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Effect of beam oscillation on electron beam welding of copper with AISI-304 stainless steel
Journal of Materials Processing Technology 233 (2016) 174–185
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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Effect of beam oscillation on electron beam welding of copper with AISI-304 stainless steel Jyotirmaya Kar ∗ , Sanat Kumar Roy, Gour Gopal Roy Department of Metallurgical and Materials Engineering, Indian Institute of Technology, Kharagpur 721302, India
a r t i c l e
i n f o
Article history: Received 10 November 2015 Received in revised form 1 March 2016 Accepted 1 March 2016 Available online 3 March 2016 Keywords: Dissimilar welding Electron beam welding Beam oscillation Stainless steel Copper
a b s t r a c t The quality of dissimilar weld joint between copper and stainless steel (SS) plates using electron beam welding with and without beam oscillation was evaluated through microstructural analysis, hardness, tensile, bend and impact tests. Microstructural analysis as well as quantitative elemental analyses demonstrated that improved mixing of copper with SS in the fusion zone could be achieved by an oscillating beam of optimum oscillation diameter. Although the tensile properties at room temperature did not vary much between the welds; high temperature measurements exhibited significant difference between them. Welds made with beam oscillation having an optimum oscillation diameter demonstrated significant increase in impact strength as well as percentage elongation compared to joints made without beam oscillation. Increased beam oscillation beyond optimum oscillation diameter was found to be detrimental promoting segregation of copper in the weld zone exhibiting deteriorated mechanical properties. A hypothesis has been put forward for better mixing during welding with beam oscillation. © 2016 Elsevier B.V. All rights reserved.
1. Introduction For the last few decades, dissimilar-metal joints are widely being used in various industries with an aim to reduce weight and cost savings due to less use of one of the components. In addition, they are capable in offering more complex functions in various engineering applications. Copper-SS dissimilar joint is one of such case, which is designed to utilize the excellent heat and electrical conductivity imparted by copper whereas, better strength, wear and corrosion resistance to be imparted by steel. Accordingly, copperSS joints find their applications in the field of power generation and transmission industries, heat transfer components, cryogenic sector, electrical and electronics sector etc. However, fusion welding of copper-SS is more challenging because of their differences in the physical, chemical and thermomechanical properties. For instance, thermal conductivity of copper is much higher than that of steel which causes difficulties in reaching its melting temperature during fusion welding. Besides, Cu and Fe have limited solid solubility among themselves (Massalski et al. (1986)) and likely to undergo liquid separation under comparatively higher cooling rate. The two materials also have large differences in their melting temperature and coefficient of thermal
∗ Corresponding author. E-mail address: [email protected] (J. Kar). http://dx.doi.org/10.1016/j.jmatprotec.2016.03.001 0924-0136/© 2016 Elsevier B.V. All rights reserved.
expansion. Fusion welding of copper-SS dissimilar couples using low energy density conventional arc welding processes like GTAW, SMAW and GMAW etc. were also found to be difficult. In such welding processes selection of proper filler metal plays a crucial role, as reported by Sajjad et al. (2012), Velu and Bhat (2013) and Roy et al. (2014). Owing to various advantages like contamination free weld, high energy density, precise positioning of the heat source, narrow heat affected zone (HAZ) etc., electron beam welding (EBW) is considered as one of the most sophisticated fusion welding technique that could be used in several component manufacturing processes where conventional arc welding fails (Cottrell (1985)). Sun and Karppi (1996) presented an overview on the applications of EBW for the joining of dissimilar metals and demonstrated the advantages of EBW over conventional fusion welding processes. Converse to EBW process, laser beam welding has certain limitation such as to cope with the high reflectivity of copper. Mai and Spowage (2004) in their study focused a laser beam 0.2 mm into the steel surface and thus obtained very limited amount (<2%) of copper in the fusion zone (FZ). Consequently, a complete metallurgical bond at the interface between copper plate and molten steel could not be achieved. Yao et al. (2009) produced a butt joint with an inclined joint interface where by off-setting the laser beam on the steel side the dilution between copper and steel in the FZ was controlled. Phanikumar et al. (2005) studied Fe-Cu dissimilar couple and reported heterogeneous microstructure in the FZ and such het-
J. Kar et al. / Journal of Materials Processing Technology 233 (2016) 174–185
erogeneity was found to decrease with lower welding speed. Chen et al. (2013) produced copper-SS dissimilar joint by laser welding and inferred that melting of copper in small amount might be beneficial for back-filling of the microcracks in the FZ; however, its larger amount might cause more inhomogeneity both in composition as well as in stress level leading to severe microcrack generation. Saha et al. (1999) investigated copper-SS butt joint using EBW and reported that better mechanical properties of the weld could be obtained by off-setting beam towards the copper side (0.5–1 mm) than running the beam along the joint interface. This resulted in melting of copper in the FZ and back-filling of the microcracks by copper. They also reported that higher welding speed (1000 mm/min and above) is beneficial in reducing the microcrack formation in the FZ. Tosto et al. (2003) successfully prepared defect free thicker (25 mm) dissimilar joints of copper-SS using EBW without beam off-setting. However, they used comparatively higher beam current (395 mA) and lower beam voltage (15 kV) than reported by earlier investigators. Magnabosco et al. (2006) reported joining of copper with SS having higher plate thickness (30–70 mm) using both single and double pass method under different set of operating parameters. They reported more heterogeneity and defects in the microstructure as the plate thicknesses were increased and sought determination of proper optimized process parameters. Unlike the use of a static beam, beam oscillation is supposed to produce a churning action of the liquid in the weld-pool, which in effect would improve the mixing and restrict the segregation providing a more uniform dispersion of copper in the weldment region. This is in contrast to the weld-pool created without using beam oscillation, where directed liquid flow due to Marangoni convection brings in directional properties. It may be pointed out that beam oscillation as a process parameter in EBW for achieving better mixing in the FZ of dissimilar metal joints, particularly for copper-SS has not yet been studied. Accordingly, the present study focuses on the microstructural evolution and its correlation to the mechanical properties of electron beam welded joints of copper (C10300) and AISI-304 SS using both oscillating and non-oscillating beam. The microstructures of the FZ, FZ-base plate interfaces and HAZ, and mechanical properties like micro-hardness, tensile strength (both room and elevated temperature), three-point bending and impact strength were investigated in details. Three specimens for each of the experimental conditions were tested to achieve reproducibility. 2. Experimental procedure 2.1. Materials and design of joints Copper (C10300) and AISI-304 SS plates with dimensions 200 × 80 × 3 mm3 (length × breadth × height) were prepared from the as received plates. The spectroscopic analyses (X-ray fluorescence spectroscopy) of the base plates, namely copper and SS are presented in Table 1. Subsequently, the joint-side face of each plate (transverse to rolling direction) were machined and polished with 240-grit emery paper. Prior to welding, the plates were thoroughly degreased and cleaned with acetone. The two abutting plates were tight fitted using clamps to ensure negligible gap between the plates. Subsequently, the plates were electron beam butt welded using the process parameters as stated in Table 2. 2.2. Process parameters used in EBW Welding of the dissimilar materials were carried out using an 80kV–12 kW EBW machine at IIT Kharagpur, indigenously designed and fabricated by the Bhabha Atomic Research Center (BARC), Department of Atomic Energy (DAE), India.
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At the beginning, EBW experiments were carried out by autogenous bead-on-plate welding of 3 mm thick 304-SS and copper plates with varying beam current, beam voltage and welding speed. Such an exercise was necessary to understand their effects on weld bead profile and to optimize the process parameters in achieving full penetration joints both for oscillating and non-oscillating beam. Three different sets of welding parameters were used to prepare the joints. Out of the three joints, two were prepared using oscillating beam with two different oscillation diameter (1 and 2 mm) and the other was prepared with a non-oscillating beam. During welding, the beam was always focused on the base plate surface and aligned to the butt line, maintaining the beam voltage and welding speed constant. It may be noted that while changing the beam oscillation parameter, current could not be maintained constant and it varied between 12–23%. This is attributed to the fact that during butt welding of copper with SS using beam oscillation, the residence time of the beam increases on both sides of the joint. Since copper conducts heat faster, prolonged dwelling of the beam on copper side favours dissipation of more heat by conduction. This situation demands more power input to compensate for the heat loss in the production of defect free welds. Since beam voltage remained constant during welding, this power compensation was achieved through increase in the beam current. It may be noted that such threshold current requirement for a particular oscillation diameter was established through several trial experiments. Apart from the above mentioned process parameters, other important process variables were maintained constant as listed in Table 3. 2.3. Characterization of the joints After welding, each prepared joint was inspected visually as well as examined by X-ray radiography for its soundness. Only after ensuring its soundness, further study on evolution of microstructures and evaluation of mechanical properties were undertaken. For microscopic observation specimens were cut perpendicular to the weld bead using a wire-cut electric discharge machining (EDM). Subsequently, each specimen was polished using a series of SiC emery papers from 240 to 4000 grit size in succession with water as the coolant, followed by fine polishing using 0.05 m alumina paste in a velvet cloth disc. For metallographic sample preparation standard ASTM E3-11 (2011) method was followed. To reveal the microstructure, aquaregia (75% HCl + 25% HNO3 ) and solution of 5gm FeCl3 + 50 ml HCl + 100 ml water were used as the etchants for SS and copper side, respectively. The etched samples were studied under both optical microscope (Laica® : DM6000 M) and scanning electron microscope (SEM) (Zeiss® : EVO-60). Distribution of chemical elements in the FZ was analyzed through elemental line scan as well as area mapping by using energy dispersive X-ray spectroscopy (EDS). For room temperature tensile tests, transverse tensile specimens (length perpendicular to welding direction) were prepared using EDM and tested according to standard ASTM E8/E8m-15a (2015). Likewise, for elevated temperature tensile tests at 400 ◦ C, transverse tensile specimens were prepared and tested as per standard ASTM E21-09 (2009). Fig. 1 (a) and (b) show the sample design and dimension for tensile tests at room and elevated temperatures, respectively. For both types of tensile tests, cross-head speed (displacement rate) of 0.5 mm/min was chosen. The room and the elevated temperature tensile tests were performed in a 100 kN (Instron® : 8862) and 250 kN (Instron® : 8800) tensile testing machine, respectively. Transverse face and root three-point bend tests were performed as per standard ASTM E190-14 (2014), to investigate the ductility and soundness of the weld joints. Standard specimens for this test were prepared, having 150 mm length and 30 mm breadth. A
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Table 1 Chemical composition of AISI-304 SS and copper (C10300) in weight%. AISI-304SS
Cr 18.06 Cu 99.97
Copper C10300
Ni 8.06 Pb 0.0001
Mn 1.802 P 0.003
Mo 0.15 Fe 0.001
Si 0.36 O 0.003
C 0.043 Ni 0.003
Al P 0.06 0.042 Total impurities Total impurities 0.03% Max.
Fe Rest
Table 2 EBW process parameters used in this study. Joint No
Beam Voltage (kV)
Beam Current (mA)
Welding Speed (mm/min)
Oscillation Frequency (Hz)
Oscillation diameter (mm)
1 2 3
60 60 60
65 73 80
1000 1000 1000
Not used 600 600
Not used 1.0 2.0
Table 3 Other important welding process variables. Gun chamber vacuum (mbar)
<5 × 10−6
Gun-specimen distance (mm)
460
Welding chamber vacuum (mbar) Beam focus point (mm)
<5 × 10−5 0
Welding pass (#) Beam off-set
1 none
Fig. 3. Sample design and dimension for Charpy impact test.
Fig. 1. Sample design and dimension for the tensile tests at: (a) room temperature, and (b) 400 ◦ C.
as per Fig. 3 and tested following the standard ASTM E23-12c (2013) method at ambient temperature of 25 ◦ C in an impact machine (Instron® : SI-1C3). Micro-hardness values were also measured following the standard ASTM E384-111 (2011) across the weld metal at an interval of 0.2 mm using a Vickers micro-hardness testing machine (UHL® : VMHT-001) under a load of 100gf/15s (load/dwell time). To identify the effect of microstructural heterogeneity such measurements were performed for both the parent metal, HAZ and FZ regions. 3. Results and discussion 3.1. Study of microstructure and elemental analysis
Fig. 2. Schematic diagram illustrating the method of three-point face/root bend test.
support plate of 304-SS having thickness of 2.5 mm was used to compensate the difference in strength between copper and SS. Furthermore, the upper plunger was shifted towards the SS side by 10 mm from the weldment region to prevent horizontal shifting of the plunger during bending towards the copper side. The test setup is shown in Fig. 2 and the three-point bend tests were performed in a 50 kN universal testing machine (Shimadzu® : AG-5000G). To evaluate the impact strength of the joints, sub-size Charpy impact specimens (due to smaller plate thickness) were prepared
Fig. 4 (a) shows an overview in the micrograph for joint-1 at a lower magnification while Fig. 4 (b)–(d) show magnified images of the different regions for the same joint obtained using nonoscillating beam. Fig. 4 (b) clearly demonstrates that the major melting had occurred on the SS side, while limited melting took place on the copper side. Moreover, copper-FZ interface seems to be rough with significant presence of iron [Fig. 4 (d)]. Fig. 4 (c) shows no traces of copper in the FZ rather comprises of re-melted and resolidified SS only. The cellular dendritic structure is well evident in this zone. The embedded figure in Fig. 4 (b) at a higher magnification shows that the FZ-SS interface is smooth and comprised of cellular dendritic morphology. The corresponding microstructures of joint-2, which was prepared with an oscillating beam of oscillation diameter 1 mm, are shown in Fig. 5 (a)–(c). Here, a significant amount of copper could be found in the FZ as shown in Fig. 5 (b). It may be noted that during beam oscillation, the beam dwells on both sides of the joint including the copper region for a certain amount of time, which not only promotes heat dissipation but also melts copper locally. It was also observed that under oscillating beam both the copper-FZ interface
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Fig. 4. Optical microscopic images for joint-1: (a) overview of the whole joint, (b) SS-FZ interface, (c) middle portion of the FZ, and (d) FZ-copper interface.
Fig. 5. Optical microscopic images for joint-2: (a) SS-FZ interface, (b) middle portion of the FZ, and (c) FZ-copper interface.
[Fig. 5 (c)] and SS-FZ interface [embedded in Fig. 5 (a)] are comparatively smoother than in the case of non-oscillating beam (Fig. 4). This might be attributed to uniform heat dissipation in all directions by the churning action of the liquid melt. Fig. 5 (b) further reveals that copper is uniformly dispersed in the FZ. Similarly, Fig. 6 (a)–(c) depict the microstructures of different regions of joint-3 (obtained with an oscillating beam of oscillation diameter 2 mm). It demonstrates that a large amount of copper is present both in the middle of the FZ as well as adjacent to copper side [Fig. 6 (b) and (c)]. It is evident from the microstructures that more copper got melted during the longer dwelling of the beam through the copper region during its oscillation. However, the molten copper in the weld metal on its solidification appears in the form of chunks and segregated mass [Fig. 6 (b)]. This may be attributed to the fact that churning action generated due to beam oscillation is not sufficient to disperse the large amount of molten
copper uniformly all over the weld metal. The SS-FZ interface is also found to be more rough compared to the joint prepared with lower beam oscillation diameter. This might be attributed to a less degree of mixing under higher beam oscillation diameter leading to directed heat flow. All joints regardless of being prepared either using a nonoscillating beam or oscillating beam were found to possess only a narrow HAZ region on the copper side adjacent to FZ as shown in Fig. 7. The HAZ region is distinguished by coarse equiaxed polyhedral grains with annealing twins compared to finer grains in the copper base metal. Copper having higher thermal conductivity allows large amount of heat to flow through it and causes grain growth in the HAZ. Furthermore, optical microscopic analysis of the joint cross-sections did not reveal any gross defects like microcracks, porosity, etc.
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Fig. 6. Optical microscopic images for joint-3: (a) SS-FZ interface, (b) middle portion of the FZ, and (c) FZ-copper interface.
Fig. 7. Optical microscopic image depicting the HAZ region on the copper side for joint-1 in etched condition.
Fig. 8. EDS profiles for joint-1: (a) SEM (BSE) image of Joint-1, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
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179
Fig. 9. EDS profiles for joint-2: (a) SEM (BSE) image of Joint-2, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
Fig. 10. EDS profiles for joint-3: (a) SEM (BSE) image of Joint-3, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
Figs. 8–10 show EDS elemental line scan and area mapping for Cu and Fe across the FZ for joint-1, 2 and 3, respectively. Fig. 8 (a)–(e) depict the BSE SEM image, line scan and area mapping for copper and iron in joint-1, prepared with non-oscillating beam. The line corresponding to the EDS scan path and the rectangle corresponding to EDS area mapping are shown in Fig. 8 (a). It may be noted
that the line scan covers the whole weld region from the Cu-weld interface to SS-weld interface. Both the line scan [Fig. 8 (b) and (c)] as well as area mapping [Fig. 8 (d) and (e)] confirm that a narrow copper rich region exists near the copper side of FZ with a negligible amount of copper dispersed across the rest of the FZ, while the majority of it is rich in iron. Fig. 9 that represents similar plots
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Fig. 11. FE-SEM (BSE) images showing microcracks: (a) joint-1, (b) enlarged view of the insert for joint-1, (c) for joint-2, (d) corresponding EDS point spectrum bitmaps, (e) joint-3, and (f) enlarged view of the insert for joint-3.
for joint-2 prepared with an oscillating beam of oscillation diameter 1 mm. This figure [Fig. 9 (b)–(e)] clearly exhibits that though the copper rich region still exists near the copper-FZ interface, but significant amount of it also appears in the FZ. Fig. 9 (b) further depicts that copper is more uniformly dispersed in the FZ. The line scan and area mapping as shown in Fig. 10 (b)–(e) shows that for joint-3, obtained with beam oscillation at higher oscillation diameter, the amount of copper in the FZ is more and segregated as, corroborated by the optical microstructure (Fig. 6). Fig. 11 (a)–(f) shows SEM image of joint-1, 2 and 3, respectively taken at different locations near the copper rich mass in the FZ. It is revealed that all the three joints have microcracks in the FZ. The formation of such microcracks might be attributed to inhomogeneity in composition and complex stress distribution in the FZ during the
Fig. 12. Micro-hardness profile across the SS-copper bimetallic joints.
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Fig. 13. Stress-strain curves for: (a) copper-SS joints, and (b) copper base plate.
solidification of the weld-pool. It is apparent from Fig. 11 (b) and (f) that the microcracks are of open type and these are not backfilled by copper, for joint-1 and 3, respectively. In contrast, Fig. 11 (c) representing the SEM image of joint-2, depicts the presence of tiny copper globules on the line of microcracks. EDS analysis of a spot on the microcrack [spectrum 1 in Fig. 11 (d)] also suggests the presence of large amount of copper within the microcrack. Thus, in case of joint-2, the microcracks are readily back-filled by copper. When copper gets mixed well in the form of tiny droplets across the weld, it increases the probability of finding copper droplets on microcracks as depicted in Fig. 11 (c). Since copper remains liquid while iron gets solidified, copper wets the microcracks generated on SS during its solidification and fills it. But if the mixing is not sufficient liquid copper cannot move further from its source. Accordingly, it appears in the form of segregated mass during solidification and finds limited scope to backfill the microcracks situated far away. Besides, segregation of copper introduces more inhomogeneity and thermal stress that generates more microcracks. Possibly for such reasons, there are large open cracks in joint -3. However, in the case of joint-1, since copper does not melt appreciably and mixing is insufficient cracks remain open.
Table 4 Room temperature tensile properties of the bimetallic welded joints and copper base plate. Material
0.2% YS (MPa)
Joint-1 Joint-2 Joint-3 Copper plate
168 171 170 205
± ± ± ±
3 4 3 4
UTS (MPa) 249 252 250 261
± ± ± ±
2 3 2 2
Elongation (%) 17 19 18 36
± ± ± ±
2 2 2 2
Failure Location copper side copper side copper side
3.2. Micro-hardness The variation in micro-hardness across the weld cross-section for each of the three prepared joints is shown in Fig. 12. It is observed that all the prepared joints possess higher hardness in the FZ than the copper base plate. Furthermore, there exists a narrow HAZ region on the copper side. A deep in hardness adjacent to the FZ/HAZ interface on the copper side might be attributed to its coarse grained structure as revealed in the microstructural analysis (Fig. 7). It may be noted that fluctuation in micro-hardness values are maximum in the FZ of joint-3; while it is minimum for the joint-1 and moderate for joint-2. Large fluctuations in hardness values in joint-3 might be attributed to greater amount of copper segregation in the FZ. In joint-2, since only moderate amount of copper is finely distributed, fluctuation becomes moderate. For joint-1 there is hardly any variation in hardness along the FZ since very little copper got melted during the joining process. Therefore, hardness variation also corroborates the microstructural investigation regarding the extent of copper mixing in the FZ. 3.3. Room temperature tensile properties The stress-strain plots for the three joints and that for the copper base plate, tested at room temperature, are shown in Fig. 13 (a)
Fig. 14. Photographs of fractured room temperature tensile samples.
and (b), respectively. Results pertaining to the tensile tests such as 0.2% off-set yield strength (YS), ultimate tensile strength (UTS) and percentage elongation up to rupture etc. were evaluated and are presented in Table 4. The relevant values for copper base plate (as it is the weakest) are also included in Table 4. Table 4 and Fig. 13 (a) clearly depict that all the joints possess almost same room temperature tensile properties. The UTS for all the joints were found to be more than 95% of copper signifying that the joints achieved adequate strength. However, the percentage elongation for all the three joints is almost half than that of copper. Fig. 14 shows the fractured test samples where the sample identification number corresponds to joint number. It is observed that for all the joints, the SS-FZ interface remained intact while fracture had occurred on the copper side, suggesting that the characteristic of the base metal remains uninfluenced by the presence or absence of beam oscillation. The tensile test results closely match the data reported by Tosto et al. (2003), Zhang et al. (2014) and Chen et al. (2015). In arc welding processes (SMAW); Velu and Bhat (2013) and Roy et al. (2014) reported that failure occurred in the HAZ region on the copper side.
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Fig. 15. Stress-strain curves at 400 ◦ C for: (a) copper-SS joints, and (b) copper base plate.
Table 5 Tensile properties of the welded joints and copper base plate at 400 ◦ C. Material
0.2% YS (MPa)
Joint-1 Joint-2 Joint-3 Copper plate
149 149 146 169
± ± ± ±
1 1 3 2
UTS (MPa) 179 183 180 189
± ± ± ±
2 1 2 2
Elongation (%) 7 14 8 48
± ± ± ±
1 2 1 2
Failure Location FZ-SS interface HAZ/ copper HAZ/ copper copper
3.4. Elevated temperature tensile properties To determine the high temperature serviceability of the joints, tensile tests were performed at 400 ◦ C and the corresponding results are presented in Table 5. For comparison, the properties of copper base plate (at 400 ◦ C) were also evaluated and are shown in the same Table 5. The corresponding stress-strain plots for the three joints as well as copper base plate are shown in Fig. 15 (a) and (b), respectively. It is clearly seen that even though all three joints demonstrated almost the same YS and UTS values, but the percentage elongation is almost twice in case of joint-2 compared to joint-1 and 3. Therefore, the area under the stress-strain curve, which represents the toughness, becomes double for joint-2. This indicates that joint2 with adequate mixing of copper in the FZ possesses better high temperature toughness rendering better serviceability. Similar to room temperature test results, here also the UTS for all the joints were 95% more than that of copper, but the percentage elongation was lower. The fractured tensile samples are shown in Fig. 16 where the sample identification number corresponds to joint number. It is observed that joint-1 which was prepared using a non-oscillating beam, fractured in the FZ-SS interface indicating improper bonding at the interface. However, joint-2 and 3 prepared using oscillating beam, fractured away from the FZ that is on copper side, while the SS-FZ interface remained intact. Incidentally, the fracture occurred at the interface of the two metallic material in joint-1, while for the other two joints it occurred away from the FZ. It can be inferred that joints prepared using oscillating beam do possess satisfactory tensile properties at elevated temperature. Since the tensile samples tested at 400 ◦ C, fractured at different locations, fracture surface morphologies were expected to be different. The fractographs of the failed samples are presented in Fig. 17 (a)–(c). Fig. 17 (a) corresponds to the fracture surface of joint1 showing some cracks. However, fractographs of joint-2 and 3 as shown in Fig. 17 (b) and (c), respectively suggest that the failure mode is ductile showing a fibrous fractured surface with dimples.
Fig. 16. Photographs of fractured tensile samples at 400 ◦ C.
3.5. Bend test The photographs of the samples after transverse face and root three-point bend tests are shown in Fig. 18 (a) and (b), respectively. All samples were found to be amenable to bending till 180◦ without the formation of any defects like separation, tearing or fracture etc. as confirmed by visual examination and liquid penetrant tests. Such observations establish the soundness and ductility of the welds prepared by EBW process. 3.6. Impact strength The Charpy impact tests were conducted to determine the resistance of the joints against shocks. Since sub-size samples were tested, the notch impact strength was calculated using the following formula: I = K/A
(1) J/cm2 ;
K = Impact energy absorbed by where, I = Impact strength in the specimen during rupture in joules; and A = Area of cross section(C/S) below the notch before the test in cm2 . The measured Charpy impact strength of the three joints and the two base plates are presented in Table 6. It is seen that the joint-1 prepared using non-oscillating beam possesses minimum impact strength which is even lower than that of copper. This might be attributed to the presence of microcracks and lack of bonding
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Fig. 17. SEM fractographs of tensile samples tested at 400 ◦ C for: (a) joint-1, (b) joint-2, and (c) joint-3.
Table 6 Charpy V-notch impact test results. Specimen Type
Energy absorbed ‘K’ in joule
C/S area at notch ‘A’ in cm2
Impact strength = K/A in J/cm2
Joint-1 Joint-2 Joint-3 Copper plate 304SS plate
14 ± 0.5 23.5 ± 1 17 ± 1 21 ± 1 50 ± 1
0.2 0.2 0.2 0.24 0.24
70 ± 2.5 118 ± 5.0 85 ± 5.0 88 ± 4.0 208 ± 4.0
Fig. 19. Photographs of fractured Charpy V-notch impact test samples.
Fig. 18. Photographs of three-point bend test samples: (a) face bend, and (b) root bend.
between copper and SS in the FZ. However, joint-2 prepared by the oscillating beam having well dispersed copper in the FZ, as evidenced through micrographs (Fig. 5) does possess much better strength. On the contrary, joint-3 also prepared using oscillating beam registered an impact strength value almost same as that of
copper. Unlike joint-2 in joint-3, large mass of segregated copper was observed in the FZ along with microcracks. The presence of such large chunks of copper in the FZ has caused inhomogeneity in the structure thus exhibiting reduced impact strength. Photographs of fractured Charpy test samples are shown in Fig. 19 where the sample identification number corresponds to joint number. For all the joints, fracture paths were traversed from FZ to the HAZ region in the copper side. This might be attributed to the prevalence of a narrow FZ and higher strength than copper. As such,
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Fig. 20. SEM (SEI) fractographs of impact tested specimens for: (a) joint-1; (b) joint-2; and (c) joint-3.
Fig. 21. Sequence of weld pool formation for (a) non-oscillating beam, (b) oscillating beam with oscillation diameter of 1 mm, and (c) oscillating beam with oscillation diameter of 2 mm. The arrows indicate the Marangoni directed flow in the weld pool.
the obtained strength values may not truly represent the strength of the weldments but are indicative only for comparative assessment. The fractured surfaces of the three joints were examined under SEM and the obtained fractographs are shown in Fig. 20 (a)–(c). These micrographs clearly demonstrate that the fracture had occurred in ductile mode since fibrous surfaces with well defined dimples were observed for all the three joints.
4. Hypothesis for better mixing at the weldment due to beam oscillation The major finding of the present study focuses on better mixing of copper with SS under beam oscillation, which has resulted in improvement in some of the mechanical properties of the welded joints. The present hypothesis has been developed to explain the phenomenon facilitating better mixing of copper in the weld pool during beam oscillation. There is clear-cut evidence that better
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mixing has occurred in the weld pool region with beam oscillation at optimum oscillation diameter as shown by Figs. 8–10. It is also noted that the response of such mixing on mechanical properties of the joints like percentage elongation at high temperature and impact strength at room temperature have shown significant improvement. The impact strength of the joint made with beam oscillation at optimum oscillation radius has enhanced by 67% than its counterpart welded without beam oscillation. Although, effect of beam oscillation on mechanical properties obtained at two different testing temperatures are different, but mixing phenomenon during welding with oscillating beam has been found to be the same in all cases as verified by EDS data. During electron beam welding without any beam oscillation, the beam moves in a straight line and the entire weldment is formed as a result of solidification of a sequence of several overlapping weld pools along a straight line as depicted in Fig. 21 (a). Under such situation, the liquid in the weld pool is driven by unidirectional Marangoni convection pushing the liquid radially outward away from the weld line (Rai et al. (2009)). Therefore, copper from the joining interface is pushed towards the periphery of the pool on SS side causing segregation. In contrast, oscillating beam (circular) is likely to produce a sequence of weld pools during its course of movement in a circular path as shown in Fig. 21 (b). For a constant linear beam velocity, it has to traverse longer distance under an oscillating beam, resulting in a higher speed of the beam. Higher beam speed will reduce the heat input per unit length and the resultant pool size would be smaller. During beam movement, several such pools are supposed to cover FZ along the circular path of the beam and there would be significant overlapping between these pools. Assisted by Marangoni convection, copper is likely to move through such intermingled pools in a zigzag path leading to intense mixing of copper in the FZ. In such case, the scale of mixing in the transverse direction would be of the order of several diameters of such tiny weld pools. For the same horizontal welding velocity, when the oscillation diameter is increased, the beam speed is further increased that lowers heat input per unit length and consequently the smaller pools (compared to the case with lower oscillation radius) are formed in sequence as depicted in Fig. 21 (c). In such case, the pool diameter might become so small that the fractional overlapping between them becomes lower than the case with lower beam oscillation radius. This might cause a lower degree of mixing resulting segregation of copper in the form of chunks. 5. Conclusions • The investigation demonstrated that electron beam welded copper-SS joints made with beam oscillation at optimum oscillation diameter (1 mm) possessed 100% more elongation at elevated temperature (400 ◦ C) and 67% more impact strength than its counterpart carried out without beam oscillation. • Adequate mixing of copper in the weld zone and backfilling of the microcracks on SS has been attributed for such improved properties. SEM and EDS analyses with line and area scans of the weld region confirmed the mixing of copper in the weld zone. Further, the EDS spectrum established the backfilling of microcracks by copper.
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• Better mixing of copper in SS with beam oscillation has been explained with a hypothesis. Such mixing in the weldment has been attributed to intense overlapping of tiny weld pools created in a circular path during beam oscillation and subsequent Marangoni directed movement of copper through such overlapped pools. Acknowledgement The authors gratefully acknowledge the financial help received from Board of Research for Nuclear Science, Department of Atomic Energy; Government of India. References ASTM E21-09, 2009. Standard test methods for elevated temperature tension tests of metallic materials. ASTM E3-11, 2011. Standard guide for preparation of metallographic specimens. ASTM E384-111 , 2011. Standard test method for Knoop and Vickers hardness of materials. ASTM E23-12c, 2013. Standard test methods for notched bar impact testing of metallic materials. ASTM E190-14, 2014. Standard test method for guided bend test for ductility of welds. ASTM E8/E8M-15a, 2015. Standard test methods for tension testing of metallic materials. Chen, S., Huang, J., Xia, J., Zhang, H., Zhao, X., 2013. Microstructural characteristics of a stainless steel/copper dissimilar joint made by laser welding. Metall. Mater. Trans. A 44A, 3690–3696. Chen, S., Huang, J., Xia, J., Zhao, X., Lin, S., 2015. Influence of processing parameters on the characteristics of stainless steel/copper laser welding. J. Mater. Process. Technol. 222, 43–51. Cottrell, C.I.M., 1985. Electron beam welding a critical review. Mater. Des. 6, 265–272. Magnabosco, I., Ferro, P., Bonollo, F., Arnberg, L., 2006. An investigation of fusion zone microstructures in electron beam welding of copper-stainless steel. Mater. Sci. Eng. A 424, 163–173. Mai, T.A., Spowage, A.C., 2004. Characterization of dissimilar joints in laser welding of steel-kovar, copper–steel and copper–aluminum. Mater. Sci. Eng. A 374, 224–233. Massalski, T.B., Murray, J.L., Bennett, L.H., Baker, H., 1986. Binary alloy phase diagrams, Volume-1. American Society for Metals Metals Park, OH. Phanikumar, G., Manjini, S., Dutt, P., 2005. Characterization of a continuous CO2 laser welded Fe-Cu dissimilar couple. Metall. Mater. Trans. A 36A, 2137–2147. Rai, R., Palmer, T.A., Elmer, J.W., Debroy, T., 2009. Heat transfer and fluid flow during electron beam welding of 304L stainless steel alloy. Weld. Res. 8, 54s–61s. Roy, C., Pavanam, V.V., Vishnu, G., Hari, P.R., Arivarasu, M., Manikandam, M., Rajkumar, D., Arivazhagan, N., 2014. Characterization of metallurgical and mechanical properties of commercially pure copper and 304 dissimilar weldments. Procedia Mater. Sci. 5, 2503–2512. Saha, T.K., Ray, A.K., Shah, B.K., Bhanumurthy, K., Kale, G., 1999. Electron beam welding of copper to AISI-304 SS. In: International Welding Conference?IWC’99, New Delhi, India, February 15–17. Sajjad, G.S., Mohsen, N., Mahmood, S., Mehdi, S.A., 2012. Gas tungsten arc welding of CP-copper to 304 stainless steel using different filler materials. Trans. Nonferrous Met. Soc. China 22, 2937–2942. Sun, Z., Karppi, R., 1996. The application of electron beam welding for the joining of dissimilar metals: an overview. J. Mater. Process. Technol. 59, 257–267. Tosto, S., Nenci, F., Jiandong Hu Corniani, G., Pierdominici, F., 2003. Microstructure of copper-AISI 304L electron beam welded alloy. Mater. Sci. Technol. 19, 519–522. Velu, M., Bhat, S., 2013. Metallurgical and mechanical examinations of steel–copper joints arc welded using bronze and nickel-base super-alloy filler materials. Mater. Des. 47, 793–809. Yao, C., Xu, B., Zhang, X., Huang, J., Fu, J., Wu, Y., 2009. Interface microstructure and mechanical properties of laser welding copper-steel dissimilar joint. Opt. Lasers Eng. 47, 807–814. Zhang, B.G., Zhao, J., Li, X., Feng, J., 2014. Electron beam welding of 304 stainless steel to QCr0.8 copper alloy with copper filler wire. Trans. Nonferrous Met. Soc. China 24, 4059–4066.
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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Effect of beam oscillation on electron beam welding of copper with AISI-304 stainless steel Jyotirmaya Kar ∗ , Sanat Kumar Roy, Gour Gopal Roy Department of Metallurgical and Materials Engineering, Indian Institute of Technology, Kharagpur 721302, India
a r t i c l e
i n f o
Article history: Received 10 November 2015 Received in revised form 1 March 2016 Accepted 1 March 2016 Available online 3 March 2016 Keywords: Dissimilar welding Electron beam welding Beam oscillation Stainless steel Copper
a b s t r a c t The quality of dissimilar weld joint between copper and stainless steel (SS) plates using electron beam welding with and without beam oscillation was evaluated through microstructural analysis, hardness, tensile, bend and impact tests. Microstructural analysis as well as quantitative elemental analyses demonstrated that improved mixing of copper with SS in the fusion zone could be achieved by an oscillating beam of optimum oscillation diameter. Although the tensile properties at room temperature did not vary much between the welds; high temperature measurements exhibited significant difference between them. Welds made with beam oscillation having an optimum oscillation diameter demonstrated significant increase in impact strength as well as percentage elongation compared to joints made without beam oscillation. Increased beam oscillation beyond optimum oscillation diameter was found to be detrimental promoting segregation of copper in the weld zone exhibiting deteriorated mechanical properties. A hypothesis has been put forward for better mixing during welding with beam oscillation. © 2016 Elsevier B.V. All rights reserved.
1. Introduction For the last few decades, dissimilar-metal joints are widely being used in various industries with an aim to reduce weight and cost savings due to less use of one of the components. In addition, they are capable in offering more complex functions in various engineering applications. Copper-SS dissimilar joint is one of such case, which is designed to utilize the excellent heat and electrical conductivity imparted by copper whereas, better strength, wear and corrosion resistance to be imparted by steel. Accordingly, copperSS joints find their applications in the field of power generation and transmission industries, heat transfer components, cryogenic sector, electrical and electronics sector etc. However, fusion welding of copper-SS is more challenging because of their differences in the physical, chemical and thermomechanical properties. For instance, thermal conductivity of copper is much higher than that of steel which causes difficulties in reaching its melting temperature during fusion welding. Besides, Cu and Fe have limited solid solubility among themselves (Massalski et al. (1986)) and likely to undergo liquid separation under comparatively higher cooling rate. The two materials also have large differences in their melting temperature and coefficient of thermal
∗ Corresponding author. E-mail address: [email protected] (J. Kar). http://dx.doi.org/10.1016/j.jmatprotec.2016.03.001 0924-0136/© 2016 Elsevier B.V. All rights reserved.
expansion. Fusion welding of copper-SS dissimilar couples using low energy density conventional arc welding processes like GTAW, SMAW and GMAW etc. were also found to be difficult. In such welding processes selection of proper filler metal plays a crucial role, as reported by Sajjad et al. (2012), Velu and Bhat (2013) and Roy et al. (2014). Owing to various advantages like contamination free weld, high energy density, precise positioning of the heat source, narrow heat affected zone (HAZ) etc., electron beam welding (EBW) is considered as one of the most sophisticated fusion welding technique that could be used in several component manufacturing processes where conventional arc welding fails (Cottrell (1985)). Sun and Karppi (1996) presented an overview on the applications of EBW for the joining of dissimilar metals and demonstrated the advantages of EBW over conventional fusion welding processes. Converse to EBW process, laser beam welding has certain limitation such as to cope with the high reflectivity of copper. Mai and Spowage (2004) in their study focused a laser beam 0.2 mm into the steel surface and thus obtained very limited amount (<2%) of copper in the fusion zone (FZ). Consequently, a complete metallurgical bond at the interface between copper plate and molten steel could not be achieved. Yao et al. (2009) produced a butt joint with an inclined joint interface where by off-setting the laser beam on the steel side the dilution between copper and steel in the FZ was controlled. Phanikumar et al. (2005) studied Fe-Cu dissimilar couple and reported heterogeneous microstructure in the FZ and such het-
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erogeneity was found to decrease with lower welding speed. Chen et al. (2013) produced copper-SS dissimilar joint by laser welding and inferred that melting of copper in small amount might be beneficial for back-filling of the microcracks in the FZ; however, its larger amount might cause more inhomogeneity both in composition as well as in stress level leading to severe microcrack generation. Saha et al. (1999) investigated copper-SS butt joint using EBW and reported that better mechanical properties of the weld could be obtained by off-setting beam towards the copper side (0.5–1 mm) than running the beam along the joint interface. This resulted in melting of copper in the FZ and back-filling of the microcracks by copper. They also reported that higher welding speed (1000 mm/min and above) is beneficial in reducing the microcrack formation in the FZ. Tosto et al. (2003) successfully prepared defect free thicker (25 mm) dissimilar joints of copper-SS using EBW without beam off-setting. However, they used comparatively higher beam current (395 mA) and lower beam voltage (15 kV) than reported by earlier investigators. Magnabosco et al. (2006) reported joining of copper with SS having higher plate thickness (30–70 mm) using both single and double pass method under different set of operating parameters. They reported more heterogeneity and defects in the microstructure as the plate thicknesses were increased and sought determination of proper optimized process parameters. Unlike the use of a static beam, beam oscillation is supposed to produce a churning action of the liquid in the weld-pool, which in effect would improve the mixing and restrict the segregation providing a more uniform dispersion of copper in the weldment region. This is in contrast to the weld-pool created without using beam oscillation, where directed liquid flow due to Marangoni convection brings in directional properties. It may be pointed out that beam oscillation as a process parameter in EBW for achieving better mixing in the FZ of dissimilar metal joints, particularly for copper-SS has not yet been studied. Accordingly, the present study focuses on the microstructural evolution and its correlation to the mechanical properties of electron beam welded joints of copper (C10300) and AISI-304 SS using both oscillating and non-oscillating beam. The microstructures of the FZ, FZ-base plate interfaces and HAZ, and mechanical properties like micro-hardness, tensile strength (both room and elevated temperature), three-point bending and impact strength were investigated in details. Three specimens for each of the experimental conditions were tested to achieve reproducibility. 2. Experimental procedure 2.1. Materials and design of joints Copper (C10300) and AISI-304 SS plates with dimensions 200 × 80 × 3 mm3 (length × breadth × height) were prepared from the as received plates. The spectroscopic analyses (X-ray fluorescence spectroscopy) of the base plates, namely copper and SS are presented in Table 1. Subsequently, the joint-side face of each plate (transverse to rolling direction) were machined and polished with 240-grit emery paper. Prior to welding, the plates were thoroughly degreased and cleaned with acetone. The two abutting plates were tight fitted using clamps to ensure negligible gap between the plates. Subsequently, the plates were electron beam butt welded using the process parameters as stated in Table 2. 2.2. Process parameters used in EBW Welding of the dissimilar materials were carried out using an 80kV–12 kW EBW machine at IIT Kharagpur, indigenously designed and fabricated by the Bhabha Atomic Research Center (BARC), Department of Atomic Energy (DAE), India.
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At the beginning, EBW experiments were carried out by autogenous bead-on-plate welding of 3 mm thick 304-SS and copper plates with varying beam current, beam voltage and welding speed. Such an exercise was necessary to understand their effects on weld bead profile and to optimize the process parameters in achieving full penetration joints both for oscillating and non-oscillating beam. Three different sets of welding parameters were used to prepare the joints. Out of the three joints, two were prepared using oscillating beam with two different oscillation diameter (1 and 2 mm) and the other was prepared with a non-oscillating beam. During welding, the beam was always focused on the base plate surface and aligned to the butt line, maintaining the beam voltage and welding speed constant. It may be noted that while changing the beam oscillation parameter, current could not be maintained constant and it varied between 12–23%. This is attributed to the fact that during butt welding of copper with SS using beam oscillation, the residence time of the beam increases on both sides of the joint. Since copper conducts heat faster, prolonged dwelling of the beam on copper side favours dissipation of more heat by conduction. This situation demands more power input to compensate for the heat loss in the production of defect free welds. Since beam voltage remained constant during welding, this power compensation was achieved through increase in the beam current. It may be noted that such threshold current requirement for a particular oscillation diameter was established through several trial experiments. Apart from the above mentioned process parameters, other important process variables were maintained constant as listed in Table 3. 2.3. Characterization of the joints After welding, each prepared joint was inspected visually as well as examined by X-ray radiography for its soundness. Only after ensuring its soundness, further study on evolution of microstructures and evaluation of mechanical properties were undertaken. For microscopic observation specimens were cut perpendicular to the weld bead using a wire-cut electric discharge machining (EDM). Subsequently, each specimen was polished using a series of SiC emery papers from 240 to 4000 grit size in succession with water as the coolant, followed by fine polishing using 0.05 m alumina paste in a velvet cloth disc. For metallographic sample preparation standard ASTM E3-11 (2011) method was followed. To reveal the microstructure, aquaregia (75% HCl + 25% HNO3 ) and solution of 5gm FeCl3 + 50 ml HCl + 100 ml water were used as the etchants for SS and copper side, respectively. The etched samples were studied under both optical microscope (Laica® : DM6000 M) and scanning electron microscope (SEM) (Zeiss® : EVO-60). Distribution of chemical elements in the FZ was analyzed through elemental line scan as well as area mapping by using energy dispersive X-ray spectroscopy (EDS). For room temperature tensile tests, transverse tensile specimens (length perpendicular to welding direction) were prepared using EDM and tested according to standard ASTM E8/E8m-15a (2015). Likewise, for elevated temperature tensile tests at 400 ◦ C, transverse tensile specimens were prepared and tested as per standard ASTM E21-09 (2009). Fig. 1 (a) and (b) show the sample design and dimension for tensile tests at room and elevated temperatures, respectively. For both types of tensile tests, cross-head speed (displacement rate) of 0.5 mm/min was chosen. The room and the elevated temperature tensile tests were performed in a 100 kN (Instron® : 8862) and 250 kN (Instron® : 8800) tensile testing machine, respectively. Transverse face and root three-point bend tests were performed as per standard ASTM E190-14 (2014), to investigate the ductility and soundness of the weld joints. Standard specimens for this test were prepared, having 150 mm length and 30 mm breadth. A
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Table 1 Chemical composition of AISI-304 SS and copper (C10300) in weight%. AISI-304SS
Cr 18.06 Cu 99.97
Copper C10300
Ni 8.06 Pb 0.0001
Mn 1.802 P 0.003
Mo 0.15 Fe 0.001
Si 0.36 O 0.003
C 0.043 Ni 0.003
Al P 0.06 0.042 Total impurities Total impurities 0.03% Max.
Fe Rest
Table 2 EBW process parameters used in this study. Joint No
Beam Voltage (kV)
Beam Current (mA)
Welding Speed (mm/min)
Oscillation Frequency (Hz)
Oscillation diameter (mm)
1 2 3
60 60 60
65 73 80
1000 1000 1000
Not used 600 600
Not used 1.0 2.0
Table 3 Other important welding process variables. Gun chamber vacuum (mbar)
<5 × 10−6
Gun-specimen distance (mm)
460
Welding chamber vacuum (mbar) Beam focus point (mm)
<5 × 10−5 0
Welding pass (#) Beam off-set
1 none
Fig. 3. Sample design and dimension for Charpy impact test.
Fig. 1. Sample design and dimension for the tensile tests at: (a) room temperature, and (b) 400 ◦ C.
as per Fig. 3 and tested following the standard ASTM E23-12c (2013) method at ambient temperature of 25 ◦ C in an impact machine (Instron® : SI-1C3). Micro-hardness values were also measured following the standard ASTM E384-111 (2011) across the weld metal at an interval of 0.2 mm using a Vickers micro-hardness testing machine (UHL® : VMHT-001) under a load of 100gf/15s (load/dwell time). To identify the effect of microstructural heterogeneity such measurements were performed for both the parent metal, HAZ and FZ regions. 3. Results and discussion 3.1. Study of microstructure and elemental analysis
Fig. 2. Schematic diagram illustrating the method of three-point face/root bend test.
support plate of 304-SS having thickness of 2.5 mm was used to compensate the difference in strength between copper and SS. Furthermore, the upper plunger was shifted towards the SS side by 10 mm from the weldment region to prevent horizontal shifting of the plunger during bending towards the copper side. The test setup is shown in Fig. 2 and the three-point bend tests were performed in a 50 kN universal testing machine (Shimadzu® : AG-5000G). To evaluate the impact strength of the joints, sub-size Charpy impact specimens (due to smaller plate thickness) were prepared
Fig. 4 (a) shows an overview in the micrograph for joint-1 at a lower magnification while Fig. 4 (b)–(d) show magnified images of the different regions for the same joint obtained using nonoscillating beam. Fig. 4 (b) clearly demonstrates that the major melting had occurred on the SS side, while limited melting took place on the copper side. Moreover, copper-FZ interface seems to be rough with significant presence of iron [Fig. 4 (d)]. Fig. 4 (c) shows no traces of copper in the FZ rather comprises of re-melted and resolidified SS only. The cellular dendritic structure is well evident in this zone. The embedded figure in Fig. 4 (b) at a higher magnification shows that the FZ-SS interface is smooth and comprised of cellular dendritic morphology. The corresponding microstructures of joint-2, which was prepared with an oscillating beam of oscillation diameter 1 mm, are shown in Fig. 5 (a)–(c). Here, a significant amount of copper could be found in the FZ as shown in Fig. 5 (b). It may be noted that during beam oscillation, the beam dwells on both sides of the joint including the copper region for a certain amount of time, which not only promotes heat dissipation but also melts copper locally. It was also observed that under oscillating beam both the copper-FZ interface
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Fig. 4. Optical microscopic images for joint-1: (a) overview of the whole joint, (b) SS-FZ interface, (c) middle portion of the FZ, and (d) FZ-copper interface.
Fig. 5. Optical microscopic images for joint-2: (a) SS-FZ interface, (b) middle portion of the FZ, and (c) FZ-copper interface.
[Fig. 5 (c)] and SS-FZ interface [embedded in Fig. 5 (a)] are comparatively smoother than in the case of non-oscillating beam (Fig. 4). This might be attributed to uniform heat dissipation in all directions by the churning action of the liquid melt. Fig. 5 (b) further reveals that copper is uniformly dispersed in the FZ. Similarly, Fig. 6 (a)–(c) depict the microstructures of different regions of joint-3 (obtained with an oscillating beam of oscillation diameter 2 mm). It demonstrates that a large amount of copper is present both in the middle of the FZ as well as adjacent to copper side [Fig. 6 (b) and (c)]. It is evident from the microstructures that more copper got melted during the longer dwelling of the beam through the copper region during its oscillation. However, the molten copper in the weld metal on its solidification appears in the form of chunks and segregated mass [Fig. 6 (b)]. This may be attributed to the fact that churning action generated due to beam oscillation is not sufficient to disperse the large amount of molten
copper uniformly all over the weld metal. The SS-FZ interface is also found to be more rough compared to the joint prepared with lower beam oscillation diameter. This might be attributed to a less degree of mixing under higher beam oscillation diameter leading to directed heat flow. All joints regardless of being prepared either using a nonoscillating beam or oscillating beam were found to possess only a narrow HAZ region on the copper side adjacent to FZ as shown in Fig. 7. The HAZ region is distinguished by coarse equiaxed polyhedral grains with annealing twins compared to finer grains in the copper base metal. Copper having higher thermal conductivity allows large amount of heat to flow through it and causes grain growth in the HAZ. Furthermore, optical microscopic analysis of the joint cross-sections did not reveal any gross defects like microcracks, porosity, etc.
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Fig. 6. Optical microscopic images for joint-3: (a) SS-FZ interface, (b) middle portion of the FZ, and (c) FZ-copper interface.
Fig. 7. Optical microscopic image depicting the HAZ region on the copper side for joint-1 in etched condition.
Fig. 8. EDS profiles for joint-1: (a) SEM (BSE) image of Joint-1, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
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179
Fig. 9. EDS profiles for joint-2: (a) SEM (BSE) image of Joint-2, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
Fig. 10. EDS profiles for joint-3: (a) SEM (BSE) image of Joint-3, (b) line scan profile for Cu-K␣, (c) line scan profile of Fe-K␣, (d) area map of Cu-K␣, and (e) area map of Fe-K␣.
Figs. 8–10 show EDS elemental line scan and area mapping for Cu and Fe across the FZ for joint-1, 2 and 3, respectively. Fig. 8 (a)–(e) depict the BSE SEM image, line scan and area mapping for copper and iron in joint-1, prepared with non-oscillating beam. The line corresponding to the EDS scan path and the rectangle corresponding to EDS area mapping are shown in Fig. 8 (a). It may be noted
that the line scan covers the whole weld region from the Cu-weld interface to SS-weld interface. Both the line scan [Fig. 8 (b) and (c)] as well as area mapping [Fig. 8 (d) and (e)] confirm that a narrow copper rich region exists near the copper side of FZ with a negligible amount of copper dispersed across the rest of the FZ, while the majority of it is rich in iron. Fig. 9 that represents similar plots
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Fig. 11. FE-SEM (BSE) images showing microcracks: (a) joint-1, (b) enlarged view of the insert for joint-1, (c) for joint-2, (d) corresponding EDS point spectrum bitmaps, (e) joint-3, and (f) enlarged view of the insert for joint-3.
for joint-2 prepared with an oscillating beam of oscillation diameter 1 mm. This figure [Fig. 9 (b)–(e)] clearly exhibits that though the copper rich region still exists near the copper-FZ interface, but significant amount of it also appears in the FZ. Fig. 9 (b) further depicts that copper is more uniformly dispersed in the FZ. The line scan and area mapping as shown in Fig. 10 (b)–(e) shows that for joint-3, obtained with beam oscillation at higher oscillation diameter, the amount of copper in the FZ is more and segregated as, corroborated by the optical microstructure (Fig. 6). Fig. 11 (a)–(f) shows SEM image of joint-1, 2 and 3, respectively taken at different locations near the copper rich mass in the FZ. It is revealed that all the three joints have microcracks in the FZ. The formation of such microcracks might be attributed to inhomogeneity in composition and complex stress distribution in the FZ during the
Fig. 12. Micro-hardness profile across the SS-copper bimetallic joints.
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Fig. 13. Stress-strain curves for: (a) copper-SS joints, and (b) copper base plate.
solidification of the weld-pool. It is apparent from Fig. 11 (b) and (f) that the microcracks are of open type and these are not backfilled by copper, for joint-1 and 3, respectively. In contrast, Fig. 11 (c) representing the SEM image of joint-2, depicts the presence of tiny copper globules on the line of microcracks. EDS analysis of a spot on the microcrack [spectrum 1 in Fig. 11 (d)] also suggests the presence of large amount of copper within the microcrack. Thus, in case of joint-2, the microcracks are readily back-filled by copper. When copper gets mixed well in the form of tiny droplets across the weld, it increases the probability of finding copper droplets on microcracks as depicted in Fig. 11 (c). Since copper remains liquid while iron gets solidified, copper wets the microcracks generated on SS during its solidification and fills it. But if the mixing is not sufficient liquid copper cannot move further from its source. Accordingly, it appears in the form of segregated mass during solidification and finds limited scope to backfill the microcracks situated far away. Besides, segregation of copper introduces more inhomogeneity and thermal stress that generates more microcracks. Possibly for such reasons, there are large open cracks in joint -3. However, in the case of joint-1, since copper does not melt appreciably and mixing is insufficient cracks remain open.
Table 4 Room temperature tensile properties of the bimetallic welded joints and copper base plate. Material
0.2% YS (MPa)
Joint-1 Joint-2 Joint-3 Copper plate
168 171 170 205
± ± ± ±
3 4 3 4
UTS (MPa) 249 252 250 261
± ± ± ±
2 3 2 2
Elongation (%) 17 19 18 36
± ± ± ±
2 2 2 2
Failure Location copper side copper side copper side
3.2. Micro-hardness The variation in micro-hardness across the weld cross-section for each of the three prepared joints is shown in Fig. 12. It is observed that all the prepared joints possess higher hardness in the FZ than the copper base plate. Furthermore, there exists a narrow HAZ region on the copper side. A deep in hardness adjacent to the FZ/HAZ interface on the copper side might be attributed to its coarse grained structure as revealed in the microstructural analysis (Fig. 7). It may be noted that fluctuation in micro-hardness values are maximum in the FZ of joint-3; while it is minimum for the joint-1 and moderate for joint-2. Large fluctuations in hardness values in joint-3 might be attributed to greater amount of copper segregation in the FZ. In joint-2, since only moderate amount of copper is finely distributed, fluctuation becomes moderate. For joint-1 there is hardly any variation in hardness along the FZ since very little copper got melted during the joining process. Therefore, hardness variation also corroborates the microstructural investigation regarding the extent of copper mixing in the FZ. 3.3. Room temperature tensile properties The stress-strain plots for the three joints and that for the copper base plate, tested at room temperature, are shown in Fig. 13 (a)
Fig. 14. Photographs of fractured room temperature tensile samples.
and (b), respectively. Results pertaining to the tensile tests such as 0.2% off-set yield strength (YS), ultimate tensile strength (UTS) and percentage elongation up to rupture etc. were evaluated and are presented in Table 4. The relevant values for copper base plate (as it is the weakest) are also included in Table 4. Table 4 and Fig. 13 (a) clearly depict that all the joints possess almost same room temperature tensile properties. The UTS for all the joints were found to be more than 95% of copper signifying that the joints achieved adequate strength. However, the percentage elongation for all the three joints is almost half than that of copper. Fig. 14 shows the fractured test samples where the sample identification number corresponds to joint number. It is observed that for all the joints, the SS-FZ interface remained intact while fracture had occurred on the copper side, suggesting that the characteristic of the base metal remains uninfluenced by the presence or absence of beam oscillation. The tensile test results closely match the data reported by Tosto et al. (2003), Zhang et al. (2014) and Chen et al. (2015). In arc welding processes (SMAW); Velu and Bhat (2013) and Roy et al. (2014) reported that failure occurred in the HAZ region on the copper side.
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Fig. 15. Stress-strain curves at 400 ◦ C for: (a) copper-SS joints, and (b) copper base plate.
Table 5 Tensile properties of the welded joints and copper base plate at 400 ◦ C. Material
0.2% YS (MPa)
Joint-1 Joint-2 Joint-3 Copper plate
149 149 146 169
± ± ± ±
1 1 3 2
UTS (MPa) 179 183 180 189
± ± ± ±
2 1 2 2
Elongation (%) 7 14 8 48
± ± ± ±
1 2 1 2
Failure Location FZ-SS interface HAZ/ copper HAZ/ copper copper
3.4. Elevated temperature tensile properties To determine the high temperature serviceability of the joints, tensile tests were performed at 400 ◦ C and the corresponding results are presented in Table 5. For comparison, the properties of copper base plate (at 400 ◦ C) were also evaluated and are shown in the same Table 5. The corresponding stress-strain plots for the three joints as well as copper base plate are shown in Fig. 15 (a) and (b), respectively. It is clearly seen that even though all three joints demonstrated almost the same YS and UTS values, but the percentage elongation is almost twice in case of joint-2 compared to joint-1 and 3. Therefore, the area under the stress-strain curve, which represents the toughness, becomes double for joint-2. This indicates that joint2 with adequate mixing of copper in the FZ possesses better high temperature toughness rendering better serviceability. Similar to room temperature test results, here also the UTS for all the joints were 95% more than that of copper, but the percentage elongation was lower. The fractured tensile samples are shown in Fig. 16 where the sample identification number corresponds to joint number. It is observed that joint-1 which was prepared using a non-oscillating beam, fractured in the FZ-SS interface indicating improper bonding at the interface. However, joint-2 and 3 prepared using oscillating beam, fractured away from the FZ that is on copper side, while the SS-FZ interface remained intact. Incidentally, the fracture occurred at the interface of the two metallic material in joint-1, while for the other two joints it occurred away from the FZ. It can be inferred that joints prepared using oscillating beam do possess satisfactory tensile properties at elevated temperature. Since the tensile samples tested at 400 ◦ C, fractured at different locations, fracture surface morphologies were expected to be different. The fractographs of the failed samples are presented in Fig. 17 (a)–(c). Fig. 17 (a) corresponds to the fracture surface of joint1 showing some cracks. However, fractographs of joint-2 and 3 as shown in Fig. 17 (b) and (c), respectively suggest that the failure mode is ductile showing a fibrous fractured surface with dimples.
Fig. 16. Photographs of fractured tensile samples at 400 ◦ C.
3.5. Bend test The photographs of the samples after transverse face and root three-point bend tests are shown in Fig. 18 (a) and (b), respectively. All samples were found to be amenable to bending till 180◦ without the formation of any defects like separation, tearing or fracture etc. as confirmed by visual examination and liquid penetrant tests. Such observations establish the soundness and ductility of the welds prepared by EBW process. 3.6. Impact strength The Charpy impact tests were conducted to determine the resistance of the joints against shocks. Since sub-size samples were tested, the notch impact strength was calculated using the following formula: I = K/A
(1) J/cm2 ;
K = Impact energy absorbed by where, I = Impact strength in the specimen during rupture in joules; and A = Area of cross section(C/S) below the notch before the test in cm2 . The measured Charpy impact strength of the three joints and the two base plates are presented in Table 6. It is seen that the joint-1 prepared using non-oscillating beam possesses minimum impact strength which is even lower than that of copper. This might be attributed to the presence of microcracks and lack of bonding
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Fig. 17. SEM fractographs of tensile samples tested at 400 ◦ C for: (a) joint-1, (b) joint-2, and (c) joint-3.
Table 6 Charpy V-notch impact test results. Specimen Type
Energy absorbed ‘K’ in joule
C/S area at notch ‘A’ in cm2
Impact strength = K/A in J/cm2
Joint-1 Joint-2 Joint-3 Copper plate 304SS plate
14 ± 0.5 23.5 ± 1 17 ± 1 21 ± 1 50 ± 1
0.2 0.2 0.2 0.24 0.24
70 ± 2.5 118 ± 5.0 85 ± 5.0 88 ± 4.0 208 ± 4.0
Fig. 19. Photographs of fractured Charpy V-notch impact test samples.
Fig. 18. Photographs of three-point bend test samples: (a) face bend, and (b) root bend.
between copper and SS in the FZ. However, joint-2 prepared by the oscillating beam having well dispersed copper in the FZ, as evidenced through micrographs (Fig. 5) does possess much better strength. On the contrary, joint-3 also prepared using oscillating beam registered an impact strength value almost same as that of
copper. Unlike joint-2 in joint-3, large mass of segregated copper was observed in the FZ along with microcracks. The presence of such large chunks of copper in the FZ has caused inhomogeneity in the structure thus exhibiting reduced impact strength. Photographs of fractured Charpy test samples are shown in Fig. 19 where the sample identification number corresponds to joint number. For all the joints, fracture paths were traversed from FZ to the HAZ region in the copper side. This might be attributed to the prevalence of a narrow FZ and higher strength than copper. As such,
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Fig. 20. SEM (SEI) fractographs of impact tested specimens for: (a) joint-1; (b) joint-2; and (c) joint-3.
Fig. 21. Sequence of weld pool formation for (a) non-oscillating beam, (b) oscillating beam with oscillation diameter of 1 mm, and (c) oscillating beam with oscillation diameter of 2 mm. The arrows indicate the Marangoni directed flow in the weld pool.
the obtained strength values may not truly represent the strength of the weldments but are indicative only for comparative assessment. The fractured surfaces of the three joints were examined under SEM and the obtained fractographs are shown in Fig. 20 (a)–(c). These micrographs clearly demonstrate that the fracture had occurred in ductile mode since fibrous surfaces with well defined dimples were observed for all the three joints.
4. Hypothesis for better mixing at the weldment due to beam oscillation The major finding of the present study focuses on better mixing of copper with SS under beam oscillation, which has resulted in improvement in some of the mechanical properties of the welded joints. The present hypothesis has been developed to explain the phenomenon facilitating better mixing of copper in the weld pool during beam oscillation. There is clear-cut evidence that better
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mixing has occurred in the weld pool region with beam oscillation at optimum oscillation diameter as shown by Figs. 8–10. It is also noted that the response of such mixing on mechanical properties of the joints like percentage elongation at high temperature and impact strength at room temperature have shown significant improvement. The impact strength of the joint made with beam oscillation at optimum oscillation radius has enhanced by 67% than its counterpart welded without beam oscillation. Although, effect of beam oscillation on mechanical properties obtained at two different testing temperatures are different, but mixing phenomenon during welding with oscillating beam has been found to be the same in all cases as verified by EDS data. During electron beam welding without any beam oscillation, the beam moves in a straight line and the entire weldment is formed as a result of solidification of a sequence of several overlapping weld pools along a straight line as depicted in Fig. 21 (a). Under such situation, the liquid in the weld pool is driven by unidirectional Marangoni convection pushing the liquid radially outward away from the weld line (Rai et al. (2009)). Therefore, copper from the joining interface is pushed towards the periphery of the pool on SS side causing segregation. In contrast, oscillating beam (circular) is likely to produce a sequence of weld pools during its course of movement in a circular path as shown in Fig. 21 (b). For a constant linear beam velocity, it has to traverse longer distance under an oscillating beam, resulting in a higher speed of the beam. Higher beam speed will reduce the heat input per unit length and the resultant pool size would be smaller. During beam movement, several such pools are supposed to cover FZ along the circular path of the beam and there would be significant overlapping between these pools. Assisted by Marangoni convection, copper is likely to move through such intermingled pools in a zigzag path leading to intense mixing of copper in the FZ. In such case, the scale of mixing in the transverse direction would be of the order of several diameters of such tiny weld pools. For the same horizontal welding velocity, when the oscillation diameter is increased, the beam speed is further increased that lowers heat input per unit length and consequently the smaller pools (compared to the case with lower oscillation radius) are formed in sequence as depicted in Fig. 21 (c). In such case, the pool diameter might become so small that the fractional overlapping between them becomes lower than the case with lower beam oscillation radius. This might cause a lower degree of mixing resulting segregation of copper in the form of chunks. 5. Conclusions • The investigation demonstrated that electron beam welded copper-SS joints made with beam oscillation at optimum oscillation diameter (1 mm) possessed 100% more elongation at elevated temperature (400 ◦ C) and 67% more impact strength than its counterpart carried out without beam oscillation. • Adequate mixing of copper in the weld zone and backfilling of the microcracks on SS has been attributed for such improved properties. SEM and EDS analyses with line and area scans of the weld region confirmed the mixing of copper in the weld zone. Further, the EDS spectrum established the backfilling of microcracks by copper.
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