appreciable effect of the loading rate on the yield strength appeared in the FL alloy, only minor effect on the ultimate compression stress was observed. It is considered to be caused by difference of the locations where the slip and the fracture occurred in the FL alloy. From the observation of the specimen surface during the compression test the slip lines were produced in all volume of grains, while the micro-shear cracks
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prefer to initiated and propagated along large interlamellar interfaces. For the slipping in the grain due to the related movement of dislocations the effect of loading rate is significant, yet for the shear cracking along the lamellae boundaries this effect is minor. For the DP alloy, due to its fine lamellar grains with ␥-phase particles located in grain boundaries the shear cracks propagate through a way mixed by the interlamellar interfaces within the lamellae matrix and the ␥ grain matrix. In this case the effect of loading rate on the dislocation movement will appreciably affect the fracture behavior. Thus, an appreciable effect of the loading rate on the ultimate compression stress appears in DP alloy. The reason why no effect of the loading rate is observed on the yield strength of DP alloy is not yet clear. 5. Conclusion Based on the results of macroscopic compression tests, observations of the macro- and micro-fracture surfaces and the FEM calculations following ideas can be drawn for the present TiAl alloys: (1) In the compression test the fracture starts from the top or/and the bottom surfaces by the shear cracking with an angle around 45◦ inclined to the compression axis. For FL specimen the shear cracking propagates through whole specimen and dominates the fracture. While for DP specimen the inclined shear cracking transforms to longitudinal normal cracking which dominates the fracture. (2) The different fracture behavior results from the higher tensile strength of the FL alloy, which is composed of large lamellae with large angle inclined to both the normal stress and the lamellae in neighboring grains.
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(3) The superior strength and plasticity measured in compression tests relative to those measured in the tensile tests are attributed to the less deteriorate effects of the micro-cracking damage and the higher resistance to the shear cracking than the resistance to the tensile cracking of TiAl alloys. (4) The effect of the loading rate on the fracture behavior can be explained by conventional concepts based on the movement of dislocations. Acknowledgements This work was financially supported by the National Nature Science Foundation of China (No. 50471109), Nature Science Foundation of Gansu Province (No. 3ZS061-A25-037) and Opening Foundation of State Key Laboratory of Gansu Advanced Non-ferrous Metal Materials (No. SKL04003). References [1] Y.-W. Kim, JOM 41 (7) (1989) 24–30. [2] D.M. Dimiduk, P.M. Hazzledine, T.A. Parthasarathy, Metall. Mater. Trans. A A29 (1) (1998) 37–47. [3] K.S. Chan, Y.-W. Kim, Metall. Trans. A 23A (6) (1992) 1663–1677. [4] K.S. Chan, Metall. Mater. Trans. A 31A (2) (2000) 71–80. [5] C.T. Liu, J.H. Schneibel, P.J. Maziasz, Intermetallics 4 (1996) 429–440. [6] K. Nonaka, K. Tanosaki, Mater. Trans. JIM 33 (9) (1992) 802–810. [7] W.F. Brace, B.W. Paulding, C. Scholz, J. Geophys. Res. 71 (16) (1966) 3939– 3953. [8] A. Bartels, H. Clemens, G. Dehm, Z. Metallkd 92 (3) (2002) 180–185. [9] A.N. Stroh, Proc. Roy. Soc. 64B (1951) 747. [10] R. Cao, H. Zhu, J.H. Chen, J. Zhang, J. Mater. Sci. 43 (2008) 299–311. [11] R. Cao, L. Li, J.H. Chen, J. Zhang. (submitted for publication). [12] J.H. Chen, R. Cao, G.Z. Wang, J. Zhang, Metall. Mater. Trans. A 35A (2004) 439–456. [13] R. Cao, J.H. Chen, J. Zhang, H. Zhu, Eng. Fract. Mech. 25 (2008) 4019–4035.