Author’s Accepted Manuscript Microstructural evolution and mechanical properties of electron beam welded dissimilar titanium alloy joints S.Q. Wang, W.Y. Li, K. Jing, X.Y. Zhang, D.L. Chen www.elsevier.com/locate/msea
PII: DOI: Reference:
S0921-5093(17)30625-1 http://dx.doi.org/10.1016/j.msea.2017.05.028 MSA35044
To appear in: Materials Science & Engineering A Received date: 6 March 2017 Revised date: 15 April 2017 Accepted date: 7 May 2017 Cite this article as: S.Q. Wang, W.Y. Li, K. Jing, X.Y. Zhang and D.L. Chen, Microstructural evolution and mechanical properties of electron beam welded dissimilar titanium alloy joints, Materials Science & Engineering A, http://dx.doi.org/10.1016/j.msea.2017.05.028 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Microstructural evolution and mechanical properties of electron beam welded dissimilar titanium alloy joints
S.Q. Wang1*, W.Y. Li2, K. Jing1, X.Y. Zhang1, D.L. Chen3* 1
School of Materials Science and Engineering, Xi’an Shiyou University, 18 Dianzi Road, Xi’an,
Shaanxi, 710065, PR China 2
State Key Laboratory of Solidification Processing, Northwestern Polytechnical University, 127
Youyi Road, Xi’an, Shaanxi 710072 , PR China, 3
Department of Mechanical and Industrial Engineering, Ryerson University, 350 Victoria Street,
Toronto, Ontario M5B 2K3, Canada *
Corresponding author. Tel.: +86 29 88382607. [email protected]
Corresponding author. Tel:
(416) 979 5000 ext. 6487. [email protected]
Abstract The aim of this study was to analyze the microstructural changes of two dissimilar Ti-6Al4V/Ti17 and Ti-6Al-4V/BT9 joints welded via electron beam welding (EBW), and establish the relationship between microstructure and properties. The results showed that the occurrence of phase transformation was mainly dependent on the peak heating temperature during EBW, and despite the same welding parameters the microstructures of two dissimilar joints were different except in the heat-affected zone (HAZ) on the Ti-6Al-4V side. Both dissimilar joints exhibited a higher strength compared with Ti-6Al-4V base metal (BM), and the strength of Ti-6Al-4V/Ti17 joint was slightly higher than that of Ti-6Al-4V/BT9 joint due to the presence of finer martensite. Fatigue life of dissimilar joints was almost the same as that of Ti-6Al-4V BM within the
experimental scatter. Fatigue failure of both dissimilar joints occurred in the HAZ, and fatigue crack basically initiated from the specimen surface.
Keywords Titanium alloy; electron beam welding; dissimilar joints; microstructure; tensile properties; fatigue resistance
With the development of science and technology and the applications of new material and new technology, the overall performance of components needs to meet more complex requirements. The parts consisted of a single metal are increasingly difficult to satisfy the intricate conditions of use, however the dissimilar metal components joined by the advanced welding technology not only provide the advantages of respective alloys to meet various needs, but also greatly reduce the overall cost of production to significantly improve the economic benefits. Therefore, dissimilar metal components are becoming the development trend of mechanical components in the future. Especially, to promote the development of integrated engines, it is necessary to connect the more expensive high-temperature titanium alloys with the relatively low-cost medium temperature titanium alloys or high-strength titanium alloys with medium strength titanium alloys in the aero gas turbine industry to increase efficiency and cost savings. Though there are a lot of advantages for the dissimilar metal components, it also brings many challenges to the welding technology. For example, although the dissimilar joints welded by electron beam welding , laser beam welding , friction welding , etc., can ensure the basic mechanical properties, in the dissimilar metal welding process, due to different thermal physical properties,
such as thermal conductivity, the inhomogeneous microstructure will be present in the joint from the heat-affected zone (HAZ) to the weld metal (WM), which will deteriorate the mechanical performance of parts. At the same time, defects formed in the welding process may become the site of crack initiation, affecting the structural integrity of the whole components. Therefore, tensile and fatigue properties have to be taken into consideration in terms of the selection, development, and design of materials for the welded joints subjected to service loading and high strain rate deformation. Particularly, the tensile and fatigue behavior of these welded joints including the WM should be closely examined and evaluated before the parts are used. Thus far, several efforts have been devoted to build the relationships between the microstructure and mechanical properties of the welded joints.
Studies have focused on the influence of welding process parameters on the structures and mechanical properties or at the given welding process parameters how the structures affect strength, plasticity, toughness, and fracture mechanism of the welded dissimilar titanium alloy joints. Zhang et al.  discussed the weldability of dissimilar joining between Ti3Al and TC4 via electron beam, and showed that the grain size significantly affected the tensile strength of the dissimilar joint, which was due to the heat inputs during welding. With increasing heat input, the grain size increased significantly, while the composition of the WM was independent of heat input. Tan et al.  observed that the electron beam welded dissimilar joint of Ti-22Al25Nb/TC11 exhibited higher tensile strength than that of the TC11 alloy base, and the impact toughness of the dissimilar joint reached about 42% of that of the TC11 base alloy. They revealed that the microstructures and the mechanical properties of dissimilar joints relied obviously on the β phase stabilizer content. Pederson et al.  evaluated tensile-, creep-, and
Charpy V-notch impact tests of the dissimilar titanium alloy joints (Ti64/Ti6246) welded via the electron beam, and observed that all tensile- and creep tested Ti64/Ti6246 welds were independent of post-weld heat-treated temperature. Their results indicated that the post-weld heat treatment could only slightly improve the ductility in the dissimilar titanium alloy welds. Wang et al.  demonstrated the relationship between the strain rate, temperature and the strain hardening behavior of the electron beam welded dissimilar Ti-6Al-4V/Ti17 joint, for example, the hardening rate was strongly dependent on the strain rate and temperature. As the strain rate increased or temperature decreased, the strain hardening rate increased at a given true stress. Though quasi-static tensile properties are required for the initial fracture, damage tolerance, actually in the real-world applications welded joints are often subjected to dynamic fatigue loading, and hence it is crucial to study the fatigue behavior of these dissimilar welded joints to guaranteeing the structural integrity of components. Wang et al. [8-11] investigated cyclic deformation behavior of the electron beam welded dissimilar Ti-6Al-4V/Ti17, Ti-6Al-4V/BT9, and Ti-6Al-4V/IMI834 joints by changing strain amplitudes and strain ratios. The results illustrated that a strain amplitude of 0.6% appeared to be a critical value, at or below which cyclic stabilization or saturation remained in the entire cyclic deformation process, and the fatigue life of the dissimilar joints were nearly the same as that of the base metals (BMs).
As reviewed above, welding process parameters, microstructure, tensile property, impact toughness, creep behavior, and strain-controlled fatigue resistance of several electron beam welded dissimilar titanium alloy joints such as Ti3Al/TC4 , Ti-22Al-25Nb/TC11 , Ti64/Ti6246 , and Ti-6Al-4V/Ti17 [7-9] have been well studied and reported in the literature. However, limited data are available on the stress-controlled fatigue behavior of electron beam
welded dissimilar titanium alloy joints. In addition, the S-N curves of cyclic stress amplitude (S) versus the number of cycles to failure (N) and fatigue limit should be also taken into consideration in terms of the selection, development, and design of materials for the manufacturing of aero-parts subjected to service loading. Particularly, there remains a need to closely evaluate the tensile and fatigue properties of these critical load-bearing parts containing the WM and HAZ prior to their actual application. Therefore, this study focuses mainly on evaluating the microstructure and stress-controlled fatigue properties of two dissimilar titanium alloy joints, and establishing the relationships among the welding process, microstructure, and property.
Material and Experimental Procedure
The alloys used in the present investigation are forged Ti-6Al-4V (TC4), Ti-5Al-4Mo-4Cr-2Sn2Zr (Ti17), and Ti-6.5Al-3.5Mo-1.5Zr (TC11) titanium alloys. The alloys were machined into plates with a thickness of 10 mm. To remove the oily substances from the surface, organic solvent such as acetone and alcohol was used. The heavier oxide films were removed by wire brushing, acid pickling before welding, then the BMs were clamped by the fixture in the vacuum chamber. After evacuating the equipment, EBW was conducted using TECHMETA LARA 52 EBW machine with the welding parameters listed in Table 1. To compare the microstructure and the mechanical properties of the dissimilar joints of Ti-6Al-4V/Ti17 and Ti-6Al-4V/BT9, the welding parameters were kept to be the same for these two dissimilar joints. Since the thickness was 10 mm, the electron beam was focused under the top surface to ensure a deep penetration.
To reveal the microstructure of different zones, metallographic specimens were cut from the EBWed workpieces perpendicular to the welding direction, then mounted, ground with SiC sand papers, and polished using 9 m diamond paste. Then the polished specimens were etched for 6 seconds in Keller’s reagent. Microstructures were observed by a scanning electron microscope (SEM). Tensile and fatigue test samples with a gauge length of 25 mm and a thickness of 1 mm were machined with the weld positioned in the middle of the gauge length using electrodischarge machining (EDM). The gauge area was ground along the loading direction with SiC sand papers up to a grit number of 600 to remove the EDM cutting traces and to obtain a smooth surface. Uniaxial tensile tests were conducted using a computerized United tensile testing machine at a strain rate of 1×10-2 s-1 and three samples were tested. To evaluate the stress-life (SN) fatigue properties of the dissimilar joints, fatigue tests were performed at room temperature in air on a fully-computerized Instron 8801 servo-hydraulic testing system under load control at a stress amplitude from 150 MPa to 500 MPa with an interval of 50 MPa. To avoid the occurrence of buckling for such welded thin sheets, tension-tension cyclic loading at a stress ratio of R=0.1 was applied at a frequency of 50 Hz and sinusoidal waveform. At least two specimens were tested at each stress amplitude. The fatigue fracture surfaces were observed via SEM to identify fatigue crack initiation sites and crack propagation mechanisms.
Results and Discussion
Microstructures of BM and joint
The microstructure of Ti-6Al-4V, Ti17, and BT9 base alloys is presented in Fig. 1, and all of them had a typical bimodal microstructure. Ti-6Al-4V and BT9 alloys consisted of equiaxed α grains and inter-granular α + β lamellae, as indicated by arrows, while Ti17 alloy was composed
of ﬁne α plates and β plates (Fig. 1(b)). Furthermore, Ti-6Al-4V alloy (Fig. 1(a)) contained more equiaxed α than BT9 alloy (Fig. 1(c)).
After EBW, the surface profile of Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint is displayed in Figs 2 and 3, respectively. The uniform ripples were observed on the front of both dissimilar joints as indicated by arrows in Fig. 2(a) and Fig. 3(a). A full penetration was achieved, as seen from the back of two dissimilar joints as shown in Fig. 2(b) and Fig. 3(b). No defects, such as undercut, porosity, and cracking, were observed on the joint surface (Figs 2 and 3), suggesting that a good surface was obtained in both dissimilar joints. An overall view of the cross section of Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint is shown in Fig. 4. It was observed that electron beam welding generated a fairly narrow WM and HAZ. The WM of both dissimilar joints exhibited a typical nail-shaped, and the microstructure could be divided into three zones: WM, HAZ, and BM. A clear borderline among BM, HAZ, and WM was clearly visible in Fig. 4, revealing a significant microstructural change that occurred across the dissimilar welded joints due to the presence of temperature gradients during EBW. In addition, the size of columnar crystals on both dissimilar joints decreased gradually from the top to the bottom. This was related to the absorbed heat. That is, the absorbed heat on both dissimilar joints decreased from the top to the bottom, giving rise to a reduced size of columnar crystals. Furthermore, the growth direction of columnar crystals on both dissimilar joints was similar, and had their own characteristic at different thicknesses. The columnar crystals at the top grew upwards, while they grew to the weld center at the middle. The different growth directions of columnar crystals were directly related to the direction of thermal conduction. The direction of thermal conduction was upwards at the top, while it was perpendicular to the puddle wall as indicated by arrows in Fig. 4
at the middle of joint. Therefore, the growth direction of columnar crystals was different at different locations.
Fig. 5 illustrates the microstructures of HAZ and WM of EBWed Ti-6Al-4V/Ti17 joint. On the Ti-6Al-4V side, in the outer-HAZ small martensite and some α phase were observed (Fig. 5(a)), while in the inner-HAZ, only martensite at the left of Fig. 5(b) formed during electron beam welding. This observation corresponded well to that reported by Rao et al.  who also observed the martensite in the electron beam welded Ti6Al4V joint. In the WM, the coarse columnar crystal was composed of fine martensite in the Fig. 5(b) and (c). On the Ti17 side, coarse β phase (Fig. 5(d)) appeared in the inner-HAZ. The blurry fine α phase and β phase were obviously visible in the outer-HAZ (Fig. 5(e)). Fig.6 shows the microstructures of HAZ and WM on EBWed Ti-6Al-4V/BT9 joint. The microstructure in the HAZ on the Ti-6Al-4V side was similar to that of in the HAZ on the Ti-6Al-4V/Ti17 joint. The outer-HAZ was also composed of martensite and α phase (Fig. 6(a)), while both the inner-HAZ on the Ti-6Al-4V side and the WM mainly contained martensite (Fig. 6(b), (c), and (d)). On the BT9 side, martensite and retained α phase were present in the HAZ as indicated by the arrows in the Fig. 6(e) and (f).
In the Fig. 5(b) and Fig. 6(b), while the size of martensite in the HAZ on the Ti-6Al-4V side for both dissimilar joints was almost same, it was remarkably finer in the WM of Ti-6Al-4V/Ti17 joint than that of Ti-6Al-4V/BT9 joint with the same welding parameters. This was associated with the element reorganization in the puddle and the thermal conductivity. The thermal conductivity of Ti17 alloy (8.2 W/m·K ) is somewhat higher than that of BT9 alloy (7.93 W/m·K ), resulting in the presence of finer martensite in the WM on Ti-6Al-4V/Ti17 joint.
In addition, the finer martensite in the WM of Ti-6Al-4V/Ti17 joint also demonstrated that the welding parameters would be more suitable for the Ti-6Al-4V/Ti17 joint.
As presented above, a complex phase transformation appeared across both dissimilar titanium alloy joints. The complicated phase transformation was mostly dependent on the initial microstructure of BM and the heating cycle during electron beam welding. In the WM of both dissimilar joints, electron beam welding led to a high peak temperature (2000oC ). Apparently, the temperature had exceeded the melting point of BMs (about 1660oC) during electron beam welding. As a result, both β phase and α phase transformed in turn into hightemperature β phase in a very short time. For instance, lamellar β phase first became hightemperature β phase, then lamellar α phase and equiaxed α phase changed into single β phase sequentially. Because of the high energy density of the electron beam, the high-temperature β phase melted, and solidified immediately after the electron beam source traveled away. At the moment of solidification, the fast cooling (about 410oC/s [15,16]) gave an insufficient time for the atoms to diffuse, and thus prevented the β phase from transforming into the stable α phase. As a result，acicular martensite was expected in the WM during cooling, which was attributed to the high cooling rate, as seen in Fig.5(b), (c) and Fig.6(c), (d). In the HAZ on the Ti-6Al-4V side for both dissimilar joints, the peak heating temperature during welding could reach or exceed the beta transus temperature of BMs (Ti-6Al-4V: 995oC , Ti17: 895oC , BT9:1020oC ). Therefore, solid-state transformation occurred in this zone, which was somewhat equivalent to quenching. Similarly, the high-temperature β phase transformed into martensite as shown in Fig. 5(b) and Fig.6(b) due to the fast cooling rates. However, it is worth noting that the high-temperature β phase in the HAZ on the Ti17 side did not transform into
martensite during cooling, it remained still β phase at room temperature as shown in Fig. 5(d). This was due to the fact that Ti17 contained more β stabilizers than that of Ti-6Al-4V BM and BT9 BM . In addition, in the HAZ on the BT9 side, the phase transformation was different from WM and HAZ on the Ti-6Al-4V side, since the peak temperature in this zone during heating could not reach the beta transus temperature of BT9, which was reported to be slightly higher than that of Ti-6Al-4V BM [17,19]. As a result, only lamellar β phase and α phase transformed into high-temperature β phase in turn, whereas equiaxed α could not completely transform into the β phase in the rapid heating process, thus martensite and residual α were observed in HAZ on the BT9 side during cooling as shown in Fig. 6(e) and (f). This also indicated that the HAZ on the BT9 side experienced temperatures mostly in the α + β domain. During the electron beam welding, as the distance increased from the center of WM to BM, the peak heating temperature decreased. Therefore, from outer-HAZ to BM with decreasing peak heating temperature, the extent of forming martensite or β phase also decreased and even no transformation occurred at all, as shown in Fig. 5(a) and (e) and Fig. 6(a) and (e).
Fig. 7 presents the stress-strain curves of Ti-6Al-4V BM, Ti17 BM, BT9 BM, Ti-6Al-4V/Ti17 joint, and Ti-6Al-4V/BT9 joint, obtained at a strain rate of 1×10-2 s-1. The evaluated tensile test results from those stress-strain curves are tabulated in Table 2. The BMs and both dissimilar joints exhibited smooth and continuous stress-strain curves, and the plastic deformation phase was gentle, suggesting the strain hardening ability of BMs and both joints was not so strong. The area under the stress-strain curve of every BM, which represents the toughness was larger than that of joints, indicating that BMs absorbed more energy before fracture with a relatively higher
toughness. The lower toughness of the weldment compared to the BMs could be related to the lower toughness of HAZ and WM due to the presence of martensite which is harder and more brittle than that of α and β phases as shown in Figs 5 and 6. The lower toughness of joints was also verified in the laser welded titanium alloy joint . It is seen that Ti17 BM had the highest yield strength (YS) and ultimate tensile strength (UTS), while Ti-6Al-4V BM had the lowest YS and UTS. The tensile properties of the weldment were a combination of those of WM, HAZ and BM. Owing to the presence of martensite (Fig. 5(a)-(c) and Fig.6) in the HAZ and WM, which was harder than that of α phase and β phase, it was anticipated that the weldment exhibited a higher YS and UTS compared to Ti-6Al-4V BM. As expected, the YS and UTS of both dissimilar joints lay in-between those of their BMs, suggesting that the applied welding parameters were within the optimal welding parameter range, and robust dissimilar joints were successfully achieved. Besides, the UTS of each alloy or joint were close to their YS, i.e., the yield ratio (the ratio of YS to UTS) was high, suggesting that the ability of resistance to deformation was strong or the plastic deformation was not easy for these alloys and both joints. Using the same welding parameters, YS and UTS of Ti-6Al-4V/Ti17 joint were obtained to be higher than those of Ti-6Al-4V/BT9 joint due to the presence of more martensite. However, in spite of the presence of martensite in the joints, the ductility of both dissimilar joints was still close to that of the higher strength alloy (Ti17 BM or BT9 BM).
Fatigue behavior of electron beam welded dissimilar titanium alloy joints (Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint) were investigated under stress control at R=0.1 and 50 Hz, and the obtained S-N results are plotted in Fig. 8. For the sake of comparison, the fatigue properties of
the BMs (Ti-6Al-4V BM, Ti17 BM, BT9 BM) in the as-received condition are also provided in Fig. 8. It is seen that with increasing stress amplitude the fatigue life of all BMs and dissimilar joints gradually decreased, and the fatigue life of both dissimilar joints was almost the same as that of Ti-6Al-4V BM. The obtained fatigue life in the present study was in good agreement with those presented in the literature [20-22].
Fatigue strength of Ti-6Al-4V/Ti17 joint was 150 MPa at 107 cycles, which was equivalent to Ti6Al-4V BM, however, fatigue strength of Ti-6Al-4V/BT9 joint was less than that of Ti-6Al-4V BM, which was attributed to the complex phase transformation (Figs 5 and 6) in the dissimilar joints as discussed above. At higher stress amplitudes (e.g., 350-400 MPa), Ti17 BM had a much longer fatigue life than that of Ti-6Al-4V BM or BT9 BM. For example, the fatigue life was 13199 cycles in the Ti17 BM, while it was 663 cycles in the Ti-6Al-4V BM, 1535 cycles in the BT9 BM at a stress amplitude of 350 MPa. It is well known that the mechanical properties of titanium alloy mainly depended on the microstructure and grain morphology. The equiaxed microstructure offered good ductility and superior fatigue resistance, while the lamellar microstructure gave high strength, good fracture toughness, excellent creep resistance, and fatigue crack growth resistance. In this study, globular α was observed in Ti-6Al-4V BM and BT9 BM, while finer lamellar α and β phase was seen in Ti17 BM as shown in Fig. 1. Therefore, it is understandable that the fatigue life of Ti17 BM was significantly longer than that of Ti-6Al4V BM and BT9 BM at the higher stress amplitudes due to the presence of finer lamellar α phase and β phase. However, this trend had an obvious change at lower stress amplitudes (150-300 MPa), that is, the difference in the fatigue life among Ti-6Al-4V BM, Ti17 BM, BT9 BM became small. The fatigue life of both dissimilar joints was nearly the same as that of Ti-6Al-4V
BM at the same stress amplitude within the experimental scatter, suggesting that the applied welding parameters were reasonable for both EBWed dissimilar joints under dynamic cyclic loading. This feature of fatigue resistance of both dissimilar joints was related to the complex microstructure as shown in Figs 5 and 6.
The fatigue limit and the fatigue ratio obtained according to Figs 7 and 8 are tabulated in Table 2. For titanium alloys, the S-N curve becomes horizontal at higher N values; or there is a limiting stress level, called the fatigue limit (also sometimes the endurance limit), below which fatigue failure will not occur . Therefore, the fatigue limit of Ti-6Al-4V BM, Ti17 BM, BT9 BM, and Ti-6Al-4V/Ti17 joint was about 150 MPa, however the value for Ti-6Al-4V/BT9 joint was lower than 150 MPa since one sample at this cyclic stress level failed, as shown in Fig. 8. This was related to the inhomogeneity of the microstructure as shown in the Fig.6 and the existence of the extended equiaxed α phase on the BT9 side at a distance of ~2 mm from the weld center as shown in Fig. 9, which was not observed in Ti-6Al-4V/Ti17 joint, resulting in a lower fatigue strength. The fatigue strength (or life) is known to be normally and roughly proportional to the UTS. That is, the higher the UTS is, the higher the fatigue strength (or life) is. It should be noted that the surface conditions, residual stresses, localized stress concentration, and severe weld concavity in workpieces also affect the fatigue life. The fatigue ratio (i.e., a ratio of fatigue limit to the UTS) of the Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint lay in-between their respective BMs as seen from Table 2. For example, they were 0.228, 0.133, 0.155 for Ti-6Al-4V BM, Ti17 BM, and Ti-6Al-4V/Ti17 joint, respectively.
The fatigue data shown in Fig. 8 could be evaluated using the following Basquin type equation ,
a 'f 2 N b
, (1) where σa is the stress amplitude, σf’ is the fatigue strength coefficient, N is the number of cycle to failure, and b is the fatigue strength exponent. The evaluated values of σf’ and b of BMs and the dissimilar joints are tabulated in Table 2. Then the fatigue life at a given stress amplitude could be estimated on the basis of both fatigue strength coefficient σf’ and fatigue strength exponent b.
Fractography of fatigued samples
The failure location of both Ti-6Al-4V/Ti17 and Ti-6Al-4V/BT9 dissimilar joints basically lay in the HAZ. It was less likely to occur at the center of weld due to the presence of martensite (Figs 5 and 6) due to the higher strength there. Fig. 10 shows an overall view of fatigue fracture surface of Ti-6Al-4V/BT9 joint tested at a stress amplitude of 200 MPa. Three typical fatigue stages could be identified on the fracture surface of tested fatigue sample (Fig. 10(a)), containing fatigue crack initiation site, crack propagation region, and final fast fracture region. It is seen from the low magnification image that the fatigue crack initiated basically from the specimen surface, and the river-line patterns appeared on the dissimilar joint which flowed along the crack propagation direction (Fig.10(b)). After initiation, cracks started to propagate towards the center of the sample. The crack propagation region exhibited a flat and striated type fracture surface (Fig. 10c). The occurrence of fatigue striations was attributed to a repeated plastic bluntingresharpening process as a result of slip of dislocations in the plastic zone ahead of the fatigue crack tip . The final rapid fracture region appeared rougher than the crack propagation region, which was characterized by typical dimples (Fig.10(d)). There was an obvious boundary between
the fast fracture region and the crack propagation region (Fig. 10), which indicated that the fatigue fracture and final tensile-like fracture were governed by different fracture mechanisms.
A study on the microstructure, tensile and fatigue behavior of electron beam welded dissimilar titanium alloy joints (Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint) was performed. The following conclusions could be drawn. 1) Two robust and defect-free dissimilar titanium alloy joints with relatively narrow WM and HAZ were achieved via electron beam welding. 2) Despite the same welding parameters applied, the dissimilar joints exhibited significantly different microstructures. Martensite in the WM of Ti-6Al-4V/Ti17 joint was finer than that of Ti-6Al-4V/BT9 joint; coarse β phase, and martensite and retained α phase were present in the HAZ of the Ti-6Al-4V/Ti17 joint and Ti-6Al-4V/BT9 joint, respectively. However, the martensite in the HAZ on both dissimilar joints was almost same. 3) The yield strength and ultimate tensile strength of both dissimilar joints lay in-between those of BMs. While the elongation of dissimilar joints was lower than that of Ti-6Al-4V BM, it appeared equivalent to that of BT9 BM or Ti17 BM. The strength of Ti-6Al-4V/Ti17 joint was a little higher than that of Ti-6Al-4V/BT9 joint due to the presence of finer martensite in the WM of the Ti-6Al-4V/Ti17 joint. 4) In spite of the microstructural change, fatigue life of both dissimilar joints was equivalent to that of Ti-6Al-4V BM. Fatigue failure of both Ti-6Al-4V/Ti17 and Ti-6Al-4V/BT9 joints basically occurred in the HAZ and fatigue crack initiated basically from the specimen surface.
These findings suggested that the welding parameters selected to weld the dissimilar joints were reasonable.
Acknowledgements This work was supported by the National Natural Science Foundation of China (NSFC, No: 51505379), the State Key Laboratory of Solidification Processing (No: SKLSP201505) of Northwestern Polytechnical University (NWPU), and Shaanxi Provincial Department of Education (16JK1615). One of the authors (D.L. Chen) is also grateful for the financial support by the Natural Sciences and Engineering Research Council of Canada (NSERC), Premier’s Research Excellence Award (PREA), NSERC-Discovery Accelerator Supplement (DAS) Award, Canada Foundation for Innovation (CFI), and Ryerson Research Chair (RRC) program. The authors would like to thank Q. Li, A. Machin, and R. Churaman for easy access to the laboratory facilities of Ryerson University and their assistance in the experiments. Special thanks are due to the
(2015QNKYCXTD02) for helpful discussion.
J.H. Tao, S.B. Hu, L.B. Ji, Effect of trace solute hydrogen on the fatigue life of electron beam welded Ti-6Al-4V alloy joints, Mater. Sci. Eng. A 684 (2017) 542-551.
A. Chamanfar, T. Pasang, A. Ventura, W.Z. Misiolek, Mechanical Properties and Microstructure of Laser Welded Ti-6Al-2Sn-4Zr-2Mo (Ti6242) Titanium Alloy, Mater. Sci. Eng. A 663 (2016) 213-224.
X.Y. Wang, W.Y. Li, T.J. Ma, A. Vairis, Characterisation studies of linear friction welded titanium joints, Mater. Des. 116 (2017) 115-126.
H.T. Zhang, P. He, J.C. Feng, H.Q. Wu, Interfacial microstructure and strength of the dissimilar joint Ti3Al/TC4 welded by the electron beam process, Mater. Sci. Eng. A 425 (2006) 255-259.
L.J. Tan, Z.K. Yao, W. Zhou, H.Z. Guo, Y. Zhao, Microstructure and properties of electron beam welded joint of Ti-22Al-25Nb/TC11, Aerosp. Sci. Technol. 14 (2010) 302306.
R. Pederson, F. Niklasson, F. Skystedt, R. Warren, Microstructure and mechanical properties of friction- and electron-beam welded Ti-6Al-4V and Ti-6Al-2Sn-4Zr-6Mo, Mater. Sci. Eng. A 552 (2012) 555- 565.
S.Q. Wang, J.H. Liu, D.L. Chen, Effect of strain rate and temperature on strain hardening behavior of a dissimilar joint between Ti-6Al-4V and Ti17 alloys, Mater. Des. 56 (2014) 174-184.
S.Q. Wang, J.H. Liu, Z.X. Lu, D.L. Chen, Cyclic deformation of dissimilar welded joints between Ti-6Al-4V and Ti17 alloys: Effect of strain ratio, Mater. Sci. Eng. A 598 (2014) 122-134.
S.Q. Wang, J.H. Liu, D.L. Chen, Strain-controlled fatigue properties of dissimilar welded joints between Ti-6Al-4V and Ti17 alloys, Mater. Des. 49 (2013) 716-727.
S.Q. Wang, J.H. Liu, D.L. Chen, Tensile and fatigue properties of electron beam welded dissimilar joints between Ti-6Al-4V and BT9 titanium alloys, Mater. Sci. Eng. A 584 (2013) 47-56.
S.Q. Wang, W.Y. Li, Y. Zhou, X. Li, D.L. Chen, Tensile and fatigue behavior of electron beam welded dissimilar joints of Ti-6Al-4V and IMI834 titanium alloys, Mater. Sci. Eng. A 649 (2016) 146-152.
K.P. Rao, K. Angamuthu, P.B. Srinivasan, Fracture toughness of electron beam welded Ti6Al4V, J. Mater. Process. Technol. 199 (2008) 185-192.
L.L. Su, Study on high speed milling of titanium alloy TC17, 2. Study on high speed milling force of TC17 alloy, Nanjing University of Aeronautics and Astronautics. 2012, 11
R. Rai, P. Burgardt, J.O. Milewski, T.J. Lienert, T.D. Roy, Heat transfer and fluid flow during electron beam welding of 21Cr-6Ni-9Mn steel and Ti-6Al-4V alloy, J. Phys. D: Appl. Phys. 42 (2009) 1-12.
M.T. Jovanović, S. Tadić, S. Zec, Z. Mišković, I. Bobić, The effect of annealing temperatures and cooling rates on microstructure and mechanical properties of investment cast Ti-6Al-4V alloy, Mater. Des. 27 (2006) 192-199.
T. Ahmed, H.J. Rack, Phase transformations during cooling in [alpha]+[beta] titanium alloys, Mater. Sci. Eng. A 243 (1998) 206-211.
S.L. Semiatin, T.R.Bieler, The effect of alpha platelet thickness on plastic flow during hot working of Ti-6Al-4V with a transformed microstructure, Acta Mater. 49 (2001) 35653573.
T. Wang, H.Z. Guo, L.J Tan, Z.K. Yao, Y. Zhao, P.H. Liu, Beta grain growth behavior of TG6 and Ti17 titanium alloys, Mater. Sci. Eng. A 528 (2011) 6375-6380.
H.W. Song, S.H. Zhang, M. Cheng, Subtransus deformation mechanisms of TC11 titanium alloy with lamellar structure. Trans. Nonferrous Met. Soc. China 20 (2010) 2168-2173.
S.Z. Xing, P.S. Dong, Fatigue of titanium weldments: S-N testing and analysis for data transferability among different joint types, Mar. Struct. 53 (2017) 1-19.
N. Hrabe, T.G. Herold, T. Quinn, Fatigue properties of a titanium alloy (Ti–6Al–4V) fabricated via electron beam melting (EBM): Effects of internal defects and residual stress, Inter. J. Fatigue 94 (2017) 202-210.
C.W. Huang, Y.Q. Zhao, S.W. Xin, C.S. Tan, W. Zhou, Q. Li, W.D. Zeng, High cycle fatigue behavior of Ti-5Al-5Mo-5V-3Cr-1Zr titanium alloy with lamellar microstructure, Mater. Sci. Eng. A 682 (2017) 107-116.
W.D. Callister, D.G. Rethwisch, Materials Science and Engineering: An Introduction, 9th edition, John Wiley & Sons, Inc. 2014, pp.272-276.
W. Xu, D. Westerbaan, S.S. Nayak, D.L. Chen, F. Goodwin, E. Biro, Y. Zhou, Microstructure and fatigue performance of single and multiple linear fiber laser welded DP980 dual-phase steel, Mater. Sci. Eng. A 553 (2012) 51-58.
C. Laird, Fatigue Crack Propagation, ASTM STP 415, ASTM International, West Conshohocken: PA; 1967.
Table 1 Details of EBW parameters used for joining dissimilar titanium alloys in the present study. Table 2 Tensile properties and fatigue parameters σf' and b for the BM and joints.
Fig. 1 Microstructures of BM of (a) Ti-6Al-4V alloy, (b) Ti17 alloy, and (c) BT9 alloy. Fig. 2 Surface profile of Ti-6Al-4V/Ti17 joint in the (a) front and (b) back side. Fig. 3 Surface profile of Ti-6Al-4V/BT9 joint in the (a) front and (b) back side. Fig. 4 Overall view of the cross section of electron beam welded (a) Ti-6Al-4V/Ti17 joint, and (b) Ti-6Al-4V/BT9 joint. Fig. 5 Magnified images of the dashed box in Fig.4(a) for (a) outer-HAZ of Ti-6Al-4V, (b) inner-HAZ of Ti-6Al-4V and WM, (c) magnified area of WM, (d) inner-HAZ of Ti17, and (e) outer-HAZ of Ti17. Fig. 6 Magnified images of the dashed box in Fig.4(b) for (a) outer-HAZ of Ti-6Al-4V, (b) innerHAZ of Ti-6Al-4V, (c) and (d) WM, and (e) and (f) HAZ of BT9.
Fig. 7 Typical stress-strain curves of BM and dissimilar joints tested at a strain rate of 1×10−2 s−1 for (a) Ti-6Al-4V BM, Ti17 BM and their joint, and (b) Ti-6Al-4V BM, BT9 BM, and their joint. Fig. 8 S-N curves of BM and dissimilar joints tested at R=0.1, 50 Hz, and RT. Fig. 9 The microstructure at a distance of ~2 mm from the weld center on the BT9 side. Fig. 10 Typical SEM images of fatigue fracture surface of Ti-6Al-4V/BT9 dissimilar joint tested at a stress amplitude of 200 MPa, (a) overall view, (b) crack initiation area, (c) crack propagation area, and (d) the fast fracture region.
Table 1 Details of EBW parameters used for joining dissimilar titanium alloys in the present study. Welding parameters Joint
Accelerating voltage Beam current Focus current welding speed UB, kV
Table 2 Tensile properties and fatigue parameters σf' and b for the BM and joints. Materials
4V/Ti17 Ti-6Al-4V/ 852
(c) α β
Fig. 1 Microstructures of BM of (a) Ti-6Al-4V alloy, (b) Ti17 alloy, and (c) BT9 alloy.
Fig.2 Surface profile of Ti-6Al-4V/Ti17 joint in the (a) front and (b) back side. (a)
Welding direction 2mm
Fig.3 Surface profile of Ti-6Al-4V/BT9 joint in the (a) front and (b) back side.
Fig. 4 Overall view of the cross section of electron beam welded (a) Ti-6Al-4V/Ti17 joint, and (b) Ti-6Al-4V/BT9 joint. (b)
Fig.5 Magnified images of the dashed box in Fig.4(a) for (a) outer-HAZ of Ti-6Al-4V, (b) innerHAZ of Ti-6Al-4V and WM, (c) magnified area of WM, (d) inner-HAZ of Ti17, and (e) outerHAZ of Ti17.
(f) Retained α
Fig.6 Magnified images of the dashed box in Fig.4(b) for (a) outer-HAZ of Ti-6Al-4V, (b) innerHAZ of Ti-6Al-4V, (c) and (d) WM, and (e) and (f) HAZ of BT9.
Engineering stress, MPa
1000 800 600 Ti-6Al-4V BM
Ti-6Al-4V/Ti17 joint Ti17 BM
200 0 0
2 3 4 5 6 7 8 Engineering strain, %
Engineering stress, MPa
Ti-6Al-4V BM BT9 BM
3 4 5 6 7 8 Engineering strain, %
Fig. 7 Typical stress-strain curves of BM and dissimilar joints tested at a strain rate of 1×10−2 s−1 for (a) Ti-6Al-4V BM, Ti17 BM and their joint, and (b) Ti-6Al-4V BM, BT9 BM, and their joint.
Stress amplitude, MPa
600 Ti-6A-l4V BM Ti17 BM BT9 BM Ti-6Al-4V/Ti17 joint Ti-6Al-4V/BT9 joint
500 400 300 200 100
0 1E+11E+21E+31E+41E+51E+61E+71E+8 Number of cycles to failure, Nf
Fig. 8 S-N curves of BM and dissimilar joints tested at R=0.1, 50 Hz, and RT.
Fig. 9 The microstructure at a distance of ~2 mm from the weld center on the BT9 side.
Fig. 10 Typical SEM images of fatigue fracture surface of Ti-6Al-4V/BT9 dissimilar joint tested at
a stress amplitude of 200 MPa, (a) overall view, (b) crack initiation area, (c) crack propagation area, and (d) the fast fracture region.