- Email: [email protected]
Study on explosive welding for manufacturing meshing bonding interface of CuCrZr to 316L stainless steel
Fusion Engineering and Design 143 (2019) 106–114
Contents lists available at ScienceDirect
Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Study on explosive welding for manufacturing meshing bonding interface of CuCrZr to 316L stainless steel
T
⁎
Ming Yanga, Honghao Maa,b, , Zhaowu Shena, Yuling Suna,c a
CAS Key Laboratory of Mechanical Behavior and Design of Materials, Department of Modern Mechanics, University of Science and Technology of China, Hefei, Anhui 230027, China b State Key Laboratory of Fire Science, University of Science and Technology of China, Hefei, Anhui 230026, China c PLA Army Academy of Artillery and Air Defense, Hefei, Anhui 230031, China
A R T I C LE I N FO
A B S T R A C T
Keywords: Explosive welding Meshing interface Mechanical properties Microstructure
In order to improve the CuCrZr/316L transition joints manufactured by explosive welding (EW), a new EW method was developed to obtain meshing bonding interfaces. The welding parameters adopted in the experiment were close to the lower limit, and the bonding quality was evaluated by systematically microstructure and mechanical tests. The results showed that the meshing interfaces can effectively enhance CuCrZr/316L EW-joints by increasing bonding area. The microstructure results indicated that metallurgical bonding of CuCrZr alloy to 316L stainless steel was created along the meshing interfaces, and two types of bond were observed at the bonding interfaces, namely metal/metal and metal/solidified melt. The EDS analysis indicated that the melt mainly consisted of Cu. Fractography studies on tensile specimens showed cleavage fracture on the CuCrZr side and ductile fracture on the 316L side near the interfaces. The tensile shear test results showed that the shear strength of the meshing interface 0° and 90° was increased by 19% and 9%, when compared to the ordinary CuCrZr/316L EW-joints. The results of microhardness test revealed that the values of microhardness decreased as the distance from the interface increased.
1. Introduction As a component directly faces the plasma, the first wall (FW) of ITER will bear extremely high heat fluxes about 20–30 MW/m2 [1,2]. CuCrZr alloy was chosen as the heat sink material due to its excellent thermal conductivity, which was bonded to an austenitic stainless steel (316 L) support structure. Recent developments in the design of the FW required that the CuCrZr/316 L joints directly faced the internal grooves, which formed the pressure boundary for the cooling water [3]. Thus renewed attention is being given to manufacturing excellent CuCrZr/316 L joints which will withstand the internal pressure, the fatigue loads and remain leak tight during the cyclic mode of operation [4]. Several methods such as friction welding [5,6], brazing [7], diffusion bonding [8], roll bonding [9], hot isostatically pressing (HIP) [10], and explosive welding (EW) [3,11–14] have been investigated for joining copper alloy to steel. When compared to other welding methods, EW shows many advantages such as high bonding strength and no long range interdiffusion [3,14]. Thus this technology is attracting more and more attention from ITER researchers [3,14]. EW is a
⁎
well-known solid state method for joining various metal materials, which employs explosive energy to make metals produce a high-speed oblique collision and achieve metallurgical bond [15]. Since external heating and large-scale melt are unnecessary for this process, it is capable of keeping the physical and chemical properties of the wrought parent components unchanged [16]. Moreover, because it enables to obtain metallurgical bonding on over the entire junction surfaces without any oxidation, explosive welding joints generally have relatively high strength [17]. Therefore, EW is very suitable for welding such CuCrZr/316 L dissimilar bulk metals. Several studies on manufacturing CuCrZr/316 L bimetals have been conducted using conventional EW technology. Goods et al [3] obtained the CuCrZr/316 L joints using two different solid state techniques, namely HIP and EW, and they found that EW could reduce or eliminate HIP’s shortcomings such as excessive grain growth, overaging and further loss of critical properties. Ma et al [14] provided a new EW method for manufacturing ITER-grade 316 L/CuCrZr hollow structural member, and the results showed that the welded specimens were sound. Wang et al [18] investigated the microstructure and mechanical properties evolution of post-weld heat treatment and final Be/CuCrZr HIP on the interface of
Corresponding author. E-mail address: [email protected] (H. Ma).
https://doi.org/10.1016/j.fusengdes.2019.03.137 Received 20 June 2018; Received in revised form 10 December 2018; Accepted 21 March 2019 0920-3796/ © 2019 Published by Elsevier B.V.
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Pmin is minimum collision pressure determined by the formula below [21].
CuCrZr/316 L EW-plates. The results indicated that EW was qualified for the manufacturing of ITER FW components in China. Although EW shows many advantages when compared to other welding methods, the applying CuCrZr/316 L EW-joints on ITER still faces a lot of challenges due to the poor working condition. During the cyclic mode of operation, the welding joints are directly exposed to the cooling water, which will be continuously impacted by high pressure fluid and become the weakest part of a component. As a result, the ordinary CuCrZr/316 L EW-joints may fail in service due to the low bonding strength. Thus, the goal of this work is to increase the bonding strength of CuCrZr/316 L EW-joints. For this purpose, we explored a new EW technology for manufacturing CuCrZr/316 L meshing bonding interface, and the microstructure and mechanical properties of the bonding interfaces were systematically investigated.
pmin =
1 c (ρσb )1/2 2
(2)
Where c, ρ, and σb are the sonic speed, density and tensile strength. The minimum collision pressure Pmin1 of the flyer plate and Pmin2 of the base plate can be calculated respectively through Eq. (2), and the Pmin is the bigger value of Pmin1 and Pmin2. There is a lower limit of collision point velocity vcmin for predicting the laminar–turbulent flow transition of the metal materials to be welded. This limit also defines the transition from straight to wavy interfaces between the flyer and base plates. Wavy interfaces can be obtained when the collision point velocity is larger than the lower limit. Cowan et al. [22] defined lower limit of collision point velocity with hydrodynamic analogy.
2. Experimental procedure
1/2
2R (H + H2) ⎤ vc min = ⎡ e 1 ⎢ ρ1 + ρ2 ⎥ ⎦ ⎣
2.1. Experimental materials
(3)
In this work, CuCrZr alloy and 316 L stainless steel plate were employed as flyer and base plates, which were designed with dimensions of 300 mm × 150 mm × 8 mm and 300 mm × 150 mm × 15 mm. The chemical composition of the CuCrZr and 316 L are shown in Table 1 and Table 2. It is important to accurately control the height and uniformity of explosive in this study. And to do that, honeycomb structure explosive was chosen as explosive material, which was consisted of aluminium honeycomb filled with emulsion explosive. As shown in Fig. 1, the thickness of aluminium foil is 60 μm, the side length of the regular hexagon cell is 8 mm. The components of the emulsion matrix are presented in Table 3. Hollow glass microballoons with a mass fraction of 15% were chosen as the sensitizer. The detonation velocity of the emulsion explosive is about 3000 m/s
Where Re is the critical Reynolds Number taking values of 10.6 [23], H1 and H2 are the Vickers hardness numbers of the flyer and base plates respectively. The material parameters used for calculating the lower limit of the dynamic parameters are listed in Table 4. For parallel set-up geometry, the collision point velocity vc is equal to the detonation velocity vd. The impact velocity vp could be calculated by the following expressions [24,25].
2.2. Welding parameters selection
R=
Good welding quality depends on careful control of the dynamic parameters [12]. There are three mainly dynamic parameters for EW, i.e., collision angle β, collision point velocity vc and impact velocity vp, which are determined by initial parameters such as detonation velocity and mass ratio of explosive to flyer plate. However, the three dynamic parameters are related to each other for certain geometric relationships and only two of them are independent. In this work, the collision point velocity vc and impact velocity vp were used to determine the initial parameters. The initial parameters were designed to meet the condition that the welding dynamic parameters were close to the lower limit of the dynamic parameters. In this condition, good welding quality was usually expected to be obtained [14,19]. In order to make impact pressure at the collision point exceed the yield stress of the materials to promote plastic deformation, the impact velocity vp of flyer plates must be larger than the lower limit of impact velocity vpmin. For welding of dissimilar metallic materials, the lower limit of impact velocity can be calculated by the following formula [20].
Where R is the mass ratio of explosive to flyer plate and E is Gurney energy of explosive. γ is the polytropic exponent of detonation products, the value for emulsion explosive is equal to 2.5. ρ0 and ρ1 are the density of the explosive and flyer plate, t0 and t1 are the thickness of the explosive and flyer plate. The empirical formula of the stand-off distance h between the flyer plate and the base plate was expressed as [26].
1 1 ⎞ vp min = pmin ⎛⎜ + ⎟ c2 ρ2 ⎠ ⎝ c1 ρ1
vp =
Mg
Al
Si
P
Fe
Cu
Content
0.8
0.1-0.8
0.1-0.25
0.1-0.25
0.5
0.1
0.5
Bal.
(4)
1 ⎛ γ ⎞ 2 ⎜ ⎟ v d − 1 ⎝ γ + 1⎠
ρ0 t0 ρ1 t1
(5)
(6)
(7)
Where t0 is the explosive thickness, and t1 is the flyer thickness. The initial parameters used for this EW are shown in Table 5 (calculation according to aforementioned procedure) and the accordingly dynamic parameters and the lower limit of the dynamic parameters are listed in Table 6. 2.3. Experimental methods The meshing bonding interfaces were obtained by prefabricating grooves in flyer and base plates. In order to make bonding quality excellent, there are two key points for manufacturing the grooves. First, the grooves will not prevent the jet and air between the flyer plate and base plate from being out. Second, the groove surfaces of base plate should be in good contact with the flyer plate during collision process. Considering the two points, dovetail grooves were cut along the detonation direction in this study. The dovetail grooves of the flyer plate were the same as those of the base plate, so which could be tightly assembled together. The dimensions of the dovetail grooves are shown in Fig. 2. The success of the experiment depends on accurately controling of the flight path of the flyer plate. For this end, four locating pins were
Table 1 Chemical composition of CuCrZr (wt%). Zr
γ2
h = 0.2(t0 + t1)
Where c1 and c2 are the sonic speed of the flyer plate and base plate respectively, ρ1 and ρ2 are the density of the flyer plate and base plate.
Cr
3R2 R2 + 5R + 4 γ
E=
(1)
Element
2E
107
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Table 2 Chemical composition of 316 L stainless steel (wt%). Element
C
Si
Mn
Cr
Ni
S
P
Mo
Fe
Content
≤0.08
≤1.00
≤2.00
16.00-18.00
10.00-14.00
≤0.03
≤0.05
≤2.00-3.00
Bal.
Fig. 2. Dimensions of dovetail grooves of the flyer and base plate (unit: mm).
Fig. 1. Honeycomb structure explosive. Table 3 Component of the emulsion matrix. Component
NH4NO3
NaNO3
H2O
C18H38
C24H44O6
C12H26
mass fraction
75%
10%
8%
4%
2%
1%
Fig. 3. Schematic diagram of the meshing interface explosive welding. Table 4 Selected material parameters used for calculating the lower limit of the dynamic parameters. Material
Ρ [kg•m−3]
Hv [Mpa]
C [m•s−1]
σb [Mpa]
CuCrZr 316 L
8960 7980
1200 2000
3800 5000
380 520
In order to reveal the interface morphology, specimens were cut parallel and vertical to the detonation direction respectively, and the cross-section of the specimens were ground with emery papers up to No. 5000 and polished to 0.5 um by diamond paste. A scanning electron microscope (GeminiSEM 500) and an optical microscope (Leica DM4M) were employed for microstructure observation of the bonding interfaces. Energy-dispersive X-ray spectrometry (EDS) analysis was also carried out to characterize the distribution of the alloy elements across the bonding interface. In order to investigate the mechanical properties of the welded plate, microhardness, tensile, and tensile-shear tests were carried out. Microhardness tests were conducted on a microhardness machine (HVS1000 M) using a 100 g load for 10 s. The tensile tests were carried out with a tensile strain rate of 1 × 10−4/s according to GB/T6396-2008 [27], and the dimensions of the specimens are shown in Fig. 4a. Fractography studies on broken specimens from tensile tests were carried out using a scanning electron microscope. According to standard test method, in order to ensure that bond separation occurs sooner than bulk material failure during tensile shear tests, the maximum permissible length of overlapped portion (L) was determined by the following equation [28]:
Table 5 Initial parameters used for explosive welding. t0 (mm)
t1 (mm)
ρ0 (g/cm3)
R
h (mm)
vd (m/s)
25
7
0.9
0.36
7
3000
Note: The original thickness of flyer plate was 8 mm, however the value of 7 mm was taken for considering the effect of grooves to mass. The calculated h is 6.4 mm, however the actual stand-off distance of 7 mm was used to ensure that the flyer would be accelerated to the maximum velocity. Table 6 Dynamic parameters for explosive welding. vpmin (m/s)
vcmin (m/s)
vp (m/s)
vc (m/s)
278
2007
311
3000
L= employed to guarantee precise meshing collision between the flyer and base plate, as shown in Fig. 3. The parallel set-up geometry was employed for this EW. A detonator was placed on the side edge of the explosive to make the detonation along the dovetail groove direction. The welding assembly was placed on a steel anvil in an explosion vessel.
σb T τ
(8)
Where σb is the yield point of softer material, T is the thickness and τ should be considered as 1.5 times higher than the estimated average shear bonding strength [28]. This equation ensures that fracture position is at bonding interface, and the dimensions of the specimens are shown in Fig. 4b. 108
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
parameters were close to the lower limit. If high explosive ratio was used, ejection will be formed in between flyer and base plate after the explosion. This results in that melting area in interface and also possible oxidation and dirty surface will not be exported to outside [11,12]. However, Fig. 5d and e show that the lower and inclined interfaces were connected by continuous melting layer, and the thickness of the melting layer was in the range of 50–200 μm. Formation of the melting layer was attributed to the adiabatic heating of gases compressed in the dovetail groove. The air inside the groove is harder to be discharged than that near the upper interface, which resulted in that the air near the lower and inclined interfaces was subjected to adiabatic compression during the collision process. Thus the continuous melting layer at the interfaces was formed under the actions of adiabatic compression heat. One way to reduce the melting layer is to reduce the detonation velocity. According to our previous experiments, the detonation velocity in the initial detonation zone is about 2600 m/s due to the unstable detonation, while it is about 3000 m/s in other zone. Fig. 6 shows the morphology of bonding interface at the initial detonation zone. It can be seen from Fig. 6 that the melting layer is extremely thin (about 8 μm) at the initial detonation zone, which shows that the melt layer can be effectively reduced by decreasing the detonation velocity. Due to the fact that the collision point velocity is equal to the detonation velocity, reducing the detonation velocity is conducive to air emission at the interface so as to reduce the melting layer. In order to describe the transition layer between the two welded materials, element line scans were conducted to determine elements distribution across the bonding interface. As shown in Table 1 and Table 2, the content of Zr in CuCrZr and 316 L is extremely low, so the distribution of Zr across the bonding interface is not presented in Figs. 7 and 8. It can be observed from Fig. 7 that the element distribution line was steep across the upper interface, while a platform appeared across the lower and inclined interfaces, which proved the results obtained by OM, namely that the upper interface was bonded in the form of direct bonding, while the lower and inclined interfaces were joined by melting layer. In addition, it can be observed from Fig. 7b and c that the diffusion layer’s thickness of the inclined interface was greater than that of the lower interface and the elements variation was more intense. This can be attributed to adiabatic shear along the inclined interface. Unlike the upper and lower interfaces, the inclined interface was bonded by inclined collision between flyer and base plates, thus stronger adiabatic shear was expected to occur along the inclined interface, which resulted
Fig. 4. Dimensions of the tensile (a) and tensile shear (b) specimens (unit: mm).
3. Results and discussion 3.1. Microstructure of bonding interfaces Figs. 5 shows optical microscope (OM) images of the microstructural morphology of bonding interface perpendicular to detonation direction. Fig. 5a is the general layout of a meshing interface after EW, which also shows five typical positions (i.e., the rectangular areas b, c, d, e, and f in Fig. 5a), and corresponding high magnification micrographs around these five positions are shown in Fig. 5b–f. As shown in Fig. 5a, the bonding interfaces are nearly straight and the groove of 316 L is filled with the CuCrZr, which indicated that precise meshing collision was accomplished between the flyer and base plate. Fig. 5b–f show that the metallurgical bonding of CuCrZr alloy to 316 L stainless steel was created along the dovetail interface, and two types of bond are observed at the interface, namely metal/metal and metal/solidified melt. As demonstrated in Fig. 5b and c, both the lower and upper corners of the meshing interface are rounded corners after EW, while they were sharp corners before EW, which indicated that large plastic deformations of welding materials emerged in the process of bonding interface formation. However, no defects such as cracks and voids were found, which shows that excellent bonding was created in the two regions. Fig. 5f suggests that the upper interfaces was welded mainly in the way of direct bonding, and only some small localized melting zones were observed at the interfaces. This result was in consistent with the previous studies [12,14] about the copper/stainless steel interface obtained by EW, and the reason for this result was that the welding
Fig. 5. Morphology of an interface perpendicular to detonation direction obtained by OM, (a) general layout of meshing interface after EW, (b) high resolution image (HRI) of position B in (a), (c) HRI of position C in (a), (d) HRI of position D in (a), (e) HRI of position E in (a), (f) HRI of position F in (a). 109
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 6. Morphology of bonding interface at the initial detonation zone, (a) general layout of an innclined interface, (b) HRI in position B in (a).
tensile strength of the bimetallic clad plate is higher than that of CuCrZr whose tensile strength is approximately 400 MPa, which is consistent with the results of previous researches, namely the tensile strength of bimetallic clad plate is higher than that of the metal plate with the relatively small tensile strength in the flyer and base plates [19,31,32]. Fig. 12 shows the fractography of the bimetallic clad plate after tensile test. Numerous ductile dimples with large size are observed on both the CuCrZr side and the 316 L side away from the interface, as shown in Fig. 12b and d, which are the typical characteristics of ductile fracture. Fig. 12a shows the unambiguous characteristics of cleavage fracture without any dimples on the CuCrZr side near the interfaces. It was reported that brittle fracture in ductile materials could be observed near the bonding interface due to high degree of shock hardening [7,33]. In this condition, owing to the low ductility of this region containing some brittle intermetallic phases and severely deformed grains, some cracks were initiated at the interface and propagated along the tensile direction [28]. Fig. 12d exhibited unambiguous characteristics of ductile fracture with dimples on the 316 L side near the interfaces, however the sizes and numbers of the dimples are lower than those away from the interface, which indicates that the ductility of the 316 L at the interface was also reduced.
in the formation of thicker melting layer. For further revealing the composition of the melting zone, element spot scans were carried out, and corresponding results are shown in Fig. 8. According to Fig. 8, the melting layer is composed of 68.15 at.% Cu, 24.39 at.% Fe and 7.46 at. % Cr. The main component of the melting layer is Cu, which can be attributed to that CuCrZr has higher thermal conductivity and lower melting point than 316 L. Figs. 9a and b show the microstructural morphology of an upper interface and a lower interface parallel to detonation direction. As illustrated in Fig. 9a and b, both the upper and lower bonding interfaces of CuCrZr/316 L bimetallic sheets showed a wavy morphology. The formation of the waves can be attributed to the results of variations in the velocity distribution at collision point and periodic disturbances of materials [15]. For the copper/stainless steel bonding interface manufactured by EW, both straight and wavy bonding interfaces have been obtained by researchers [11–14]. The wavy bonding interface is usually preferred due to better mechanical properties and more bonding area [12,14]. It was also reported that the bonding interface transformed from straight to wavy interface with the increasing explosive loading [11,29]. The morphology of the interface given in Fig. 9a and b indicates that the explosive loading was high enough to obtain a wavy interface, Fig.9a and b also show that drastic grain refinement occurs at the interface and grains are elongated in the direction of the explosion. The is due to the tangential component of impact velocity and the shearing stresses produced between the impacting metals, which induces excessive plastic deformation at the interface during explosion [28,30]. As shown in the SEM image using higher magnification (see Fig. 10), there is no defects such as micropores and cracks at the interface, which indicates that the welding parameters were reasonable and the bonding interfaces were excellent.
3.3. Tensile shear tests Shear test is important to assess transition joints and one of the purposes of this study is to improved the shear strength of the bonding interface. For the meshing interface, the tensile shear resistance may vary in different tensile directions. Therefore, we performed tensile shear tests on the meshing interface in two tensile directions. In addition, the tensile shear tests of the ordinary interface were also carried out to evaluate the effect of meshing interface on shear strength. For comparison, the ordinary interface was obtained by using the same welding parameters with the meshing interface. Fig. 13 shows typical stress–distance curves obtained from the tensile shear tests, and the shear strength values for each group of tests are listed in Table 7.
3.2. Tensile tests The typical stress–strain curves obtained by tensile tests are shown in Fig. 11, and the average tensile strength is 598 MPa. Obviously, the
Fig. 7. Distribution of elements across the bonding interface, (a) distribution of alloy elements (DAE) along upper interface, (b) DAE along lower interface, (c) DAE along inclined interface. 110
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 8. Element spot scans results of local melting zone, (a) target areas of EDS analysis, (b) EDS analysis result (EAR) for rectangle 1 in.(a), (c) EAR for rectangle 2 in. (a), and (d) EAR for rectangle 3 in.(a).
Fig. 9. Morphology of bonding interface parallel to detonation direction obtained by OM, (a) upper inerface, (b) lower interface.
Fig. 10. Morphology of bonding interface obtained by SEM, (a) typical wave (b) high resolution image in (a).
indicates that the shear strength of the meshing interface may be increased further by parameters optimization. The average tensile-shear strength of 328 MPa is obtained for meshing interface 90°, which is 9% higher than that of ordinary interface. This happened because mechanical locking improves the shear strength. As shown in Fig. 13, the meshing interface can prevent crack initiation and propagation under external force along the meshing interface 90° direction. Meanwhile, this mechanical locking may also help to prevent interface cracking when it is subjected to thermal stress caused by high temperature fluid or possible butt welding process [4,35]. The tensile-shear strength values of meshing interface in two direction are both higher than that of ordinary interface, which indicates that the meshing interface effectively improved the shear strength of the bonding interface.
According to Table 7, the average tensile-shear strength of the meshing interface 0° was 359 MPa, which was approximately 19% higher than the 301 MPa of ordinary interface. The reson for the increase is that the bonding area of the meshing interface was 31% higher than that of ordinary interface. As a result, higher macroscopic shear strength was obtained. However, the shear strength of the unit bonding area was lower than that of ordinary interface, which could be attributed to that the bonding strength of the lower and inclined interfaces was lower than that of the upper interfaces. This is because continuous melting layers were formed at the lower and inclined interface (See Fig. 5). It was reported that a hard and brittle intermetallic was often formed in the melting zones, which affected the bonding quality and the mechanical properties with a negative manner [12,34]. This result also 111
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 13. Typical stress–distance curves obtained by tensile shear tests.
Fig. 11. Typical stress–strain curves obtained by tensile tests.
Table 7 Tensile shear test results of three bonding interfaces.
3.4. Microhardness tests Fig. 14a and b respectively show the test positions and results of the microhardness near the interface. As shown in Fig. 14b, the microhardness generally decreased as the distance from the welding interface increased, which could be attributed to work hardening due to the severe plastic deformation in the welding zones. The result is in consistent with the previous work [12–14,19]. However, the increased value of the hardness for 316 L was more than the CuCrZr. This phenomenon might be explained by the fact that 316 L has higher strain hardening with respect to the CuCrZr [12]. In addition, Fig. 14b shows that the microhardness values at the upper interface are higher than those at lower interface. This is attributed to the air in the groove which affects the collision speed with a negative manner. For this reson, the cold deformation at the upper interface was greater than that at lower interface. The microhardness values at the inclined interface were significantly lower than those at the upper and lower interfaces. This happened because direct collision between CuCrZr and 316 L occurred
No
στo/MPa
στm0/MPa
στm90/MPa
1 2 3
305 292 306
350 365 361
316 330 339
Note: στo is the tensile shear strength of ordinary interface, στm0 is the tensile shear strength of meshing interface 0° στm90 is the tensile shear strength of meshing interface 90°.
at the upper and lower interfaces during EW process, while oblique collision and slide occurred at the inclined interface. So the degree of work hardening at the inclined interface was lower than that at upper and lower interfaces. Fig. 15 shows microhardness test results near the melt layer. The average microhardness of the melt layer at the bonding interface was 191.7 HV, which was slightly higher than the 162.0 HV of the CuCrZr, and much less than 346.7 HV of the 316 L near the melt
Fig. 12. Fractography after tensile test, (a) the fractography of CuCrZr near the interface, (b) the fractography of CuCrZr away from the interface, (c) the fractography of 316 L near the interface, (d) the fractography of 316 L away from the interface. 112
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 14. Microhardness profiles near the bonding interfaces.
interface increased. The upper interface had the highest microhardness value, and the microhardness values at the inclined interface were significantly lower than those at the upper and lower interfaces. Acknowledgements The reported research is supported by the China National Nature Science (no. 51674229 and no. 51874267), and Fundamental Research Funds for Central Universities (WK2480000002). References [1] D.W. Lee, Y.D. Bae, S.K. Kim, et al., Design evaluation of the semi-prototype for the ITER blanket first wall qualification, Thin Solid Films 518 (2010) 6676–6681. [2] I. Mazul, A. Alekseev, V. Belyakov, et al., Russian development of enhanced heat flux technologies for ITER first wall, Fusion Eng. Des. 87 (2012) 437–442. [3] S.H. Goods, J.D. Puskar, Solid state bonding of CuCrZr to 316L stainless steelfor ITER applications, Fusion Eng. Des. 86 (2011) 1634–1638. [4] R. Wei, S.X. Zhao, H. Dong, et al., Enhancing the cucrzr/316l hip-joint by Ni electroplating, Fusion Eng. Des. 117 (2017) 58–62. [5] K. Tsuchiya, H. Kawamura, Mechanical properties of Cu-Cr-Zr alloy and SS316 joints fabricated by friction welding method, J. Nucl. Mater. 233–237 (233) (1996) 913–917. [6] V. Shokri, A. Sadeghi, M.H. Sadeghi, Effect of friction stir welding parameters on microstructure and mechanical properties of DSS–Cu joints, Mater. Sci. Eng. A. 693 (2017) 111–120. [7] Y. Zhang, J. Huang, H. Chi, et al., Study on welding–brazing of copper and stainless steel using tungsten/metal gas suspended arc welding, Mater. Lett. 156 (2015) 7–9. [8] J.T. Xiong, Q. Xie, J.L. Li, et al., Diffusion bonding of stainless steel to copper with Tin bronze and Gold interlayers, J. Mater. Eng. Perform. 21 (1) (2012) 33–37. [9] Khalid A. Al-Ghamdi, G. Hussain, On the comparison of formability of roll-bonded steel-Cu composite sheet metal in incremental forming and stamping processes, Int. J. Adv. Manuf. Tech. 87 (1-4) (2016) 1–12. [10] R. Wei, Q. Li, W.J. Wang, et al., Microstructure and properties of W-Cu/CuCrZr/ 316L joint bonded by one-step HIP technique, Fusion Eng. Des. 128 (2018) 47–52. [11] Ahmet Durgutlu, B. Gülenç, F. Findik, Examination of copper/stainless steel joints formed by explosive welding, Mater. Des. 26 (6) (2005) 497–507. [12] Bina Mohammad Hosein, F. Dehghani, M. Salimi, Effect of heat treatment on bonding interface in explosive welded copper/stainless steel, Mater. Des. 45 (2013) 504–509. [13] Ahmet Durgutlu, H. Okuyucu, B. Gulenc, Investigation of effect of the stand-off distance on interface characteristics of explosively welded copper and stainless steel, Mater. Des. 29 (7) (2008) 1480–1484. [14] R. Ma, Y. Wang, J. Wu, et al., Explosive welding method for manufacturing ITERgrade 316L(N)/CuCrZr hollow structural member, Fusion Eng. Des. 89 (12) (2014) 3117–3124. [15] Mousavi, S.A.A. Akbari, P.F. Sartangi, Experimental investigation of explosive welding of cp-titanium/AISI 304 stainless steel, Mater. Des. 30 (3) (2009) 459–468. [16] R. Mendes, J.B. Ribeiro, A. Loureiro, Effect of explosive characteristics on the explosive welding of stainless steel to carbon steel in cylindrical configuration, Mater. Des. 51 (51) (2013) 182–192. [17] P. Bazarnik, B. Adamczyk-Cieślak, A. Gałka, et al., Mechanical and microstructural characteristics of Ti6Al4V/AA2519 and Ti6Al4V/AA1050/AA2519 laminates manufactured by explosive welding, Mater. Des. 111 (2016) 146–157. [18] P. Wang, J. Chen, Q. Li, et al., Study on the microstructure and properties evolution of CuCrZr/316LN-IG explosion bonding for ITER first wall components, Fusion Eng. Des 124 (2017) 1135–1139. [19] Xuejiao Li, H. Ma, Z. Shen, Research on explosive welding of aluminum alloy to steel with dovetail grooves, Mater. Des. 87 (2015) 815–824.
Fig. 15. Microhardness near the melt layer.
layer. This is because the melt layer is mainly composed of CuCrZr, as shown in Fig. 8. 4. Conclusions In this study, CuCrZr/316 L meshing bonding interfaces were obtained though EW and their microstructural and mechanical properties were investigated. The following conclusions can be drawn from this study: (1) Meshing bonding interfaces can be obtained though EW, which is a new way to enhance CuCrZr/316 L EW-joints by increasing bonding area and producing mechanical locking. (2) Due to careful control of welding parameters, excellent metallurgical bonding of CuCrZr alloy and 316 L stainless steel was achieved along the dovetail interface. The upper interfaces was welded mainly in the way of direct bonding, and continuous melting layer appeared at the lower and inclined interfaces. The EDS analysis indicated that the melting zone mainly consisted of Cu. (3) The ideal wave bonding was created at the upper and lower interface parallel to detonation direction, and no defects such as micropores and cracks was found in these regions. (4) The tensile fracture specimen was characterized by cleavage fractures on the CuCrZr side and ductile fractures on the 316 L side near the interface. (5) The average shear strength of the meshing interface 0° and 90° was increased by 19% and 9%, when compared to the ordinary CuCrZr/ 316 L EW-joints. This indicates that the meshing interface can effectively improve the shear strength of the bonding interface. (6) The values of microhardness decreased as the distance from the 113
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
[20] S.W. Stivers, R.H. Wittman, Computer selection of the optimum explosive loading andwelding geometry, High energy rate fabrication, Proc. Semin., of the 5nd Int. Symp. 4 (2) (1975) 1–16. [21] T.Z. Blazynski, Explosive Welding Forming and Compaction. London, (1983) ISBN 0-85334-166-4. [22] G. Cowan, O. Bergmann, A. Holtzman, Mechanism of bond zone wave formation in explosion-clad metals, Metall. Mater. Trans. B. 2 (11) (1971) 3145–3155. [23] P.V. Vaidyanathan, A. Ramanathan, Computer-aided design of explosive welding systems, J. Mater. Process. Tech 38 (3) (1993) 501–516. [24] M.M. Athar, B. Hoseini, Tolaminejad, weldability window and the effect of interface morphology on the properties of Al/Cu/Al laminated composites fabricated by explosive welding, Mater. Des. 86 (2015) 516–525. [25] A. Koch, N. Arnold, M.A. Estermann, Simple relation between the detonation velocity of an explosive and its gurney energy, Propellants Explos. Pyrotech. 27 (6) (2002) 365–368. [26] S.W. Stivers, R.H. Wittman, Computer selection of the optimum explosive loading and welding geometry, Proc 5th Int on High Energy Rate Fabrication. 4 (2) (1975) 1–16. [27] GB/T6396-2008, Clad steel plates-mechanical and technological test, China National Standardization Management Committee (2008). [28] M.M. Hoseini-Athar, B. Tolaminejad, Interface morphology and mechanical
[29] [30]
[31]
[32] [33] [34]
[35]
114
properties of Al-Cu-Al laminated composites fabricated by explosive welding and subsequent rolling process, Met. Mater. Int. 22 (2016) 670–680. F. Fındık, Recent developments in explosive welding, Mater. Des. 32 (3) (2011) 1081–1093. J.E. Lee, D.H. Bae, W.S. Chung, K.H. Kim, J.H. Lee, Y.R. Cho, Effects of annealing on the mechanical and interface properties of stainless steel/aluminum/copper cladmetal sheets, J. Mater. Process. Technol. 187–188 (2007) 546–549. M. Acarer, B. Demir, An investigation of mechanical and metallurgical properties of explosive welded aluminum–dual phase steel, Mater. Lett. 62 (25) (2008) 4158–4160. H. Xia, S. Wang, H. Ben, Microstructure and mechanical properties of Ti/Al explosive cladding, Mater. Des. 56 (2014) 1014–1019. R. Kacar, M. Acarer, Microstructure–property relationship in explosively welded duplex stainless steel–steel, Mater. Sci. Eng. A 363 (2003) 290–296. S.M. Aceves, F. Espinosa-Loza, J.W. Elmer, et al., Comparison of Cu, Ti and Ta interlayer explosively fabricated aluminum to stainless steel transition joints for cryogenic pressurized hydrogen storage, Int. J. Hydrog. Energy 40 (2015) 1490–1503. J. Ning, L.J. Zhang, G.C. Jiang, et al., Narrow gap multi-pass laser butt welding of explosion welded CP-Ti/Q235B bimetallic sheet by using a copper interlayer, J. Alloys Compd. 701 (2017) 587–602.
Contents lists available at ScienceDirect
Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Study on explosive welding for manufacturing meshing bonding interface of CuCrZr to 316L stainless steel
T
⁎
Ming Yanga, Honghao Maa,b, , Zhaowu Shena, Yuling Suna,c a
CAS Key Laboratory of Mechanical Behavior and Design of Materials, Department of Modern Mechanics, University of Science and Technology of China, Hefei, Anhui 230027, China b State Key Laboratory of Fire Science, University of Science and Technology of China, Hefei, Anhui 230026, China c PLA Army Academy of Artillery and Air Defense, Hefei, Anhui 230031, China
A R T I C LE I N FO
A B S T R A C T
Keywords: Explosive welding Meshing interface Mechanical properties Microstructure
In order to improve the CuCrZr/316L transition joints manufactured by explosive welding (EW), a new EW method was developed to obtain meshing bonding interfaces. The welding parameters adopted in the experiment were close to the lower limit, and the bonding quality was evaluated by systematically microstructure and mechanical tests. The results showed that the meshing interfaces can effectively enhance CuCrZr/316L EW-joints by increasing bonding area. The microstructure results indicated that metallurgical bonding of CuCrZr alloy to 316L stainless steel was created along the meshing interfaces, and two types of bond were observed at the bonding interfaces, namely metal/metal and metal/solidified melt. The EDS analysis indicated that the melt mainly consisted of Cu. Fractography studies on tensile specimens showed cleavage fracture on the CuCrZr side and ductile fracture on the 316L side near the interfaces. The tensile shear test results showed that the shear strength of the meshing interface 0° and 90° was increased by 19% and 9%, when compared to the ordinary CuCrZr/316L EW-joints. The results of microhardness test revealed that the values of microhardness decreased as the distance from the interface increased.
1. Introduction As a component directly faces the plasma, the first wall (FW) of ITER will bear extremely high heat fluxes about 20–30 MW/m2 [1,2]. CuCrZr alloy was chosen as the heat sink material due to its excellent thermal conductivity, which was bonded to an austenitic stainless steel (316 L) support structure. Recent developments in the design of the FW required that the CuCrZr/316 L joints directly faced the internal grooves, which formed the pressure boundary for the cooling water [3]. Thus renewed attention is being given to manufacturing excellent CuCrZr/316 L joints which will withstand the internal pressure, the fatigue loads and remain leak tight during the cyclic mode of operation [4]. Several methods such as friction welding [5,6], brazing [7], diffusion bonding [8], roll bonding [9], hot isostatically pressing (HIP) [10], and explosive welding (EW) [3,11–14] have been investigated for joining copper alloy to steel. When compared to other welding methods, EW shows many advantages such as high bonding strength and no long range interdiffusion [3,14]. Thus this technology is attracting more and more attention from ITER researchers [3,14]. EW is a
⁎
well-known solid state method for joining various metal materials, which employs explosive energy to make metals produce a high-speed oblique collision and achieve metallurgical bond [15]. Since external heating and large-scale melt are unnecessary for this process, it is capable of keeping the physical and chemical properties of the wrought parent components unchanged [16]. Moreover, because it enables to obtain metallurgical bonding on over the entire junction surfaces without any oxidation, explosive welding joints generally have relatively high strength [17]. Therefore, EW is very suitable for welding such CuCrZr/316 L dissimilar bulk metals. Several studies on manufacturing CuCrZr/316 L bimetals have been conducted using conventional EW technology. Goods et al [3] obtained the CuCrZr/316 L joints using two different solid state techniques, namely HIP and EW, and they found that EW could reduce or eliminate HIP’s shortcomings such as excessive grain growth, overaging and further loss of critical properties. Ma et al [14] provided a new EW method for manufacturing ITER-grade 316 L/CuCrZr hollow structural member, and the results showed that the welded specimens were sound. Wang et al [18] investigated the microstructure and mechanical properties evolution of post-weld heat treatment and final Be/CuCrZr HIP on the interface of
Corresponding author. E-mail address: [email protected] (H. Ma).
https://doi.org/10.1016/j.fusengdes.2019.03.137 Received 20 June 2018; Received in revised form 10 December 2018; Accepted 21 March 2019 0920-3796/ © 2019 Published by Elsevier B.V.
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Pmin is minimum collision pressure determined by the formula below [21].
CuCrZr/316 L EW-plates. The results indicated that EW was qualified for the manufacturing of ITER FW components in China. Although EW shows many advantages when compared to other welding methods, the applying CuCrZr/316 L EW-joints on ITER still faces a lot of challenges due to the poor working condition. During the cyclic mode of operation, the welding joints are directly exposed to the cooling water, which will be continuously impacted by high pressure fluid and become the weakest part of a component. As a result, the ordinary CuCrZr/316 L EW-joints may fail in service due to the low bonding strength. Thus, the goal of this work is to increase the bonding strength of CuCrZr/316 L EW-joints. For this purpose, we explored a new EW technology for manufacturing CuCrZr/316 L meshing bonding interface, and the microstructure and mechanical properties of the bonding interfaces were systematically investigated.
pmin =
1 c (ρσb )1/2 2
(2)
Where c, ρ, and σb are the sonic speed, density and tensile strength. The minimum collision pressure Pmin1 of the flyer plate and Pmin2 of the base plate can be calculated respectively through Eq. (2), and the Pmin is the bigger value of Pmin1 and Pmin2. There is a lower limit of collision point velocity vcmin for predicting the laminar–turbulent flow transition of the metal materials to be welded. This limit also defines the transition from straight to wavy interfaces between the flyer and base plates. Wavy interfaces can be obtained when the collision point velocity is larger than the lower limit. Cowan et al. [22] defined lower limit of collision point velocity with hydrodynamic analogy.
2. Experimental procedure
1/2
2R (H + H2) ⎤ vc min = ⎡ e 1 ⎢ ρ1 + ρ2 ⎥ ⎦ ⎣
2.1. Experimental materials
(3)
In this work, CuCrZr alloy and 316 L stainless steel plate were employed as flyer and base plates, which were designed with dimensions of 300 mm × 150 mm × 8 mm and 300 mm × 150 mm × 15 mm. The chemical composition of the CuCrZr and 316 L are shown in Table 1 and Table 2. It is important to accurately control the height and uniformity of explosive in this study. And to do that, honeycomb structure explosive was chosen as explosive material, which was consisted of aluminium honeycomb filled with emulsion explosive. As shown in Fig. 1, the thickness of aluminium foil is 60 μm, the side length of the regular hexagon cell is 8 mm. The components of the emulsion matrix are presented in Table 3. Hollow glass microballoons with a mass fraction of 15% were chosen as the sensitizer. The detonation velocity of the emulsion explosive is about 3000 m/s
Where Re is the critical Reynolds Number taking values of 10.6 [23], H1 and H2 are the Vickers hardness numbers of the flyer and base plates respectively. The material parameters used for calculating the lower limit of the dynamic parameters are listed in Table 4. For parallel set-up geometry, the collision point velocity vc is equal to the detonation velocity vd. The impact velocity vp could be calculated by the following expressions [24,25].
2.2. Welding parameters selection
R=
Good welding quality depends on careful control of the dynamic parameters [12]. There are three mainly dynamic parameters for EW, i.e., collision angle β, collision point velocity vc and impact velocity vp, which are determined by initial parameters such as detonation velocity and mass ratio of explosive to flyer plate. However, the three dynamic parameters are related to each other for certain geometric relationships and only two of them are independent. In this work, the collision point velocity vc and impact velocity vp were used to determine the initial parameters. The initial parameters were designed to meet the condition that the welding dynamic parameters were close to the lower limit of the dynamic parameters. In this condition, good welding quality was usually expected to be obtained [14,19]. In order to make impact pressure at the collision point exceed the yield stress of the materials to promote plastic deformation, the impact velocity vp of flyer plates must be larger than the lower limit of impact velocity vpmin. For welding of dissimilar metallic materials, the lower limit of impact velocity can be calculated by the following formula [20].
Where R is the mass ratio of explosive to flyer plate and E is Gurney energy of explosive. γ is the polytropic exponent of detonation products, the value for emulsion explosive is equal to 2.5. ρ0 and ρ1 are the density of the explosive and flyer plate, t0 and t1 are the thickness of the explosive and flyer plate. The empirical formula of the stand-off distance h between the flyer plate and the base plate was expressed as [26].
1 1 ⎞ vp min = pmin ⎛⎜ + ⎟ c2 ρ2 ⎠ ⎝ c1 ρ1
vp =
Mg
Al
Si
P
Fe
Cu
Content
0.8
0.1-0.8
0.1-0.25
0.1-0.25
0.5
0.1
0.5
Bal.
(4)
1 ⎛ γ ⎞ 2 ⎜ ⎟ v d − 1 ⎝ γ + 1⎠
ρ0 t0 ρ1 t1
(5)
(6)
(7)
Where t0 is the explosive thickness, and t1 is the flyer thickness. The initial parameters used for this EW are shown in Table 5 (calculation according to aforementioned procedure) and the accordingly dynamic parameters and the lower limit of the dynamic parameters are listed in Table 6. 2.3. Experimental methods The meshing bonding interfaces were obtained by prefabricating grooves in flyer and base plates. In order to make bonding quality excellent, there are two key points for manufacturing the grooves. First, the grooves will not prevent the jet and air between the flyer plate and base plate from being out. Second, the groove surfaces of base plate should be in good contact with the flyer plate during collision process. Considering the two points, dovetail grooves were cut along the detonation direction in this study. The dovetail grooves of the flyer plate were the same as those of the base plate, so which could be tightly assembled together. The dimensions of the dovetail grooves are shown in Fig. 2. The success of the experiment depends on accurately controling of the flight path of the flyer plate. For this end, four locating pins were
Table 1 Chemical composition of CuCrZr (wt%). Zr
γ2
h = 0.2(t0 + t1)
Where c1 and c2 are the sonic speed of the flyer plate and base plate respectively, ρ1 and ρ2 are the density of the flyer plate and base plate.
Cr
3R2 R2 + 5R + 4 γ
E=
(1)
Element
2E
107
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Table 2 Chemical composition of 316 L stainless steel (wt%). Element
C
Si
Mn
Cr
Ni
S
P
Mo
Fe
Content
≤0.08
≤1.00
≤2.00
16.00-18.00
10.00-14.00
≤0.03
≤0.05
≤2.00-3.00
Bal.
Fig. 2. Dimensions of dovetail grooves of the flyer and base plate (unit: mm).
Fig. 1. Honeycomb structure explosive. Table 3 Component of the emulsion matrix. Component
NH4NO3
NaNO3
H2O
C18H38
C24H44O6
C12H26
mass fraction
75%
10%
8%
4%
2%
1%
Fig. 3. Schematic diagram of the meshing interface explosive welding. Table 4 Selected material parameters used for calculating the lower limit of the dynamic parameters. Material
Ρ [kg•m−3]
Hv [Mpa]
C [m•s−1]
σb [Mpa]
CuCrZr 316 L
8960 7980
1200 2000
3800 5000
380 520
In order to reveal the interface morphology, specimens were cut parallel and vertical to the detonation direction respectively, and the cross-section of the specimens were ground with emery papers up to No. 5000 and polished to 0.5 um by diamond paste. A scanning electron microscope (GeminiSEM 500) and an optical microscope (Leica DM4M) were employed for microstructure observation of the bonding interfaces. Energy-dispersive X-ray spectrometry (EDS) analysis was also carried out to characterize the distribution of the alloy elements across the bonding interface. In order to investigate the mechanical properties of the welded plate, microhardness, tensile, and tensile-shear tests were carried out. Microhardness tests were conducted on a microhardness machine (HVS1000 M) using a 100 g load for 10 s. The tensile tests were carried out with a tensile strain rate of 1 × 10−4/s according to GB/T6396-2008 [27], and the dimensions of the specimens are shown in Fig. 4a. Fractography studies on broken specimens from tensile tests were carried out using a scanning electron microscope. According to standard test method, in order to ensure that bond separation occurs sooner than bulk material failure during tensile shear tests, the maximum permissible length of overlapped portion (L) was determined by the following equation [28]:
Table 5 Initial parameters used for explosive welding. t0 (mm)
t1 (mm)
ρ0 (g/cm3)
R
h (mm)
vd (m/s)
25
7
0.9
0.36
7
3000
Note: The original thickness of flyer plate was 8 mm, however the value of 7 mm was taken for considering the effect of grooves to mass. The calculated h is 6.4 mm, however the actual stand-off distance of 7 mm was used to ensure that the flyer would be accelerated to the maximum velocity. Table 6 Dynamic parameters for explosive welding. vpmin (m/s)
vcmin (m/s)
vp (m/s)
vc (m/s)
278
2007
311
3000
L= employed to guarantee precise meshing collision between the flyer and base plate, as shown in Fig. 3. The parallel set-up geometry was employed for this EW. A detonator was placed on the side edge of the explosive to make the detonation along the dovetail groove direction. The welding assembly was placed on a steel anvil in an explosion vessel.
σb T τ
(8)
Where σb is the yield point of softer material, T is the thickness and τ should be considered as 1.5 times higher than the estimated average shear bonding strength [28]. This equation ensures that fracture position is at bonding interface, and the dimensions of the specimens are shown in Fig. 4b. 108
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
parameters were close to the lower limit. If high explosive ratio was used, ejection will be formed in between flyer and base plate after the explosion. This results in that melting area in interface and also possible oxidation and dirty surface will not be exported to outside [11,12]. However, Fig. 5d and e show that the lower and inclined interfaces were connected by continuous melting layer, and the thickness of the melting layer was in the range of 50–200 μm. Formation of the melting layer was attributed to the adiabatic heating of gases compressed in the dovetail groove. The air inside the groove is harder to be discharged than that near the upper interface, which resulted in that the air near the lower and inclined interfaces was subjected to adiabatic compression during the collision process. Thus the continuous melting layer at the interfaces was formed under the actions of adiabatic compression heat. One way to reduce the melting layer is to reduce the detonation velocity. According to our previous experiments, the detonation velocity in the initial detonation zone is about 2600 m/s due to the unstable detonation, while it is about 3000 m/s in other zone. Fig. 6 shows the morphology of bonding interface at the initial detonation zone. It can be seen from Fig. 6 that the melting layer is extremely thin (about 8 μm) at the initial detonation zone, which shows that the melt layer can be effectively reduced by decreasing the detonation velocity. Due to the fact that the collision point velocity is equal to the detonation velocity, reducing the detonation velocity is conducive to air emission at the interface so as to reduce the melting layer. In order to describe the transition layer between the two welded materials, element line scans were conducted to determine elements distribution across the bonding interface. As shown in Table 1 and Table 2, the content of Zr in CuCrZr and 316 L is extremely low, so the distribution of Zr across the bonding interface is not presented in Figs. 7 and 8. It can be observed from Fig. 7 that the element distribution line was steep across the upper interface, while a platform appeared across the lower and inclined interfaces, which proved the results obtained by OM, namely that the upper interface was bonded in the form of direct bonding, while the lower and inclined interfaces were joined by melting layer. In addition, it can be observed from Fig. 7b and c that the diffusion layer’s thickness of the inclined interface was greater than that of the lower interface and the elements variation was more intense. This can be attributed to adiabatic shear along the inclined interface. Unlike the upper and lower interfaces, the inclined interface was bonded by inclined collision between flyer and base plates, thus stronger adiabatic shear was expected to occur along the inclined interface, which resulted
Fig. 4. Dimensions of the tensile (a) and tensile shear (b) specimens (unit: mm).
3. Results and discussion 3.1. Microstructure of bonding interfaces Figs. 5 shows optical microscope (OM) images of the microstructural morphology of bonding interface perpendicular to detonation direction. Fig. 5a is the general layout of a meshing interface after EW, which also shows five typical positions (i.e., the rectangular areas b, c, d, e, and f in Fig. 5a), and corresponding high magnification micrographs around these five positions are shown in Fig. 5b–f. As shown in Fig. 5a, the bonding interfaces are nearly straight and the groove of 316 L is filled with the CuCrZr, which indicated that precise meshing collision was accomplished between the flyer and base plate. Fig. 5b–f show that the metallurgical bonding of CuCrZr alloy to 316 L stainless steel was created along the dovetail interface, and two types of bond are observed at the interface, namely metal/metal and metal/solidified melt. As demonstrated in Fig. 5b and c, both the lower and upper corners of the meshing interface are rounded corners after EW, while they were sharp corners before EW, which indicated that large plastic deformations of welding materials emerged in the process of bonding interface formation. However, no defects such as cracks and voids were found, which shows that excellent bonding was created in the two regions. Fig. 5f suggests that the upper interfaces was welded mainly in the way of direct bonding, and only some small localized melting zones were observed at the interfaces. This result was in consistent with the previous studies [12,14] about the copper/stainless steel interface obtained by EW, and the reason for this result was that the welding
Fig. 5. Morphology of an interface perpendicular to detonation direction obtained by OM, (a) general layout of meshing interface after EW, (b) high resolution image (HRI) of position B in (a), (c) HRI of position C in (a), (d) HRI of position D in (a), (e) HRI of position E in (a), (f) HRI of position F in (a). 109
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 6. Morphology of bonding interface at the initial detonation zone, (a) general layout of an innclined interface, (b) HRI in position B in (a).
tensile strength of the bimetallic clad plate is higher than that of CuCrZr whose tensile strength is approximately 400 MPa, which is consistent with the results of previous researches, namely the tensile strength of bimetallic clad plate is higher than that of the metal plate with the relatively small tensile strength in the flyer and base plates [19,31,32]. Fig. 12 shows the fractography of the bimetallic clad plate after tensile test. Numerous ductile dimples with large size are observed on both the CuCrZr side and the 316 L side away from the interface, as shown in Fig. 12b and d, which are the typical characteristics of ductile fracture. Fig. 12a shows the unambiguous characteristics of cleavage fracture without any dimples on the CuCrZr side near the interfaces. It was reported that brittle fracture in ductile materials could be observed near the bonding interface due to high degree of shock hardening [7,33]. In this condition, owing to the low ductility of this region containing some brittle intermetallic phases and severely deformed grains, some cracks were initiated at the interface and propagated along the tensile direction [28]. Fig. 12d exhibited unambiguous characteristics of ductile fracture with dimples on the 316 L side near the interfaces, however the sizes and numbers of the dimples are lower than those away from the interface, which indicates that the ductility of the 316 L at the interface was also reduced.
in the formation of thicker melting layer. For further revealing the composition of the melting zone, element spot scans were carried out, and corresponding results are shown in Fig. 8. According to Fig. 8, the melting layer is composed of 68.15 at.% Cu, 24.39 at.% Fe and 7.46 at. % Cr. The main component of the melting layer is Cu, which can be attributed to that CuCrZr has higher thermal conductivity and lower melting point than 316 L. Figs. 9a and b show the microstructural morphology of an upper interface and a lower interface parallel to detonation direction. As illustrated in Fig. 9a and b, both the upper and lower bonding interfaces of CuCrZr/316 L bimetallic sheets showed a wavy morphology. The formation of the waves can be attributed to the results of variations in the velocity distribution at collision point and periodic disturbances of materials [15]. For the copper/stainless steel bonding interface manufactured by EW, both straight and wavy bonding interfaces have been obtained by researchers [11–14]. The wavy bonding interface is usually preferred due to better mechanical properties and more bonding area [12,14]. It was also reported that the bonding interface transformed from straight to wavy interface with the increasing explosive loading [11,29]. The morphology of the interface given in Fig. 9a and b indicates that the explosive loading was high enough to obtain a wavy interface, Fig.9a and b also show that drastic grain refinement occurs at the interface and grains are elongated in the direction of the explosion. The is due to the tangential component of impact velocity and the shearing stresses produced between the impacting metals, which induces excessive plastic deformation at the interface during explosion [28,30]. As shown in the SEM image using higher magnification (see Fig. 10), there is no defects such as micropores and cracks at the interface, which indicates that the welding parameters were reasonable and the bonding interfaces were excellent.
3.3. Tensile shear tests Shear test is important to assess transition joints and one of the purposes of this study is to improved the shear strength of the bonding interface. For the meshing interface, the tensile shear resistance may vary in different tensile directions. Therefore, we performed tensile shear tests on the meshing interface in two tensile directions. In addition, the tensile shear tests of the ordinary interface were also carried out to evaluate the effect of meshing interface on shear strength. For comparison, the ordinary interface was obtained by using the same welding parameters with the meshing interface. Fig. 13 shows typical stress–distance curves obtained from the tensile shear tests, and the shear strength values for each group of tests are listed in Table 7.
3.2. Tensile tests The typical stress–strain curves obtained by tensile tests are shown in Fig. 11, and the average tensile strength is 598 MPa. Obviously, the
Fig. 7. Distribution of elements across the bonding interface, (a) distribution of alloy elements (DAE) along upper interface, (b) DAE along lower interface, (c) DAE along inclined interface. 110
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 8. Element spot scans results of local melting zone, (a) target areas of EDS analysis, (b) EDS analysis result (EAR) for rectangle 1 in.(a), (c) EAR for rectangle 2 in. (a), and (d) EAR for rectangle 3 in.(a).
Fig. 9. Morphology of bonding interface parallel to detonation direction obtained by OM, (a) upper inerface, (b) lower interface.
Fig. 10. Morphology of bonding interface obtained by SEM, (a) typical wave (b) high resolution image in (a).
indicates that the shear strength of the meshing interface may be increased further by parameters optimization. The average tensile-shear strength of 328 MPa is obtained for meshing interface 90°, which is 9% higher than that of ordinary interface. This happened because mechanical locking improves the shear strength. As shown in Fig. 13, the meshing interface can prevent crack initiation and propagation under external force along the meshing interface 90° direction. Meanwhile, this mechanical locking may also help to prevent interface cracking when it is subjected to thermal stress caused by high temperature fluid or possible butt welding process [4,35]. The tensile-shear strength values of meshing interface in two direction are both higher than that of ordinary interface, which indicates that the meshing interface effectively improved the shear strength of the bonding interface.
According to Table 7, the average tensile-shear strength of the meshing interface 0° was 359 MPa, which was approximately 19% higher than the 301 MPa of ordinary interface. The reson for the increase is that the bonding area of the meshing interface was 31% higher than that of ordinary interface. As a result, higher macroscopic shear strength was obtained. However, the shear strength of the unit bonding area was lower than that of ordinary interface, which could be attributed to that the bonding strength of the lower and inclined interfaces was lower than that of the upper interfaces. This is because continuous melting layers were formed at the lower and inclined interface (See Fig. 5). It was reported that a hard and brittle intermetallic was often formed in the melting zones, which affected the bonding quality and the mechanical properties with a negative manner [12,34]. This result also 111
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 13. Typical stress–distance curves obtained by tensile shear tests.
Fig. 11. Typical stress–strain curves obtained by tensile tests.
Table 7 Tensile shear test results of three bonding interfaces.
3.4. Microhardness tests Fig. 14a and b respectively show the test positions and results of the microhardness near the interface. As shown in Fig. 14b, the microhardness generally decreased as the distance from the welding interface increased, which could be attributed to work hardening due to the severe plastic deformation in the welding zones. The result is in consistent with the previous work [12–14,19]. However, the increased value of the hardness for 316 L was more than the CuCrZr. This phenomenon might be explained by the fact that 316 L has higher strain hardening with respect to the CuCrZr [12]. In addition, Fig. 14b shows that the microhardness values at the upper interface are higher than those at lower interface. This is attributed to the air in the groove which affects the collision speed with a negative manner. For this reson, the cold deformation at the upper interface was greater than that at lower interface. The microhardness values at the inclined interface were significantly lower than those at the upper and lower interfaces. This happened because direct collision between CuCrZr and 316 L occurred
No
στo/MPa
στm0/MPa
στm90/MPa
1 2 3
305 292 306
350 365 361
316 330 339
Note: στo is the tensile shear strength of ordinary interface, στm0 is the tensile shear strength of meshing interface 0° στm90 is the tensile shear strength of meshing interface 90°.
at the upper and lower interfaces during EW process, while oblique collision and slide occurred at the inclined interface. So the degree of work hardening at the inclined interface was lower than that at upper and lower interfaces. Fig. 15 shows microhardness test results near the melt layer. The average microhardness of the melt layer at the bonding interface was 191.7 HV, which was slightly higher than the 162.0 HV of the CuCrZr, and much less than 346.7 HV of the 316 L near the melt
Fig. 12. Fractography after tensile test, (a) the fractography of CuCrZr near the interface, (b) the fractography of CuCrZr away from the interface, (c) the fractography of 316 L near the interface, (d) the fractography of 316 L away from the interface. 112
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
Fig. 14. Microhardness profiles near the bonding interfaces.
interface increased. The upper interface had the highest microhardness value, and the microhardness values at the inclined interface were significantly lower than those at the upper and lower interfaces. Acknowledgements The reported research is supported by the China National Nature Science (no. 51674229 and no. 51874267), and Fundamental Research Funds for Central Universities (WK2480000002). References [1] D.W. Lee, Y.D. Bae, S.K. Kim, et al., Design evaluation of the semi-prototype for the ITER blanket first wall qualification, Thin Solid Films 518 (2010) 6676–6681. [2] I. Mazul, A. Alekseev, V. Belyakov, et al., Russian development of enhanced heat flux technologies for ITER first wall, Fusion Eng. Des. 87 (2012) 437–442. [3] S.H. Goods, J.D. Puskar, Solid state bonding of CuCrZr to 316L stainless steelfor ITER applications, Fusion Eng. Des. 86 (2011) 1634–1638. [4] R. Wei, S.X. Zhao, H. Dong, et al., Enhancing the cucrzr/316l hip-joint by Ni electroplating, Fusion Eng. Des. 117 (2017) 58–62. [5] K. Tsuchiya, H. Kawamura, Mechanical properties of Cu-Cr-Zr alloy and SS316 joints fabricated by friction welding method, J. Nucl. Mater. 233–237 (233) (1996) 913–917. [6] V. Shokri, A. Sadeghi, M.H. Sadeghi, Effect of friction stir welding parameters on microstructure and mechanical properties of DSS–Cu joints, Mater. Sci. Eng. A. 693 (2017) 111–120. [7] Y. Zhang, J. Huang, H. Chi, et al., Study on welding–brazing of copper and stainless steel using tungsten/metal gas suspended arc welding, Mater. Lett. 156 (2015) 7–9. [8] J.T. Xiong, Q. Xie, J.L. Li, et al., Diffusion bonding of stainless steel to copper with Tin bronze and Gold interlayers, J. Mater. Eng. Perform. 21 (1) (2012) 33–37. [9] Khalid A. Al-Ghamdi, G. Hussain, On the comparison of formability of roll-bonded steel-Cu composite sheet metal in incremental forming and stamping processes, Int. J. Adv. Manuf. Tech. 87 (1-4) (2016) 1–12. [10] R. Wei, Q. Li, W.J. Wang, et al., Microstructure and properties of W-Cu/CuCrZr/ 316L joint bonded by one-step HIP technique, Fusion Eng. Des. 128 (2018) 47–52. [11] Ahmet Durgutlu, B. Gülenç, F. Findik, Examination of copper/stainless steel joints formed by explosive welding, Mater. Des. 26 (6) (2005) 497–507. [12] Bina Mohammad Hosein, F. Dehghani, M. Salimi, Effect of heat treatment on bonding interface in explosive welded copper/stainless steel, Mater. Des. 45 (2013) 504–509. [13] Ahmet Durgutlu, H. Okuyucu, B. Gulenc, Investigation of effect of the stand-off distance on interface characteristics of explosively welded copper and stainless steel, Mater. Des. 29 (7) (2008) 1480–1484. [14] R. Ma, Y. Wang, J. Wu, et al., Explosive welding method for manufacturing ITERgrade 316L(N)/CuCrZr hollow structural member, Fusion Eng. Des. 89 (12) (2014) 3117–3124. [15] Mousavi, S.A.A. Akbari, P.F. Sartangi, Experimental investigation of explosive welding of cp-titanium/AISI 304 stainless steel, Mater. Des. 30 (3) (2009) 459–468. [16] R. Mendes, J.B. Ribeiro, A. Loureiro, Effect of explosive characteristics on the explosive welding of stainless steel to carbon steel in cylindrical configuration, Mater. Des. 51 (51) (2013) 182–192. [17] P. Bazarnik, B. Adamczyk-Cieślak, A. Gałka, et al., Mechanical and microstructural characteristics of Ti6Al4V/AA2519 and Ti6Al4V/AA1050/AA2519 laminates manufactured by explosive welding, Mater. Des. 111 (2016) 146–157. [18] P. Wang, J. Chen, Q. Li, et al., Study on the microstructure and properties evolution of CuCrZr/316LN-IG explosion bonding for ITER first wall components, Fusion Eng. Des 124 (2017) 1135–1139. [19] Xuejiao Li, H. Ma, Z. Shen, Research on explosive welding of aluminum alloy to steel with dovetail grooves, Mater. Des. 87 (2015) 815–824.
Fig. 15. Microhardness near the melt layer.
layer. This is because the melt layer is mainly composed of CuCrZr, as shown in Fig. 8. 4. Conclusions In this study, CuCrZr/316 L meshing bonding interfaces were obtained though EW and their microstructural and mechanical properties were investigated. The following conclusions can be drawn from this study: (1) Meshing bonding interfaces can be obtained though EW, which is a new way to enhance CuCrZr/316 L EW-joints by increasing bonding area and producing mechanical locking. (2) Due to careful control of welding parameters, excellent metallurgical bonding of CuCrZr alloy and 316 L stainless steel was achieved along the dovetail interface. The upper interfaces was welded mainly in the way of direct bonding, and continuous melting layer appeared at the lower and inclined interfaces. The EDS analysis indicated that the melting zone mainly consisted of Cu. (3) The ideal wave bonding was created at the upper and lower interface parallel to detonation direction, and no defects such as micropores and cracks was found in these regions. (4) The tensile fracture specimen was characterized by cleavage fractures on the CuCrZr side and ductile fractures on the 316 L side near the interface. (5) The average shear strength of the meshing interface 0° and 90° was increased by 19% and 9%, when compared to the ordinary CuCrZr/ 316 L EW-joints. This indicates that the meshing interface can effectively improve the shear strength of the bonding interface. (6) The values of microhardness decreased as the distance from the 113
Fusion Engineering and Design 143 (2019) 106–114
M. Yang, et al.
[20] S.W. Stivers, R.H. Wittman, Computer selection of the optimum explosive loading andwelding geometry, High energy rate fabrication, Proc. Semin., of the 5nd Int. Symp. 4 (2) (1975) 1–16. [21] T.Z. Blazynski, Explosive Welding Forming and Compaction. London, (1983) ISBN 0-85334-166-4. [22] G. Cowan, O. Bergmann, A. Holtzman, Mechanism of bond zone wave formation in explosion-clad metals, Metall. Mater. Trans. B. 2 (11) (1971) 3145–3155. [23] P.V. Vaidyanathan, A. Ramanathan, Computer-aided design of explosive welding systems, J. Mater. Process. Tech 38 (3) (1993) 501–516. [24] M.M. Athar, B. Hoseini, Tolaminejad, weldability window and the effect of interface morphology on the properties of Al/Cu/Al laminated composites fabricated by explosive welding, Mater. Des. 86 (2015) 516–525. [25] A. Koch, N. Arnold, M.A. Estermann, Simple relation between the detonation velocity of an explosive and its gurney energy, Propellants Explos. Pyrotech. 27 (6) (2002) 365–368. [26] S.W. Stivers, R.H. Wittman, Computer selection of the optimum explosive loading and welding geometry, Proc 5th Int on High Energy Rate Fabrication. 4 (2) (1975) 1–16. [27] GB/T6396-2008, Clad steel plates-mechanical and technological test, China National Standardization Management Committee (2008). [28] M.M. Hoseini-Athar, B. Tolaminejad, Interface morphology and mechanical
[29] [30]
[31]
[32] [33] [34]
[35]
114
properties of Al-Cu-Al laminated composites fabricated by explosive welding and subsequent rolling process, Met. Mater. Int. 22 (2016) 670–680. F. Fındık, Recent developments in explosive welding, Mater. Des. 32 (3) (2011) 1081–1093. J.E. Lee, D.H. Bae, W.S. Chung, K.H. Kim, J.H. Lee, Y.R. Cho, Effects of annealing on the mechanical and interface properties of stainless steel/aluminum/copper cladmetal sheets, J. Mater. Process. Technol. 187–188 (2007) 546–549. M. Acarer, B. Demir, An investigation of mechanical and metallurgical properties of explosive welded aluminum–dual phase steel, Mater. Lett. 62 (25) (2008) 4158–4160. H. Xia, S. Wang, H. Ben, Microstructure and mechanical properties of Ti/Al explosive cladding, Mater. Des. 56 (2014) 1014–1019. R. Kacar, M. Acarer, Microstructure–property relationship in explosively welded duplex stainless steel–steel, Mater. Sci. Eng. A 363 (2003) 290–296. S.M. Aceves, F. Espinosa-Loza, J.W. Elmer, et al., Comparison of Cu, Ti and Ta interlayer explosively fabricated aluminum to stainless steel transition joints for cryogenic pressurized hydrogen storage, Int. J. Hydrog. Energy 40 (2015) 1490–1503. J. Ning, L.J. Zhang, G.C. Jiang, et al., Narrow gap multi-pass laser butt welding of explosion welded CP-Ti/Q235B bimetallic sheet by using a copper interlayer, J. Alloys Compd. 701 (2017) 587–602.