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The effect of microstructures on the fatigue crack growth in Ti6Al4V laser welds
Int. J. Fatigue Vol. 19. No. 10, pp. 713-720, 1997 <;~ 1998 Elsevier Science Ltd. All rights reserved Printed in Great Britain {)142 I 123/97/$17.00+.00
ELSEVIER
PIh S0142-1123(97)00113-8
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds L. W. Tsay* and C. Y. Tsay Institute of Materials Engineering, National Taiwan Ocean University, Keelung, Taiwan, Republic of China (Received 20 May 1997; revised 18 August 1997; accepted 21 August 1997) The effect of microstructures on fatigue crack growth rates (FCGRs) in Ti-6AI-4V plates and laser welds was investigated. Experimental results revealed that the location of tensile fracture of laserwelded specimens was always in the base metal, owing to the presence of martensite in the narrow fusion zone (FZ). For welds aged at various temperatures, only minor differences in tensile properties were observed. The various heat treatments of such plate did not affect FCGRs significantly. The fracture surface was characterized by fatigue striation at high AK values. When the R ratio was increased from 0.1 to 0.4, higher FCGRs would be resulted. The decomposition of martensite in the aged FZ causes retardation of FCGRs in a weld, which is characterized by a mixed mode fracture. The more tortuous crack path within the FZ and HAZ was also indicative of high crack growth resistance. Roughness-induced crack closure caused by crystallographic fracture was responsible fi~r decreased FCGRs. © 1998 Elsevier Science Ltd. All rights reserved (Keywords: Ti-6AI--4V; laser welding; fatigue c r a c k growth rates; fatigue striation; cleavage fracture)
INTRODUCTION
larger packet size in the structure results in lower FCGRs s,'. Because of the high temperature reactivity of titanium alloys, electron beam (EB) welding is the most popular method for joining Ti-6AI~4V with high quality"( However, EB welding is often limited by the size of the vacuum chamber. Laser beam (LB) welding can provide welds of similar quality as compared to those of EB welds". Raising the oxygen content in the T i - 6 - 4 alloy is known to increase the tensile strength while decreasing tensile ductility ~2. Interstitial element (O, N) contamination can further decrease tensile ductility of the weld metal ~. it was found that no significant oxygen contamination occurred in laser welding with effective shielding in the experiment '4. Preparation of the surface to be welded, which removes the titanium oxide from the surface, does not affect the oxygen content in the titanium welds ~5. Relatively few works were performed to investigated the fatigue crack growth behavior of T i - 6 - 4 welds. The present work was undertaken to investigate the effect of microstructures on the mechanical properties in T i - 6 - 4 plates and laser welds. FCGRs were measured in various regions of the welds and particularly paid attention to the regions near the fusion zone (FZ). Both the microstructural and the stress ratio (R) effects were also evaluated. Fractographic observations were performed on various specimens after different heat treatments. Meanwhile, the variations of FCGRs in the welds could be correlated well with the changes in the fracture mode.
Ti-6A1-4V alloy is widely used in the aerospace industries for its remarkable strength:weight ratio and resistance to high temperature creep. This alloy can be heattreated to obtain various microstructures. The shape and distribution of the alpha (o0, beta (/3) and martensite (c~') are considered to be closely related to the mechanical properties of the material. Peter and Williams L indicated that microstructural changes can produce significant variations in crack growth rates at low AK values for Ti-6A1-2V-Mo. Above a critical value of AK, the crack growth is found to become structure insensitive 2. The criterion for the fracture transition from structure sensitive to insensitive is considered to be a change from primary single to multiple slip within the individual grain at the crack tip 2. The growth mode also changes when the reversed plastic zone size exceeds the scale of the microstructure, that is, the Widmanstfitten colonies or alpha laths ~:. Different microstructures exhibit different characteristic growth rates in structure sensitive regions '2-4 % Previous research indicated that significant reduction in region II growth rates for titanium alloys can be obtained by microstructure modification 7 ". Fatigue crack growth rates (FCGRs) can be substantially reduced for /3-anneal specimens, in which crack bifurcation occurs in the Widmanstfitten packet 6,7. The
*Author Ik)r correspondence.
713
L. W. Tsay and C. Y. Tsay
714
welds at various temperatures in vacuum tk~llowed by air cooling. The suggested PWHT is in the temperature range of 538-593°C (1000-1100°F) for 2 4 h ~ % In the present study, PWHTs were at 500, 550 or 600°C for 3 h. The mechanical tests in the study included microhardness measurements, tensile and fatigue crack growth tests. Microhardness measurements were taken across the weld to compare hardness variations in the FZ, heat-affected zone (HAZ) and unaffected base metal (BM). Tensile specimens, with a gage length of 25 mm and a nominal width of 6 ram, were made according to the specifications of ASTM E8M-90a (subsize, 2.5 mm thick) 17. The tensile strain rate was set at 7 x 10 4s ' and the results were the average of at least three specimens for each testing condition. Compact-tension (CT) specimens with the rolling direction (RD) normal to the crack path (see Figure 2) were employed. The experimental procedures met the standard of ASTM E647-91 ~* specifications. The applied load had a constant amplitude sinusoidal waveform and the loading frequency was 20 Hz. The stress ratio, that is, the ratio of minimum to maximum load, was set at 0.1 or 0.4. The crack growth directions were aligned either parallel or normal to the welding direction in the welded CT specimens (Figure 2). When the crack path was normal to the welding direction, the distance between the notch tip and the bottom fusion line of the weld was about 6.0 mm (Figure 2b). Specimens used in the above tests were prepared as follows: 12.5 + 0.1 Figure 1 SEM metallographs of T i ~ - 4 plates in the (a) as-received and (b) solution-treated conditions 30 0.1
M A T E R I A L AND E X P E R I M E N T A L PROCEDURES The chemical composition by weight percentage of the 3.1 m m thick T i - 6 A 1 4 V alloy was 6.20 A1, 4.22 V, 0.14 Fe, 0.01 C, 0.0136 O and 0.007 N and balanced Ti. The microstructure of the as-received T i - 6 - 4 consisted of small percentage of /3 phase distributed at the elongated c~ grain boundaries (Figure la). The alloy was solution-treated in vacuum at 950°C for 40 min and then cooled subsequently by Ar quench. Because the material was held below the /3 transus temperature, the microstructure consists equiaxed primary c~ and transformed /3 (Figure lb). All specimens were welded in the solution-treated condition. A Rofin-Sinar RS 850 5 KW CO2 laser was utilized for laser welding. Table 1 lists the laser welding variables used in this study. After welding, postweld heat treatments (PWHTs) were performed on the
Table 1
Laser welding parameter used in the experiment
Laser power Travel speed Focal lens Focal length Focal point Plasma assist gas flow rate Backing gas flow rate
2500 W 1500 m m rain ZnSe 190.5 mm 0.5 m m underfocus (0.5 mm below the surface) 25 lpm He 15 lpm Ar
,37s-+0, / RD
~
n
--
30_+0,1
1 62.54- 0.1
(b)
RD
1] Unit • mm Figure 2 Schematic representation showing CT specimens employed in the fatigue crack growth test crack growth (a) parallel and (b) normal to welding direction. RD is the rolling direction o1" the plate and FZ represents the fusion zone
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
715
the specimen's profile was cut using a electrodischarge machine with a wire electrode after which the specimens was ground to the required thickness. Metallographic observations and microhardness measurements were made to identity each region of the weld. Crack growth across the FZ can reveal the variation of crack growth rates in the different regions of the weld. The compliance function proposed by Saxena and Hudak ~° is used to determine the crack length, , / W = C,, + C,U, + C~_U~ + C~U~ + CaU ~, + C~U 5, where U, = I/[(BEWP)°5 + I] Here ~l is the crack length, W the specimen width, B the specimen thickness, E Young's modulus, P the load, V the crack opening displacement and Co, CI, C2, C~, Ca, C5 are compliance coefficients. The calculation of AK was made according to2°: A K = [A P / B W 1/2][f(a/W) ] f ( a / W ) = [(2 + a/W)/(I - a/W)3/2][0.886 + 4.64(a/W) -
13.32(a/W) 2 + 14.72(a/W) 3 - 5.6(a/W) a]
where AP is the loading range. Fractographic examinations were carried out by a scanning electron microscope (SEM). Thin foil specimens for transmission electron microscope (TEM) were prepared by a standard jet-polisher in a electrolyte consisting of 95% methyl alcohol and 5% sulfuric acid before being examined by a JEOL-2000EX microscope operated at 160 kV. RESULTS AND DISCUSSION MicrostructuraI observation The metallograph of the as-welded FZ Figure 3a), reveals the microstructure is a coarse /3 grain with internally acicular morphology. TEM micrographs show that the FZ consists of mainly oe' martensite (Figure 3b) as compared to granular a and refined lamellar c~ + /3 in the BM (Figure 3c). The acicular oe' in the FZ of laser welds was attributed to rapid cooling of the molten weld pool during solidification. In addition, the microstructure in the HAZ would also depend on the thermal history of welding. The microstructure of the HAZ in the vicinity of the fusion boundary was almost the same as the FZ, while regions of the HAZ far away from the fusion boundary were heated to relatively low temperatures and had a microstructure similar to that of the BM. The structure in the middle region of the HAZ was found to be a mixture of primary oe and oe'. Microstructures of the FZ aged at various temperatures were similar and revealed the presence of decomposed martensite. Figure 4 shows that the precipitation of /3 from the martensite was located primarily at the lath boundaries. Hardness measurement Figure 5 is a series of schematic representations showing the microhardness distribution in the welds. The hardness of the FZ and HAZ is higher than that of the BM. The highest hardness is within the FZ while the hardness in the HAZ drops as the distance from the fusion boundaries increases. These results are consistent with the microstructural observations, that
Figure 3 Photographs revealing the: (a) as-welded structure of the FZ; (b) TEM micrograph of (at; and (cl TEM microgruph of the BM in the solution-treated condition is, there is a refined microstructure with a higher hardness in the FZ. In addition, the hardness in the FZ could be raised slightly after PWHTs. The slight increase in hardness in the FZ after PWHTs was associated with the decomposition of ¢x' martensite to form c~ + /3, as shown in Figure 4. it is noted that only minor differences in hardness of the FZ were found among welds aged at different temperatures. The BM was also slightly hardened after aging, and this could be correlated with the precipitation of fine a from small quantities of metastable /32t, but the BM remained softer than the other region. Thus, fracture of welded specimens in this region can be anticipated.
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Tensile properties The room temperature tensile properties of T i - 6 - 4 base plates are listed in Table 2. As-received specimens with the RD normal to the tensile axis exhibited the lowest strength among the various specimens being tested, while the base plates aged in the temperature range of 500-600°C had only a slightly higher strength than the solution-treated specimens. The increase in hardness in the BM after aging, as shown in Figure 5, accounts for a slightly increase in strength in the aging temperature range of concern. Furthermore, the variations in tensile properties is minor for the variously aged specimens. Table 3 displays the tensile results of laser welds aged at different temperatures. It is important to mention that necking of tensile specimens was observed for all aged welds and that the fracture was located within the BM regardless of PWHTs. Tensile properties remained more or less the same for all the welds. During tensile testing, the FZ, HAZ and BM across the gage length underwent a nonuniform deformation which led to a slight loss in ductility. However, it implied that laser welds have consistent tensile properties as compared to those of the base plates. In the case of T i - 6 - 4 electroslag welds, ductility loss is correlated to the large prior /3 grain size and the microstructure produced by the high heat input j3. Laser welds show superior ductility and this has been attributed to fine prior /3 grain size 22. In the present study, the presence of martensite in the narrow FZ of laser welds might be the reason for these
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Figure 5 Schemalic diagrams showing microhardness distributions of the welds (a) as-welded, (b) 500 °C, (c) 550 °C and (d) 600 °C per 3 h aged conditions
excellent properties. Tensile-fractured surfaces examined by SEM revealed a ductile dimple fracture for all the specimens.
Fatigue crack propagation The FCGR (da/dN) plotted against the stress intensity factor range (AK) curve of T i - 6 - 4 base plates subjected to different heat treatments are shown in Figure 6. The fatigue crack growth tests reveal da/dN values fall within a narrow band for various specimens being tested (Figure 6a). In general, the 600°C aged specimens demonstrated the lowest crack growth rates, while the 500°C aged specimens had slightly higher crack growth rates than those of other specimens. Figure 6b shows the effect of increasing R from 0.1
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
717
Table 2 Tensile properties of Ti-6AI~4V plate Specimen condition
U.T.S. (MPa)
As-received~' As-received~' Solution 950°C per 40 min Aged 500°C per 3 h Aged 550°C per 3 h Aged 600°C per 3 h
954 936 957 980 977 970
Y.S. (MPa)
Elongation (%)
910 916 854 894 896 895
Reduction in area (~/~)
17.2 17.6 17.1 15.2 15.6 16.3
31.2 31.9 30.7
28.2
32.7 33,4
"Tensile axis parallel rolling direction bTensilc axis normal to rolling direction Table 3
Tensile properties of Ti-6AIqV laser welds
Specimen condition
U.T.S (MPa)
Y.S. (MPa)
Aged 500°C Aged 550°C Aged 600°(7
971 978 974
890 891 886
Elongation (~/~) 14.5 14.6 15.0
Reductionin area Ic,4) 30.6 3 I.I 31.9
Fracturelocation Base metal Base metal Base metal
All tensile specimen had been laser welded then aging treatmenl to 0.4 on FCGRs of as-received and 550°C aged specimens with similar ultimate tensile strength. Clearly FCGRs increase as R is increased. Fatigue crack growth behavior of Ti-6--4 laser welds after aging is shown in Figure 7. In Figure 7a the crack path is normal to the welding direction. The total width of the FZ and HAZ, as measured in Figure 5, is about 4 ram. Therefore, as the crack propagates across the FZ. the region of concern is only within a limited range of AK. When the crack grows from the BM into the HAZ, the FCGRs decrease immediately. After further growth of the crack, lower FCGRs are observed in the FZ. As the crack re-enters into the BM, the FCGRs return to levels similar to those for base plates subjected to the same aging treatment. Thus the results reveal that the FZ retards the crack growth most among various regions in a weld. Retardation of crack growth in the FZ is not so obvious at R = 0.4, however. As mentioned above, the microstructure of the HAZ in various locations would depends on its thermal history. The low temperature H A Z has a microstructure corresponding to the BM and thus has similar fatigue properties to that of the BM. However, the microstructure of the HAZ in the vicinity of the FZ consists predominantly of martensite, hence its fatigue properties will correspond to those of the FZ. As can be seen in Figure 6a, the initial lower FCGRs at low AK values are obtained in the BM, an area which does not undergo any microstructual changes. The HAZ and FZ always show residual tensile stress in the direction parallel to the welding direction, while residual compressive stresses are developed in the BM 23 25. Residual welding stresses can be reduced by P W H T at high temperature 25. The residual compressive stresses that exist in the BM can, however, effectively reduce the crack growth rate in this region 25. In addition, at smaller AK values or lower R ratios, the effect of residual stresses on the da/dN vs AK relationship is more pronounced 26. Horikawa et al. 2v also indicated that under these conditions residual welding stresses have a great influence on the crack initiation. As shown in Figure 7a the crack growth rates are low at low AK ranges, and this can be attributed to the existence of residual compressive stresses in the BM.
Thomas et al. 2~ indicated that a low temperature PWHT, such as stress relief or aging treatment, results in a deterioration in the fracture toughness of the FZ for T i - 6 4 EB welds. However, an investigation of T i - 6 A I - 2 V - M o TIG welds ~ revealed lower strength and ductility, resistance to fatigue crack growth in the HAZ and FZ was higher than in the BM. In the present study, T i - 6 4 laser welds were found to have not only good tensile properties but also high resistance to fatigue crack growth. As the crack grows in the FZ and parallel to welding direction for the 550°C aged welds (Figure 7b), crack growth resistance in the FZ is significantly better than in the same aged base plate. However, the difference in FCGRs between base plate and 550°C aged welds becomes less marked with increasing AK. The results also revealed an obvious increase in FCGRs was observed for the 550°C aged welds as R was increased from 0.1 to 0.4.
Fractographic observations SEM fractographs of T i - 6 - 4 base plates and laser welds were shown in Figure 8. Fatigue fractographs of T i - 6 - 4 base plates after different heat treatment were characterized by transgranular fractures (Figure 8a). Fatigue striations become clearer at high AK values. As the crack grows across the FZ, low magnification of the fracture surface reveals that the FZ has a remarkable rough surface for aged laser welds (Figure 8b). However, the crack path ill the same aged T i - 6 - 4 base plate is fairly smooth (Figure 8c). A mixed mode fracture with coarse cleavage facets was observed in the FZ (Figure 8d), as opposed to the striations that covered a large part of the fracture surface in the BM. In the investigation of fatigue crack growth transition in T i - 6 - 4 alloy, significant portions of cleavage facets was characterized by a high level of crack closure 3. The fracture surface of the coarse grain HAZ also shows a mixed mode fracture with less cleavage facets (Figure 8e), as compared to that of the FZ. In summary, the crack path is more irregular and tortuous in the FZ and HAZ than in the BM. The changes of growth direction are assisted by crystallographic fracture. Thus, the overall decrease in FCGRs can be attributed to the effect of roughness-induced
L. W. Tsay and C. Y, Tsay
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crack closure. As the crack propagated along the FZ, the overall fracture appearance at low magnification also suggested that there was a relatively unfavorable orientation for crack propagation (Figure 9a). Transgranular and cleavage fractures coexisted on the fracture surface (Figure 9b). However, The extent of crystallographic-induced fracture is decreased with increasing_stress intensity factor range. When AK is > 20 MPa gm, it is mainly transgranular fracture that is observed in the FZ (Figure 9c). The differences between the crack growth rates in the FZ and the BM, as shown in Figure 7b, become smaller at high AK
10
20
30
aK (MPafm)
Figure7
Fatigue crack growth behavior for T i - 6 - 4 laser welds with the crack growth (a) normal to and (b) parallel to the welding direction
values. The reduced closure level, which is induced by crystallographic fracture, could account for this convergence of crack growth rates. CONCLUSIONS 1. The excellent tensile properties of the laser welds are comparative to those of the BM, and can be attributed to the presence of martensite in the narrow FZ.
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
719
Figure8 SEM fractographs showing (a) transgranular fracture in the base plate, (b) the FZ at low magnification, (c) the BM at low magnification, (dl crystallographic induced fracture in the FZ and (e) the fracture surface of the HAZ
2. The FCGRs of the T i - 6 - 4 plates subjected to different heat treatment conditions fall within a narrow data band. In general, the 500 and 600°C aged specimens exhibited the highest and lowest FCGRs, respectively. In addition, the FCGRs increased obviously when R was increased form 0.1 to 0.4. The fracture surfaces were characterized by fatigue striations at high AK values. 3. The decomposition of martensite in the aged FZ retards the crack growth most among various regions in a weld, which is characterized by a mixed mode fracture. While, the HAZ reveals less
portions of cleavage facets as compared to those in the FZ. 4. The crack path in the FZ and HAZ is more tortuous than in the BM. Hence, the roughness-induced crack closure which is caused by crystallographic fracture would result in the decrease in FCGRs. ACKNOWLEDGEMENTS The authors gratefully acknowledge the support of the republic of China National Science Council (Contract No.85 NSC-2216-E-019-002).
L. W. Tsay and C. Y. Tsay
720
10
II
12
13
14
15 16 17
18
19
20
21
22
23 Figure 9 SEM fractographs of Ti-6--4 laser welds showing (a) a low magnification of the fracture appearancc, (b) mixed fracturc modes at low AK ranges and (c) transgranular fracture at high AK ranges (notc that crack growth parallel to the welding direction)
24
REFERENCES 1
2
3
4
Peters, M. and Williams, J. C., Microstructure and mechanical properties of a welded (o~ + /3) Ti alloy. Metallurgical Tronsactions. 1984, 15A(8), 1589-1596. Yuen, A., Hopkins, S. W., Leverant, G. R. and Rau, C. A.. Correlations between fracture surface appearance and fracture mechanics parameters for stage 11 fatigue crack propagation in Ti-6AI V. Meiallurgical Tran,vaclions. 1974, 5(8), 1833 1842. Ravichandran, K. S. and Dwarakadasa, E. S.. Fatigue crack growth transitions in Ti 6A1-4V alloy. Scripta Metallur, q,ica, 1989, 23(10), 1685-1690. Ravichandran, K. S., Fracture mode transitions during fatigue crack growth in Ti 6AI-4V alloy. Scripta Metallul~ica, 1990, 24(7), 1275-1280.
25
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Ravichandran, K. S., Dwarakadasa, E. S. and Baneljec, D., Mechanisms of cleavage during fatigue crack growth in Ti 6AI4V alloy. Seripta Metallurgica, 1991. 25(9), 21 I5 2120. Yoder, G, R., Cooley, L. A. and Crookcr, T. W., Observations on microstructurally sensitive fatigue crack growth in Widmanstiitten Ti-6AI 4V alloy. Metallurgiccd Transactions'. 1977, 8A(ll), 1737-1743. Yoder. G. R., Cooley, L. A. and Crooker. T. W., Enhancement of fatigue crack growth and fracture resistance in Ti 6A14V and Ti-6AI~-V-2Sn through micrnstrucmral modification. Transactions q/'the ASME, 1977, 99(10), 313-318. Yoder. G. R., Cooley, L. A. and Crooker, T. W.. Fatigue crack propagation resistance of beta-annealed Ti 6AI-4V alloys of differing interstitial oxygen contents. Metalhtr~ica/"])'an.Saclions, 1978, 9A(10), 1413-1420. Yoder, G. R., Cooley, L. A. and Crooker. T. W.. 50-Fold difference in region-ll fatigue crack propagation rcsistancc of titanium alloys: a grain-size effect. Tram'aetion~ o/flu, ASME, 1979. 101(l), 86 90. Thomas, G., Ramachandra, V.. Nagaralan. K. V.. Pant, B., Sarkur, B. K. and Vasudcvan. R., Electron beam welding studies of 25 mm thick Ti-6AIMV sections. WehlilLq .IottpTtal. 1989. 68(8), 336s 341s. Mazumder, J. and Stcen, W. M., Welding of Ti 6AI-4V by a continuous wave CO~ laser. Metal Conxlrtwtion, 1980, 12(9), 423~1-27. Welsch, G. and Bunk, W., Deformation modes of the (x-phase of Ti-6AI-4V as a function of oxygen concentration and aging temperature. Metallurgical Tram'actions, 1982, 13A(5). 889-899. Chen, S. J. and Devletian, J. H., Microstructure and mechanical properties of electroslag welds in Ti 6AI-4V alloy. WelditL~ Journal, 1990, 69(9), 319s-325s. Mazumder, J. and Steen, W. M., Microstructure and mechanical properties of laser welded titanium 6AI-4V. Metalhtr~iea/Transactions, 1982, 13A(5), 865-871. Denney. P. E. and Metzbower, E. A., l,aser beam welding of titanimn. WeMing Journal, 1989, 68(8), 342s 346s. AWS, Welding Handbook, Metals aml Their Weldttbilitv. 7th edn, Voh 4. AWS, Miami, 1982. pp. 434-470. ASTM. Stamlard I),st Method ,/or 7"ensio~l Tcsli&q O/Metallic Mawrial.~ [Metric], Annual Book of ASTM Standards, ESM90a, Vol. 03.01. ASTM. Pennsylvania, 1991, pp. 146 160. ASTM, Standard Test Method .fi)r meas'urement q! I:ati~ue Crack Growth Rates, Annual Book of ASTM Standards, E647 91, Vol. 03.01. ASTM, Pennsylvania, 1991, pp. 654 681. Saxena, A. and Hudak, S. J., Review and extension of compliance information |or common crack growth specimens, hm'rnational Journal of Fracture, 1978, 14(5), 453-468. ASTM, Stgmdard Test Method./or Phme Stroin I:ructurc ]?m~h*tess of Metallic Material~, Annual Book of ASTM Slaudards, E399-90, Voh 03.01. ASTM, Pennsylvania, 1991, pp. 485 515. Thomas, G., Ramachandra. V.. Nair, M. J., Nagarz~ian, K. V. and Vasudevan, R., Effect of preweld and poslweld hcat treatment on the properties of GTA wclds in Ti-6AI 4V shecL Wehlinv, Journal, 1992, 71(1), 15s 20s. Baeslack, W. A. and Banas, C. M., A comparative evaluation of laser and gas tungsten arc weldments in high-tcmperalure titanium alloys. Weldin~ ,Ioltr/ldl. 1981. 60(7k 1215--1305. Huang, C. C., Pan, Y. C. and Chuang. T. H., Efl~_,cls of po,,I weld heat treatments on the residual stress and inechanical prnperties of electron bean welded SAE 4130 steel plates. ,hmr #ml q/' Materials En
ELSEVIER
PIh S0142-1123(97)00113-8
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds L. W. Tsay* and C. Y. Tsay Institute of Materials Engineering, National Taiwan Ocean University, Keelung, Taiwan, Republic of China (Received 20 May 1997; revised 18 August 1997; accepted 21 August 1997) The effect of microstructures on fatigue crack growth rates (FCGRs) in Ti-6AI-4V plates and laser welds was investigated. Experimental results revealed that the location of tensile fracture of laserwelded specimens was always in the base metal, owing to the presence of martensite in the narrow fusion zone (FZ). For welds aged at various temperatures, only minor differences in tensile properties were observed. The various heat treatments of such plate did not affect FCGRs significantly. The fracture surface was characterized by fatigue striation at high AK values. When the R ratio was increased from 0.1 to 0.4, higher FCGRs would be resulted. The decomposition of martensite in the aged FZ causes retardation of FCGRs in a weld, which is characterized by a mixed mode fracture. The more tortuous crack path within the FZ and HAZ was also indicative of high crack growth resistance. Roughness-induced crack closure caused by crystallographic fracture was responsible fi~r decreased FCGRs. © 1998 Elsevier Science Ltd. All rights reserved (Keywords: Ti-6AI--4V; laser welding; fatigue c r a c k growth rates; fatigue striation; cleavage fracture)
INTRODUCTION
larger packet size in the structure results in lower FCGRs s,'. Because of the high temperature reactivity of titanium alloys, electron beam (EB) welding is the most popular method for joining Ti-6AI~4V with high quality"( However, EB welding is often limited by the size of the vacuum chamber. Laser beam (LB) welding can provide welds of similar quality as compared to those of EB welds". Raising the oxygen content in the T i - 6 - 4 alloy is known to increase the tensile strength while decreasing tensile ductility ~2. Interstitial element (O, N) contamination can further decrease tensile ductility of the weld metal ~. it was found that no significant oxygen contamination occurred in laser welding with effective shielding in the experiment '4. Preparation of the surface to be welded, which removes the titanium oxide from the surface, does not affect the oxygen content in the titanium welds ~5. Relatively few works were performed to investigated the fatigue crack growth behavior of T i - 6 - 4 welds. The present work was undertaken to investigate the effect of microstructures on the mechanical properties in T i - 6 - 4 plates and laser welds. FCGRs were measured in various regions of the welds and particularly paid attention to the regions near the fusion zone (FZ). Both the microstructural and the stress ratio (R) effects were also evaluated. Fractographic observations were performed on various specimens after different heat treatments. Meanwhile, the variations of FCGRs in the welds could be correlated well with the changes in the fracture mode.
Ti-6A1-4V alloy is widely used in the aerospace industries for its remarkable strength:weight ratio and resistance to high temperature creep. This alloy can be heattreated to obtain various microstructures. The shape and distribution of the alpha (o0, beta (/3) and martensite (c~') are considered to be closely related to the mechanical properties of the material. Peter and Williams L indicated that microstructural changes can produce significant variations in crack growth rates at low AK values for Ti-6A1-2V-Mo. Above a critical value of AK, the crack growth is found to become structure insensitive 2. The criterion for the fracture transition from structure sensitive to insensitive is considered to be a change from primary single to multiple slip within the individual grain at the crack tip 2. The growth mode also changes when the reversed plastic zone size exceeds the scale of the microstructure, that is, the Widmanstfitten colonies or alpha laths ~:. Different microstructures exhibit different characteristic growth rates in structure sensitive regions '2-4 % Previous research indicated that significant reduction in region II growth rates for titanium alloys can be obtained by microstructure modification 7 ". Fatigue crack growth rates (FCGRs) can be substantially reduced for /3-anneal specimens, in which crack bifurcation occurs in the Widmanstfitten packet 6,7. The
*Author Ik)r correspondence.
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welds at various temperatures in vacuum tk~llowed by air cooling. The suggested PWHT is in the temperature range of 538-593°C (1000-1100°F) for 2 4 h ~ % In the present study, PWHTs were at 500, 550 or 600°C for 3 h. The mechanical tests in the study included microhardness measurements, tensile and fatigue crack growth tests. Microhardness measurements were taken across the weld to compare hardness variations in the FZ, heat-affected zone (HAZ) and unaffected base metal (BM). Tensile specimens, with a gage length of 25 mm and a nominal width of 6 ram, were made according to the specifications of ASTM E8M-90a (subsize, 2.5 mm thick) 17. The tensile strain rate was set at 7 x 10 4s ' and the results were the average of at least three specimens for each testing condition. Compact-tension (CT) specimens with the rolling direction (RD) normal to the crack path (see Figure 2) were employed. The experimental procedures met the standard of ASTM E647-91 ~* specifications. The applied load had a constant amplitude sinusoidal waveform and the loading frequency was 20 Hz. The stress ratio, that is, the ratio of minimum to maximum load, was set at 0.1 or 0.4. The crack growth directions were aligned either parallel or normal to the welding direction in the welded CT specimens (Figure 2). When the crack path was normal to the welding direction, the distance between the notch tip and the bottom fusion line of the weld was about 6.0 mm (Figure 2b). Specimens used in the above tests were prepared as follows: 12.5 + 0.1 Figure 1 SEM metallographs of T i ~ - 4 plates in the (a) as-received and (b) solution-treated conditions 30 0.1
M A T E R I A L AND E X P E R I M E N T A L PROCEDURES The chemical composition by weight percentage of the 3.1 m m thick T i - 6 A 1 4 V alloy was 6.20 A1, 4.22 V, 0.14 Fe, 0.01 C, 0.0136 O and 0.007 N and balanced Ti. The microstructure of the as-received T i - 6 - 4 consisted of small percentage of /3 phase distributed at the elongated c~ grain boundaries (Figure la). The alloy was solution-treated in vacuum at 950°C for 40 min and then cooled subsequently by Ar quench. Because the material was held below the /3 transus temperature, the microstructure consists equiaxed primary c~ and transformed /3 (Figure lb). All specimens were welded in the solution-treated condition. A Rofin-Sinar RS 850 5 KW CO2 laser was utilized for laser welding. Table 1 lists the laser welding variables used in this study. After welding, postweld heat treatments (PWHTs) were performed on the
Table 1
Laser welding parameter used in the experiment
Laser power Travel speed Focal lens Focal length Focal point Plasma assist gas flow rate Backing gas flow rate
2500 W 1500 m m rain ZnSe 190.5 mm 0.5 m m underfocus (0.5 mm below the surface) 25 lpm He 15 lpm Ar
,37s-+0, / RD
~
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1 62.54- 0.1
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1] Unit • mm Figure 2 Schematic representation showing CT specimens employed in the fatigue crack growth test crack growth (a) parallel and (b) normal to welding direction. RD is the rolling direction o1" the plate and FZ represents the fusion zone
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
715
the specimen's profile was cut using a electrodischarge machine with a wire electrode after which the specimens was ground to the required thickness. Metallographic observations and microhardness measurements were made to identity each region of the weld. Crack growth across the FZ can reveal the variation of crack growth rates in the different regions of the weld. The compliance function proposed by Saxena and Hudak ~° is used to determine the crack length, , / W = C,, + C,U, + C~_U~ + C~U~ + CaU ~, + C~U 5, where U, = I/[(BEWP)°5 + I] Here ~l is the crack length, W the specimen width, B the specimen thickness, E Young's modulus, P the load, V the crack opening displacement and Co, CI, C2, C~, Ca, C5 are compliance coefficients. The calculation of AK was made according to2°: A K = [A P / B W 1/2][f(a/W) ] f ( a / W ) = [(2 + a/W)/(I - a/W)3/2][0.886 + 4.64(a/W) -
13.32(a/W) 2 + 14.72(a/W) 3 - 5.6(a/W) a]
where AP is the loading range. Fractographic examinations were carried out by a scanning electron microscope (SEM). Thin foil specimens for transmission electron microscope (TEM) were prepared by a standard jet-polisher in a electrolyte consisting of 95% methyl alcohol and 5% sulfuric acid before being examined by a JEOL-2000EX microscope operated at 160 kV. RESULTS AND DISCUSSION MicrostructuraI observation The metallograph of the as-welded FZ Figure 3a), reveals the microstructure is a coarse /3 grain with internally acicular morphology. TEM micrographs show that the FZ consists of mainly oe' martensite (Figure 3b) as compared to granular a and refined lamellar c~ + /3 in the BM (Figure 3c). The acicular oe' in the FZ of laser welds was attributed to rapid cooling of the molten weld pool during solidification. In addition, the microstructure in the HAZ would also depend on the thermal history of welding. The microstructure of the HAZ in the vicinity of the fusion boundary was almost the same as the FZ, while regions of the HAZ far away from the fusion boundary were heated to relatively low temperatures and had a microstructure similar to that of the BM. The structure in the middle region of the HAZ was found to be a mixture of primary oe and oe'. Microstructures of the FZ aged at various temperatures were similar and revealed the presence of decomposed martensite. Figure 4 shows that the precipitation of /3 from the martensite was located primarily at the lath boundaries. Hardness measurement Figure 5 is a series of schematic representations showing the microhardness distribution in the welds. The hardness of the FZ and HAZ is higher than that of the BM. The highest hardness is within the FZ while the hardness in the HAZ drops as the distance from the fusion boundaries increases. These results are consistent with the microstructural observations, that
Figure 3 Photographs revealing the: (a) as-welded structure of the FZ; (b) TEM micrograph of (at; and (cl TEM microgruph of the BM in the solution-treated condition is, there is a refined microstructure with a higher hardness in the FZ. In addition, the hardness in the FZ could be raised slightly after PWHTs. The slight increase in hardness in the FZ after PWHTs was associated with the decomposition of ¢x' martensite to form c~ + /3, as shown in Figure 4. it is noted that only minor differences in hardness of the FZ were found among welds aged at different temperatures. The BM was also slightly hardened after aging, and this could be correlated with the precipitation of fine a from small quantities of metastable /32t, but the BM remained softer than the other region. Thus, fracture of welded specimens in this region can be anticipated.
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Tensile properties The room temperature tensile properties of T i - 6 - 4 base plates are listed in Table 2. As-received specimens with the RD normal to the tensile axis exhibited the lowest strength among the various specimens being tested, while the base plates aged in the temperature range of 500-600°C had only a slightly higher strength than the solution-treated specimens. The increase in hardness in the BM after aging, as shown in Figure 5, accounts for a slightly increase in strength in the aging temperature range of concern. Furthermore, the variations in tensile properties is minor for the variously aged specimens. Table 3 displays the tensile results of laser welds aged at different temperatures. It is important to mention that necking of tensile specimens was observed for all aged welds and that the fracture was located within the BM regardless of PWHTs. Tensile properties remained more or less the same for all the welds. During tensile testing, the FZ, HAZ and BM across the gage length underwent a nonuniform deformation which led to a slight loss in ductility. However, it implied that laser welds have consistent tensile properties as compared to those of the base plates. In the case of T i - 6 - 4 electroslag welds, ductility loss is correlated to the large prior /3 grain size and the microstructure produced by the high heat input j3. Laser welds show superior ductility and this has been attributed to fine prior /3 grain size 22. In the present study, the presence of martensite in the narrow FZ of laser welds might be the reason for these
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excellent properties. Tensile-fractured surfaces examined by SEM revealed a ductile dimple fracture for all the specimens.
Fatigue crack propagation The FCGR (da/dN) plotted against the stress intensity factor range (AK) curve of T i - 6 - 4 base plates subjected to different heat treatments are shown in Figure 6. The fatigue crack growth tests reveal da/dN values fall within a narrow band for various specimens being tested (Figure 6a). In general, the 600°C aged specimens demonstrated the lowest crack growth rates, while the 500°C aged specimens had slightly higher crack growth rates than those of other specimens. Figure 6b shows the effect of increasing R from 0.1
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
717
Table 2 Tensile properties of Ti-6AI~4V plate Specimen condition
U.T.S. (MPa)
As-received~' As-received~' Solution 950°C per 40 min Aged 500°C per 3 h Aged 550°C per 3 h Aged 600°C per 3 h
954 936 957 980 977 970
Y.S. (MPa)
Elongation (%)
910 916 854 894 896 895
Reduction in area (~/~)
17.2 17.6 17.1 15.2 15.6 16.3
31.2 31.9 30.7
28.2
32.7 33,4
"Tensile axis parallel rolling direction bTensilc axis normal to rolling direction Table 3
Tensile properties of Ti-6AIqV laser welds
Specimen condition
U.T.S (MPa)
Y.S. (MPa)
Aged 500°C Aged 550°C Aged 600°(7
971 978 974
890 891 886
Elongation (~/~) 14.5 14.6 15.0
Reductionin area Ic,4) 30.6 3 I.I 31.9
Fracturelocation Base metal Base metal Base metal
All tensile specimen had been laser welded then aging treatmenl to 0.4 on FCGRs of as-received and 550°C aged specimens with similar ultimate tensile strength. Clearly FCGRs increase as R is increased. Fatigue crack growth behavior of Ti-6--4 laser welds after aging is shown in Figure 7. In Figure 7a the crack path is normal to the welding direction. The total width of the FZ and HAZ, as measured in Figure 5, is about 4 ram. Therefore, as the crack propagates across the FZ. the region of concern is only within a limited range of AK. When the crack grows from the BM into the HAZ, the FCGRs decrease immediately. After further growth of the crack, lower FCGRs are observed in the FZ. As the crack re-enters into the BM, the FCGRs return to levels similar to those for base plates subjected to the same aging treatment. Thus the results reveal that the FZ retards the crack growth most among various regions in a weld. Retardation of crack growth in the FZ is not so obvious at R = 0.4, however. As mentioned above, the microstructure of the HAZ in various locations would depends on its thermal history. The low temperature H A Z has a microstructure corresponding to the BM and thus has similar fatigue properties to that of the BM. However, the microstructure of the HAZ in the vicinity of the FZ consists predominantly of martensite, hence its fatigue properties will correspond to those of the FZ. As can be seen in Figure 6a, the initial lower FCGRs at low AK values are obtained in the BM, an area which does not undergo any microstructual changes. The HAZ and FZ always show residual tensile stress in the direction parallel to the welding direction, while residual compressive stresses are developed in the BM 23 25. Residual welding stresses can be reduced by P W H T at high temperature 25. The residual compressive stresses that exist in the BM can, however, effectively reduce the crack growth rate in this region 25. In addition, at smaller AK values or lower R ratios, the effect of residual stresses on the da/dN vs AK relationship is more pronounced 26. Horikawa et al. 2v also indicated that under these conditions residual welding stresses have a great influence on the crack initiation. As shown in Figure 7a the crack growth rates are low at low AK ranges, and this can be attributed to the existence of residual compressive stresses in the BM.
Thomas et al. 2~ indicated that a low temperature PWHT, such as stress relief or aging treatment, results in a deterioration in the fracture toughness of the FZ for T i - 6 4 EB welds. However, an investigation of T i - 6 A I - 2 V - M o TIG welds ~ revealed lower strength and ductility, resistance to fatigue crack growth in the HAZ and FZ was higher than in the BM. In the present study, T i - 6 4 laser welds were found to have not only good tensile properties but also high resistance to fatigue crack growth. As the crack grows in the FZ and parallel to welding direction for the 550°C aged welds (Figure 7b), crack growth resistance in the FZ is significantly better than in the same aged base plate. However, the difference in FCGRs between base plate and 550°C aged welds becomes less marked with increasing AK. The results also revealed an obvious increase in FCGRs was observed for the 550°C aged welds as R was increased from 0.1 to 0.4.
Fractographic observations SEM fractographs of T i - 6 - 4 base plates and laser welds were shown in Figure 8. Fatigue fractographs of T i - 6 - 4 base plates after different heat treatment were characterized by transgranular fractures (Figure 8a). Fatigue striations become clearer at high AK values. As the crack grows across the FZ, low magnification of the fracture surface reveals that the FZ has a remarkable rough surface for aged laser welds (Figure 8b). However, the crack path ill the same aged T i - 6 - 4 base plate is fairly smooth (Figure 8c). A mixed mode fracture with coarse cleavage facets was observed in the FZ (Figure 8d), as opposed to the striations that covered a large part of the fracture surface in the BM. In the investigation of fatigue crack growth transition in T i - 6 - 4 alloy, significant portions of cleavage facets was characterized by a high level of crack closure 3. The fracture surface of the coarse grain HAZ also shows a mixed mode fracture with less cleavage facets (Figure 8e), as compared to that of the FZ. In summary, the crack path is more irregular and tortuous in the FZ and HAZ than in the BM. The changes of growth direction are assisted by crystallographic fracture. Thus, the overall decrease in FCGRs can be attributed to the effect of roughness-induced
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crack closure. As the crack propagated along the FZ, the overall fracture appearance at low magnification also suggested that there was a relatively unfavorable orientation for crack propagation (Figure 9a). Transgranular and cleavage fractures coexisted on the fracture surface (Figure 9b). However, The extent of crystallographic-induced fracture is decreased with increasing_stress intensity factor range. When AK is > 20 MPa gm, it is mainly transgranular fracture that is observed in the FZ (Figure 9c). The differences between the crack growth rates in the FZ and the BM, as shown in Figure 7b, become smaller at high AK
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Fatigue crack growth behavior for T i - 6 - 4 laser welds with the crack growth (a) normal to and (b) parallel to the welding direction
values. The reduced closure level, which is induced by crystallographic fracture, could account for this convergence of crack growth rates. CONCLUSIONS 1. The excellent tensile properties of the laser welds are comparative to those of the BM, and can be attributed to the presence of martensite in the narrow FZ.
The effect of microstructures on the fatigue crack growth in Ti-6AI-4V laser welds
719
Figure8 SEM fractographs showing (a) transgranular fracture in the base plate, (b) the FZ at low magnification, (c) the BM at low magnification, (dl crystallographic induced fracture in the FZ and (e) the fracture surface of the HAZ
2. The FCGRs of the T i - 6 - 4 plates subjected to different heat treatment conditions fall within a narrow data band. In general, the 500 and 600°C aged specimens exhibited the highest and lowest FCGRs, respectively. In addition, the FCGRs increased obviously when R was increased form 0.1 to 0.4. The fracture surfaces were characterized by fatigue striations at high AK values. 3. The decomposition of martensite in the aged FZ retards the crack growth most among various regions in a weld, which is characterized by a mixed mode fracture. While, the HAZ reveals less
portions of cleavage facets as compared to those in the FZ. 4. The crack path in the FZ and HAZ is more tortuous than in the BM. Hence, the roughness-induced crack closure which is caused by crystallographic fracture would result in the decrease in FCGRs. ACKNOWLEDGEMENTS The authors gratefully acknowledge the support of the republic of China National Science Council (Contract No.85 NSC-2216-E-019-002).
L. W. Tsay and C. Y. Tsay
720
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13
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15 16 17
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23 Figure 9 SEM fractographs of Ti-6--4 laser welds showing (a) a low magnification of the fracture appearancc, (b) mixed fracturc modes at low AK ranges and (c) transgranular fracture at high AK ranges (notc that crack growth parallel to the welding direction)
24
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2
3
4
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