Effect of the flyer material on the interface phenomena in aluminium and copper explosive welds

Accepted Manuscript Effect of the flyer material on the interface phenomena in aluminium and copper explosive welds

G.H.S.F.L. Carvalho, R. Mendes, R.M. Leal, I. Galvão, A. Loureiro PII: DOI: Reference:

S0264-1275(17)30229-0 doi: 10.1016/j.matdes.2017.02.087 JMADE 2827

To appear in:

Materials & Design

Received date: Revised date: Accepted date:

23 January 2017 14 February 2017 27 February 2017

Please cite this article as: G.H.S.F.L. Carvalho, R. Mendes, R.M. Leal, I. Galvão, A. Loureiro , Effect of the flyer material on the interface phenomena in aluminium and copper explosive welds. The address for the corresponding author was captured as affiliation for all authors. Please check if appropriate. Jmade(2017), doi: 10.1016/j.matdes.2017.02.087

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ACCEPTED MANUSCRIPT EFFECT OF THE FLYER MATERIAL ON THE INTERFACE PHENOMENA IN ALUMINIUM AND COPPER EXPLOSIVE WELDS

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G.H.S.F.L. Carvalho1, R. Mendes2, R.M. Leal1, 3, I. Galvão1, 4, A. Loureiro1*

1 - CEMUC, Department of Mechanical Engineering, University of Coimbra,

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Portugal

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2 - ADAI, LEDAP, Department of Mechanical Engineering, University of

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Coimbra, Portugal

3 - ESAD.CR, Polytechnic Institute of Leiria, Portugal

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4 - ISEL, Department of Mechanical Engineering, Polytechnic Institute of Lisbon, Portugal

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*email - [email protected]; tel. - +351 239 790 700

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ACCEPTED MANUSCRIPT ABSTRACT The effect of physical and mechanical properties of three different flyers on the interface phenomena of partially overlapped explosive welds, using the same base plate material, was studied. Flyers of Copper Cu-DHP and aluminium alloy

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6082 (tempers T6 and O) were welded to AA6082-T6 base plates. The morphology of the weld interface is strongly influenced by the physical and

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mechanical properties of the flyer. In the interface of the aluminium welds, the

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use of a flyer of lower hardness and yield strength than the base plate results in

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asymmetrical waves, with bigger amplitude and smaller wavelength than the weld series of similar temper, and higher mechanical properties. The copper-

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aluminium welds presented flat interfaces, mainly because of the significant differences in melting temperature and density between the copper flyer and the

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aluminium base plate. Considering these results and analysing several dissimilar

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welds carried out by other authors it was found that when the product of density and melting temperature ratios between the flyer and the base plate exceeds a

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certain value, there is no formation of waves at the interface of the metals.

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Furthermore, for the Cu-Al welds, the CuAl2 (θ) intermetallic phase was formed on the bond zone.

Keywords: Explosive welding; Aluminium; Copper; Interface phenomena; Intermetallic compounds.

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ACCEPTED MANUSCRIPT 1. INTRODUCTION

Welding is considered one of the most important manufacturing processes in engineering since the joining of materials is an essential step in the development

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of various components for engineering projects. However, traditional arc welding processes change the microstructure and mechanical properties of base materials

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due to the application of severe thermal cycles and cause, among others,

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problems like loss of toughness, hot and cold cracking, formation of deleterious

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phases and a decrease in resistance to corrosion.

These problems are even more prominent in the welding of dissimilar

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materials because of their differences in physical properties, for instance the melting temperature. So, solid-state welding processes, such as explosive

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welding (EXW), are a viable alternative for joining dissimilar materials since

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they do not require the extensive melting of the materials to be joined. Nevertheless, dissimilar EXW has difficulties yet to be overcome, related to the

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phenomena occurring at the weld interface, mainly the formation of brittle

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intermetallic compounds (IMC) when certain metals, like copper and aluminium alloys, are joined together and also the interface phenomena regarding the wave formations [1–3].

Explosive welding is a solid-state welding process characterized by a highvelocity impact between two or more materials as the result of the controlled detonation of an explosive. The collision between the plates forms a metallic jet, removing contaminants and oxides from the surface of the plates. This jet is

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ACCEPTED MANUSCRIPT essential to clean the surfaces and guarantee a strong bond between the plates. After the jet, the surfaces are tightly compressed and held in direct contact under extreme pressure making the weld possible [4]. The explosive used (type, density and the presence of sensitizers) has a

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considerable influence in the process. Besides the explosive, there are also several parameters that influence the process. To represent the influence of these

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parameters, a tool called "weldability window" was developed that defines the

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weldability window is shown in Figure 1.

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conditions that must be met to accomplish good welds [5]. A schematic

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Figure 1 - Weldability window, reproduced from Blazynski [5] with Springer’s permission.

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ACCEPTED MANUSCRIPT Studies conducted by De Rosset [6], Ribeiro el al. [7] and Zlobin et al. [8] have aided the understanding of the related concepts, equations and applicability of the weldability window. This tool is related mainly with phenomena occurring at the interface and each line represents a limit that is related to a specific

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phenomenon. Put simply, the main conditions are as follows. The first condition is related to the formation of the metallic jet at the

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collision point. As mentioned above, this jet is essential to obtain a strong and

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consistent bond. Some authors relate this condition with the speed of sound in the

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material and with the β angle. Walsh et al. [9] were some of the researchers that formulated an equation for this limit, which is represented by the a-a’ line in

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Figure 1. The second limit is the minimum collision velocity (VC) which defines the minimum transition speed above which a wavy interface is obtained and it is

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represented by the e-e’ line [10]. The third condition is related to the impact

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velocity (VP) above which the resulting impact pressure exceeds the yield stress of the materials [7]. This limit corresponds to the f-f’ line, according to Deribas

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and Zakharenko [11]. The last threshold, identified as the g-g’ curve in the

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image, was proposed to avoid excessive melt at the bond zone [12]. Most of these curves are empirical and have limited application. It is important to emphasize that despite the weldability window indicating that a wavy interface is needed to establish a consistent bonding, Zlobin [13] mentioned that the formation of waves and the bonding of the materials are independent physical processes, assuming that the basic condition for welding the plates is to exceed the yield strength of the weakest material, forcing it to

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ACCEPTED MANUSCRIPT deform plastically. The same author also mentions that in some experiments, bonding occurs for a collision velocity below the one required for the formation of the waves while for other cases bonding is achieved for velocities higher than the wavy transition velocity but the bond line remains smooth. Several

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studies [1,14–17] have shown welds of dissimilar materials with a straight bond zone but displaying a consistent bond. This shows that, at least in some cases, the

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waves are not strictly necessary to achieve the welding.

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The phenomena arising at the interface of explosive welds are not

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completely understood and knowledge concerning dissimilar welds is limited. The interaction between different materials is problematic because of the

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metallurgical transformations at the bond zone. Regarding the alloys studied in this work, copper and aluminium are two of the three most used metals on earth

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(after iron) [18], and a combination of these alloys may have extensive

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applications in the electrical, refrigeration and transportation industries. Still, the

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physical properties of these two metals are quite different, as shown in

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ACCEPTED MANUSCRIPT Table 1, which hinders the establishment of the weld between them. A big difference in melting temperature can lead to metallurgical problems like hot cracks; the difference in thermal conductivity may result in abnormal heat flow during welding while differences in thermal expansion can lead to high residual

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stresses in the welds.

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ACCEPTED MANUSCRIPT Table 1 - Physical properties of Cu and Al with steel (1.0) as reference [19].

Melting temperature Thermal conductivity Thermal expansion

Copper 0.7 5.9 1.5

Aluminium 0.4 3.1 2.1

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Copper and aluminium are not mutually soluble in each other and the combination of these elements can result in the formation of harmful phases such

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as CuAl2 (θ)(blue), CuAl (η2) (green), Cu4Al3 (ζ2) (red) and Cu9Al4 (γ2) (yellow),

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as illustrated in Figure 2. Furthermore, the image shows a eutectic formation for

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an alloy composition of 67%Al in weight. Eutectic compositions always have a lower melting temperature than the pure materials and this usually complicates

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the welding process.

Figure 2 - Binary phase diagram of the Cu-Al system (Adapted from Drits et al. [20]).

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Some authors like Ashani and Bagheri [21]; Berski et al. [22] and Gülenç [15] have shown that copper can be joined to aluminium by explosive welding but the interface phenomena including intermetallic formation and the

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bond line morphology have not been explored deeply. Research addressing this combination usually focuses on mechanical properties. Albeit some find a typical

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wavy interface when aluminium is the flyer and copper the base plate [15], some

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authors [1,2,23] have found a smooth interface when projecting copper into

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aluminium. Since the alloys are the same, the differences found in the bond zone when the alloys switch positions may help us to understand the factors that lead

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to wavy or flat welds.

The aim of this research is to study the effect of the physical and

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mechanical properties of different flyers on the interface phenomena of explosive

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welds in order to better understand the main factors controlling the interface morphology and the characteristics that differentiate similar and dissimilar

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joining. For this purpose, the tests were performed using similar welding

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parameters but different flyers: two aluminium flyers with the same composition but different temper (different mechanical properties) and a copper flyer.

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ACCEPTED MANUSCRIPT 2. EXPERIMENTAL PROCEDURE

Three series of explosive welds with the same explosive and similar welding parameters were performed to evaluate the differences appeared on the interface

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when using different flyer plates. The same base plate material was used for all welds, a 3 mm thick aluminium alloy 6082-T6 and three different flyer plates of

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Cu-DHP (deoxidized high phosphorus copper), 6082-T6 aluminium alloy (“T6”

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temper is solution heat treated and artificially aged) and 6082-O aluminium alloy

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(“O” temper is annealed). Table 2 shows the average mechanical properties for each material, before welding, obtained by tensile tests and Vickers hardness

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measurements.

Alloy

Vickers Hardness 42 114 94

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AA6082-O AA6082-T6 Cu-DHP

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Table 2 - Mechanical properties of materials used in the experiments. Yield stress (MPa) 112 283 221

Tensile Strength (MPa) 118 341 249

Elongation on fracture (%) 9 17 26

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In order to simplify the discussion, the welds were identified as SWAA (Similar Welding - Aluminium-T6 flyer and Aluminium-T6 base plate), DWAA (Dissimilar Welding - Aluminium-O flyer and Aluminium-T6 base plate) and DWCA (Dissimilar Welding - Copper flyer and Aluminium-T6 base plate). The weld parameters and conditions of each weld are presented in Table 3. Most of the parameters were measured, except the flyer impact velocity (Vp) and the collision angle (β), which were calculated. 10

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Table 3 - Welding parameters.

DWAA

DWCA

Flyer alloy Base plate alloy Configuration Explosive

AA6082-T6

AA6082-O AA6082-T6 Partially overlapped Emulsion Explosive

Cu-DHP

ρ Explosive [kg/m³]

768

768

762

90 x 120

95 x 120

95 x 120

Thickness flyer [mm]

3

3

1

Thickness base [mm]

3 2.37

Stand-off distance [mm] Vd = Vc [m/s] Vp [m/s]* β [°]*

4 3523 1051 17

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Explosive Ratio

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Length x Width flyer/base [mm]

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SWAA

3

3

2.37

2.13

4 3477 1037 17

1.35 3430 973 16

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*Vp and β values were estimated according to Kennedy [24] and Cooper [25].

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The estimation of these values was made using the Gurney equation for a one dimensional problem in parallel configuration (Eq. 01), presented by

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Kennedy [24] and rearranged by Blazynski [5] which relates the terminal flyer

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velocity with the explosive ratio, and with the relation between the impact velocity and the β angle (Eq. 02) [5]. Where Vp is the flyer velocity, R is the explosive to flyer mass ratio, √2𝐸 is the Gurney characteristic velocity for a given explosive, Vd is the detonation velocity and β is the angle:

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ACCEPTED MANUSCRIPT 1/2

3𝑅 2 𝑉𝑝 = √2𝐸 ( 2 ) 𝑅 + 5𝑅 + 4

(𝐸𝑞. 02)

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𝛽 𝑉𝑝 = 2𝑉𝑑 𝑠𝑖𝑛 ( ) 2

(𝐸𝑞. 01)

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The value of √2𝐸 is different for each explosive and it is not always available in the literature. For most high explosives the Gurney characteristic

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velocity has already been evaluated [5,25–27]. However, for ammonium nitrate-

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based emulsion explosives, these data are not commonly available. Therefore, the methodology proposed by Cooper [25] which correlates the Gurney velocity with

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the detonation velocity: √2𝐸 = 𝑉𝑑 /2.97, was used to estimate the Gurney

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characteristic velocity. Nonetheless, it is essential to understand that this

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approximation presented by Cooper [25] is an average value derived from various explosives used by the author. Hence the values estimated in Table 3

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should not be considered as exact values, but estimated ones. But this estimation is essential to locate the welds in the weldability window and to give an idea of

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the impact velocity and the β angle (which is related to the impact velocity). The explosive used for all welds was an ammonium nitrate-based emulsion explosive with hollow glass microspheres. Emulsion explosives can be sensitized with different materials such as hollow glass or polymeric microspheres or mixed with metal particles such as aluminium [28,29]. The addition of these materials changes the density of the explosive mixture and consequently its detonation velocity. For DWCA, the explosive ratio (R) was 12

ACCEPTED MANUSCRIPT approximately 10% smaller than in the other series to compensate the differences in flyer thickness and density. The plates were placed in a partially overlapped configuration to obtain specimens that could be used for load transfer applications, since completely overlapped plates are mainly used for cladding

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applications. Figure 3 presents the configuration used in this work. The explosive ratio (R) was calculated considering only the plate section

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beneath the explosive, in this work. However, in partially overlapped welds the

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width of the flyer propelled by the detonation products is higher than the width of

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the explosive charge. As a consequence, the real velocity of the flyer plate should

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be lower than the estimated flyer velocity in Table 3.

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Explosive

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Base plate

Flyer plate

Support for the flyer

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Figure 3 - Schematic representation of the welding configuration.

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Samples were taken longitudinally to the direction of the detonation run and etched with two different etchants, Weck’s reagent and the Graff-Sargent reagent, for metallographic analysis. Weck’s reagent is indicated to analyse the grain boundaries and the presence of plastic deformation and the Graff-Sargent reagent, since it does not colour the microstructure, is good for analysing the morphology of the bond line, the presence of cracks and metallurgical transformations. Scanning electron microscopy (SEM) with Energy-dispersive 13

ACCEPTED MANUSCRIPT X-ray spectroscopy (EDS) were conducted with a Zeiss Merlin Compact/VP equipment to characterize the morphology of the interfaces and the new layer formed in the weld interface of the DWCA series. Microhardness profiles (HV0.2) were performed, as may be seen in

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Figure 4, in a straight line, oblique to the bond line, starting on the cladding and finishing on the base plate. This method allows smaller distance intervals from

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the bond line since each indentation must have a minimum distance from the

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other. The indentations were located 200μm from each other with one indentation

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positioned exactly on the interface. A more detailed analysis of the interface was

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done using a smaller load (HV0.025), in order to test the phases with smaller area.

Flyer plate

Base plate

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200μm

Figure 4 - Location of microhardness measurements.

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ACCEPTED MANUSCRIPT 3. RESULTS AND DISCUSSION

3.1. Weldability window

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The weldability window was studied for all the three welds, as shown in Figure 5: SWAA in green, DWAA in red and DWCA in blue. Each weld is also

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positioned in the respective window in the same image. The equations used for

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each limit were: Cowan (left limit) [10], Deribas and Zakharenko (lower

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limit) [11], Wittman (upper limit) [12] and Walsh (right limit) [9]. The value used for constant “N” in Wittman’s equation was 0.06 [7] since when using 0.11

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the welds were positioned far above the upper limit, which would not truly represent the experimental results since no excessive melt was found in any of

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the welds. It is possible to see that all three welds are inside three of the four

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boundaries that, in theory, correspond to a good weld. The welds were slightly above the upper limit defined by the Wittman equation but it must be taken into

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account that it is only an approximate position for the welds considering the

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limitations for the flyer plate velocity approximation and the ratio (R) calculation. The main phenomenon represented by the upper limit is the excessive melting and if no continuous molten layer is found, the welds are good from this approach. Since wave formation is one of the topics in this work, it is pertinent to describe the Cowan equation (left limit) [10], presented in (Eq. 03), where Rt is the Reynolds number which is 10.6; ρ is the density; H is the hardness and Vct is the transition collision velocity. According to Cowan [10],

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ACCEPTED MANUSCRIPT despite the density and hardness differences, the transition from the smooth to wavy bond zone occurs at nearly the same Reynolds values. So, it is essentially constant and 10.6 is the average for all the systems studied.

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DWCA SWAA

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β[°]

DWAA

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30

15

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10

2600

3000

3400

3800

4200

4600

5000

5400

Vc [m/s]

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2200

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5 0 1800

(𝐸𝑞. 03)

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(𝜌𝑏𝑎𝑠𝑒 + 𝜌𝑓𝑙𝑦𝑒𝑟 )𝑉𝑐𝑡 2(𝐻𝑏𝑎𝑠𝑒 + 𝐻𝑓𝑙𝑦𝑒𝑟 )

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𝑅𝑡 =

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Figure 5 - Weldability window for each weld conducted. SWAA represented in

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green, DWAA in red and DWCA in blue.

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ACCEPTED MANUSCRIPT 3.2. Microhardness

Figure 6 shows the hardness profile of each weld. The average hardness of each alloy before welding (BW) is also illustrated by dashed lines. It is possible to see

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that the hardness of AA6082-T6 is more similar to the Cu-DHP flyer than the AA6082-O flyer, which is softer than the others. All welded materials presented

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an increase in hardness caused by the mechanical work from the impact of the

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flyer plate.

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The SWAA series presents a hardness profile without great changes between the flyer and the base plate, because the main hardening mechanism for

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these materials is the formation of strengthening precipitates. The DWAA series presented great difference in hardness between the projected plate and the base

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plate because of the difference in the heat treatment between the aluminium

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alloys, but no substantial increase in hardness was observed considering the same reason abovementioned and because there is no formation of intermetallic

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phases. On the other hand, DWCA presented some differences between the plates

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and, most important, presented peaks in hardness precisely at the interface. Since this hardness is much higher than the strain hardening capability of these copper and aluminium alloys, this peak suggests that a harder phase was formed in the explosion welding process.

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ACCEPTED MANUSCRIPT FLYER PLATE

BASE PLATE

170

AA6082-T6

130

AA6082-T6

110 90

Cu-DHP

70

AA6082-O

-1.00

-0.75

50 -0.50

-0.25

30 0.00

0.25

0.50

Distance to interface (mm) SWAA

DWCA

O

T6

1.00

Cu

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DWAA

0.75

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Hardness (HV0.2)

150

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Figure 6 - Microhardness profiles.

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3.3. Interface analysis - aluminium weld series

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The main feature to be analysed in the aluminium alloy welds is the morphology

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of the weld since there is no intermetallic phase formation. Since the SWAA and DWAA aluminium welds presented similar morphology, they will be analysed

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together in order to facilitate the comparison of their characteristics. Figure 7 illustrates the characteristics of the interface of each weld series.

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It is possible to see that the welds between aluminium alloys display a similar wavy morphology regardless of the differences in the mechanical properties of the flyer. Some plastic deformation may also be observed next to the waves for the aluminium welds, especially for the flyer in O temper, see Figure 7b. Some melted areas are also present near the top and bottom of the waves, as indicated by arrows in Figure 7a and Figure 7b. Regarding the melted areas along the

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ACCEPTED MANUSCRIPT interface, the DWAA series presented more regions of melted material although not significantly different from the SWAA.

AA6082-T6

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AA6082-O

AA6082-T6

b

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a

AA6082-T6

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Figure 7 - Interface of welds SWAA (a) and DWAA (b). Optical microscopy -

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Weck’s reagent.

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In the SWAA weld it is possible to see transformed zones, presumably melted zones, next to the wave edge and sometimes at the wave crest, as

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indicated by an arrow in Figure 8a. Some small cracks were found inside these zones, as indicated by an arrow in Figure 8b. The etching with the Graff-Sargent

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reagent is not as intense as with Weck’s reagent, permitting a better detection of the cracks. Figure 9a and Figure 9b show that this transformation leads to a slight decrease in hardness compared to the average hardness value for the flyer and the base plate, which is of 133 HV0.2. This decrease in hardness suggests that a high temperature was achieved, at least above its recrystallization temperature, probably causing dissolution or coalescence of the aluminium strengthening precipitates. 19

ACCEPTED MANUSCRIPT The aluminium alloys of the 6xxx series are sensitive to solidification cracks because of the presence of Magnesium Silicide (Mg2Si) and the formation of a eutectic, which always has the lowest melting temperature. Cracks found in Figure 8b have the characteristics typical of solidification cracking, as suggested

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by Kah et al. [30] and Mathers [31]. Thus, the presence of these cracks, which are always related with problems in solidification, proves that localized melting

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occurred at the interface. Moreover, the presence of elongated grains in the

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deformation of the material during welding.

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neighbourhood of the interface, see Figure 8a, suggests a strong plastic

AA6082-T6

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AA6082-T6

b

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AA6082-T6

AA6082-T6

Figure 8 - Detail of a wave in the SWAA series. Optical microscopy - Weck’s reagent (a) and the Graff-Sargent etch (b).

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ACCEPTED MANUSCRIPT AA6082-T6

AA6082-T6

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AA6082-T6

b

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a

AA6082-T6

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Figure 9 - Microhardness indentations (HV0.025) on melted regions of two different waves (a and b) in the SWAA series. Optical microscopy - Graff-

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Sargent etch.

For the dissimilar aluminium-aluminium weld (DWAA) there is also some

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evidence of plastic deformation for both aluminium alloys, as illustrated in

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Figure 10a. The grains are clearly defined in the AA6082-T6 alloy in the lower part of the image, while the material flow is well marked at the top. Melted areas

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marked by arrows are still visible in the same image, containing shrinkage cavities resulting from shrinkage during solidification. In Figure 10b, these

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solidification cracks in another wave of the same weld and etched with GraffSargent reagent are even clearer. Though not so different, the DWAA series presented more melted zones than the SWAA. The hardness measurements in these zones (Figure 11a and Figure 11b) showed values between the average hardness found in the profile hardness tests for the flyer in “O” in an annealed condition (46 HV0.2) and the “T6”base plate solution heat treated and artificially aged (134 HV0.2). 21

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AA6082-O

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AA6082-O

AA6082-T6

b

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a

AA6082-T6

Figure 10 - Detail of a wave in the DWAA series. Optical microscopy - Weck’s

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reagent (a) and Graff-Sargent etch (b).

AA6082-O

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AA6082-O

AA6082-T6

a

AA6082-T6

b

Figure 11 - Microhardness indentations (HV0.025) in melted regions of two different waves (a and b) in the DWAA series. Optical microscopy - GraffSargent etch.

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ACCEPTED MANUSCRIPT Besides the general similarities in morphology, these two aluminium welds presented differences at the interface when carefully analysed. For a better comparison between the two morphologies, the profile of the wave geometry at the interface in both weld series is shown in Figure 12. Comparing the interface

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of both welds, it is possible to see that the shapes and sizes of the waves are different. In the similar welds (in a solid blue line) the waves have a smaller

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average amplitude, bigger average wavelength and are more symmetric than

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those in the dissimilar weld series (in a dashed red line).

λ = 407 μm (average)

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DWAA

λ = 421 μm (average)

A = 149 μm (average)

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SWAA

A = 164 μm (average)

Figure 12 - Comparison of the wave geometries between the similar (solid

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blue line) and dissimilar (dashed red line) aluminium welds series.

Jaramillo et al. [32] studied the transition from smooth to wavy bond lines in EXW for three different systems (Cu-Cu, Fe-Fe and Al-Al) and reported that the most common theories for wave formation in explosion welding are based on hydrodynamic analogy of metals under high pressure, shock and rarefaction wave interaction and dynamic plasticity. The hydrodynamic theory proposed by Cowan and Holtzman [33], which describes the formation of bond zone waves in

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ACCEPTED MANUSCRIPT explosion clads is analogous to fluid flow around an obstacle, is the most widespread for interface analysis. Since the waves are also the result of plastic deformation, which is directly related to the yield strength of the material, this mechanical property is related to

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the wave formation. Albeit having used the same alloy for the base plate and flyer plate, the mechanical properties of the flyers were very distinct. These

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differences shifted the left limit of the weldability window (Cowan equation) to

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the right when the flyer is changed from the “O” temper to the “T6” temper, see

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Figure 5. This is related to the differences in hardness and yield strength, caused by the different temper conditions. Figure 12 shows that conducting the weld

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with the same alloy but with lower hardness and yield strength for the flyer plate, the wave starts to become asymmetrical, bigger in amplitude and with a higher

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frequency (smaller wavelength).

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It may be seen that despite the relation between the yield strength and the wave formation, this characteristic is not considered in most common equations

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for the interface, although some equations, like the one introduced by

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Cowan [10], include the hardness of the materials.

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ACCEPTED MANUSCRIPT 3.4. Interface analysis - Copper-Aluminium weld series

The interface of the copper-aluminium (DWCA) welds is more complex, mainly because dissimilar combinations often involve complex microstructures and

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metallurgical transformations. The combination of aluminium and copper, according to the phase diagram in Figure 2, can form several intermediate

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compounds.

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The first characteristic that can be detected by the metallographic analysis

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is that even using similar parameters and the same base plate in all welds, the bond zone morphology of the copper-aluminium weld DWCA is quite different

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from the other abovementioned series. Figure 13a shows the characteristic morphology of the weld, which is flat, without effective waves. Besides the flat

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interface, the same figure shows the formation of a new layer between the two

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materials. This layer is nearly continuous but with small breaks over the length and a variable thickness of about a few microns. This suggests that a molten layer

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is formed at the interface.

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In Figure 13b, which illustrates a piece of a weld interface under higher magnification, the presence of a new phase can be observed as well as some cracks, marked with arrows in the picture. These cracks suggest that the phase, unlike the aluminium and copper alloys, is very brittle. It is natural that when using higher magnification as in Figure 13b, the boundary between the flyer and base plate does not seem completely smooth.

25

ACCEPTED MANUSCRIPT Cu-DHP

AA6082-T6

AA6082-T6

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Cu-DHP

b

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a

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Figure 13 - Characteristic morphology (a) and detail (b) of the copperaluminium interface. The arrows indicate the cracks in the new phase formed in

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the welding process. Optical microscopy - Weck’s reagent.

Microhardness was measured in several regions of the boundary to

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characterize this new compound. Figure 14a and Figure 14b show two of these

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regions, as well as the hardness values measured at each site. The great difference in hardness between the copper flyer and the aluminium base plate and

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the new compound is made evident by the Vickers pyramidal indentation size. The extremely high hardness of this compound suggests the presence of

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intermetallic compounds in this region. This intermetallic layer has an average thickness of 16μm, ranging from 0 (zones without intermetallic) to 55μm.

26

ACCEPTED MANUSCRIPT Cu-DHP

Cu-DHP

AA6082-T6

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AA6082-T6

b

RI

a

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Figure 14 - Microhardness indentations (HV0.025) in the intermetallic phase of two different locations on the interface (a and b). Cu-flyer (above) and Al-base

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plate (below). Optical microscopy - unetched.

The identification of these compounds through their crystallographic

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structure is difficult due to their diminutive size. The chemical composition of

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the compounds, together with the hardness measurements, was used to distinguish the phase (or phases) present in the interface. Hence, the EDS method

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was used to evaluate the chemical composition at several locations (A, B, C, D

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and E) next to the hardness indentations, as shown in Figure 15.

27

ACCEPTED MANUSCRIPT

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Cu-DHP

SC

RI

AA6082-T6

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Figure 15 - Microstructure of the DWCA weld interface, revealed by SEM, indicating the location of EDS analyses related to microhardness measurements -

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unetched.

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The results of chemical analysis for each location are shown in Table 4.

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Regions D and E represent respectively the 6082 aluminium alloy base plate and

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the copper alloy flyer.

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Table 4 - Chemical composition at different regions (EDS analysis). Spectrum Label A B C D E CuAl2 (θ) [34]

Al 61.99 62.78 61.22 98.65 -

Atomic % Cu Si 36.94 0.45 35.92 0.74 37.99 0.45 0.74 100 31.9-33.0

Mg 0.62 0.57 0.35 0.61 -

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ACCEPTED MANUSCRIPT The chemical compositions obtained for all the three regions, A, B and C, which are located inside the intermediate phase are similar and close to the CuAl2 (θ) intermetallic phase, according to Murray [34] analysis of the Al-Cu system. The hardness measured (between 602 and 800 HV0.025) is higher than the

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values found by Chen and Hwang [35] (between approximately 580-650 HV0.010), Ouyang et al. [36] (between 486-557 HV0.2) and Mróz et al. [37]

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(between 449.4 and 685.9 HV0.05). Nonetheless, Chen and Hwang [35]; Guo et

SC

al. [38]; Wei et al. [39], in their studies of Al-Cu intermetallics and diffusion,

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reported that CuAl2 is the first phase to be formed, only after the next reaction is the formation of Cu9Al4 (γ2). Moreover, CuAl2 is always reported as being found

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in Cu-Al explosive welding, even when the other two most common phases CuAl (η2) and Cu9Al4 (γ2) are found. In fact, Mróz et al. [37] and Dyja et al. [40]

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also found only CuAl2 intermetallic in Cu-Al explosive welds of round products.

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For this reason, it is suggested that CuAl2 (θ) is the intermetallic phase found. Still, intermetallic phases in explosive welding are not expected to be formed

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under equilibrium conditions as represented by the phase diagram, considering

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the high velocity of the process and severe cooling rate. Therefore, the formation of each phase may be different from the one represented in the phase diagram, since it depends not only on temperature but also on the time and cooling rate. It sometimes happens that in melted areas in an explosion welding interface, the chemical composition is not uniform, suggesting the formation of several intermetallic compounds, as reported by Loureiro et al. [1]. Particularly, the phases near the copper are richer in copper and those near the aluminium are

29

ACCEPTED MANUSCRIPT richer in aluminium. Chen and Hwang [35] did an interesting investigation, reporting the stages of intermetallic formation in Al/Cu bimetal. They report that first CuAl2 forms on the aluminium side, and after that Cu9Al4 forms on the copper side. A chemical mapping of copper and aluminium was performed by

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EDS to verify if there were any changes in the chemical composition along the intermetallic thickness. Figures 17a and b show the mapping for copper and

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aluminium respectively while Figure 16c shows both the copper and aluminium

SC

maps overlapped. Except for some small islands of pure alloys, the distribution

NU

of elements does not present changes that could indicate the presence of more than one IMC. The distribution of the two elements in the melted zone is

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D

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uniform, as can be seen in Figure 16c.

b

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a

c

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ACCEPTED MANUSCRIPT Figure 16 - EDS analysis presenting the map of copper in green (a), the map of aluminium in red (b) and both maps overlapped (c).

As said previously, CuAl2 is reported as the easiest IMC to form. Actually,

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aluminium-rich intermetallic compounds are more likely to form first because copper atoms will diffuse into the aluminium alloy and will be surrounded by

RI

several aluminium atoms. This can be explained by two main factors: the first is

SC

that the diffusion of copper into aluminium is much easier than the diffusion of

NU

aluminium into copper, considering that, in general, it is favoured in metals with a lower density and melting temperature. The second factor is that aluminium has

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a much lower melting temperature (the composition of the eutectic point is also rich in aluminium), hence the aluminium will melt first and will also contribute

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to the flow of copper towards the aluminium alloy. However, the melting of the

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copper is also very likely. Indeed, Wei et al. [39] studied the intermetallic formation at an Al-Cu joint and through the Gibbs free energy, concluded that

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this is the intermetallic compound usually expected to be formed first.

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There is no doubt that there is a mixture of elements but all the phenomena occurring during the welding process are not completely understood. Since there is a new phase, with a new crystallographic structure (and since it is an intermetallic phase) it is undeniable that there is an atomic interaction between the elements. For that reason, atomic diffusion is one of the possible phenomena that are likely to occur in the process. It may seem contradictory to suggest diffusion since it needs time to occur and explosive welding is an extremely fast

31

ACCEPTED MANUSCRIPT process. Nevertheless, it is necessary to understand all aspects acting in the process. When dealing with solid-state processes, especially explosion welding, the process includes factors like high temperature, high pressure, plastic deformation and direct contact between the materials. All these factors not only

PT

lead to diffusion, but also accelerate any atomic diffusion occurring in the system. Hence, the time required to have an atomic mixture between the plate

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elements will be reduced. Moreover, a capital process characteristic that shall be

SC

taken into account is the presence of fusion. Regardless of being called a solid-

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state process, it is known that it is almost impossible to avoid some melting in most impact welding processes [22,41,42], particularly in systems with the

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formation of eutectic compositions like Al-Cu. And when it comes to atomic interaction, it is known that the diffusion is much easier in liquid media. Thereby,

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there is a strong possibility that diffusion is the main mechanism promoting the

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interaction and mixture of elements.

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3.5. Wave formation

Three different flyer materials were projected against the same base plate material with similar welding parameters. Both welds with aluminium flyers, besides the differences in mechanical properties, presented wavy interfaces while the weld using a copper flyer presented a smooth interface. The mechanisms, theories and physics of wave formations will not be discussed here but some

32

ACCEPTED MANUSCRIPT characteristics can be highlighted to understand the differences in interface morphology considering that the welds were conducted with the same base plate. The mechanical properties of the copper and AA6082-T6 used as flyers in the SWAA and DWCA are not extremely different. So, this difference only by

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itself cannot justify the discrepancies found in the interface morphology of the aluminium welds (SWAA) and the copper weld (DWCA). By analysing these

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two flyer alloys it was possible to identify four major differences between them:

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melting temperature, thermal and electrical conductivity and density. Analysing

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other dissimilar explosive welds that obtained a flat interface from other researchers (Cu-Al [1,2]; Ta-Al [2]; Nb-Al [43] and Ta-Cu [44]), some aspects

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should be highlighted. In these cases, shown in Table 5, the flyer material always had a higher density and melting temperature than the base plate material. The

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flyer density in these cases is at least 1.9 times higher than the base plate density,

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and the melting temperature at least 1.6 times higher. Even with a high detonation velocity (4500 m/s) in Cu-Al welds, Amani and Soltanieh [3] did not

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find a typically wavy bond, only very low amplitude waves. Thus, the

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weldability window may not be suitable for all materials combinations, especially for combinations of materials with significant physical differences. Table 5 lists several dissimilar explosive welds found in the literature in order to compare their density and melting temperature. Despite the welds with a straight bond line always having a flyer with a higher density and melting temperature than the base plate, there are also explosive welds with that same characteristic that did not result in smooth bond lines. Still, the ratio between the

33

ACCEPTED MANUSCRIPT flyer and the base plate densities in these wavy welds was always equal or below 1.8, which is the case of stainless steel-Ti [2]. For the melting temperature, the ratio was always below 2.5, the Ti-Al [2] case. Therefore, when the flyer is denser than the base plate and has a much

PT

higher melting temperature, it is possible that above a certain ratio, the substantial difference in density and melting point will prevent the formation of a

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wavy interface, even when the collision velocity of the weld is above the

SC

transition velocity calculated by the Cowan equation. In fact, after analysing this

NU

equation (Eq. 03), it is clear that by increasing the density (regardless of whether it is the base or flyer plate) the collision velocity needed for a wavy bond

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decreases. However, this does not seem to be observed in the experiments in Table 5. The melting temperature, on the other hand, does not have an influence

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on the Cowan equation (Eq. 03). Therefore, it is possible that the difference in

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the density and the melting temperature between the flyer and the base plate may dictate the morphology of the interface. One example of this assertion is when

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the aluminium and copper plates switch positions. Gülenç [15] studied this

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combination but with aluminium as the flyer plate and copper as the base plate. With aluminium (lower melting point and density) as the flyer plate instead of copper, it was possible to achieve a wavy bond zone instead of a smooth line. In order to investigate the influence of these two physical properties further and according to (Eq. 04), the density ratio (ρR) and melting temperature ratio (TR) were multiplied. This product was called the wave interface factor (WIF) and is also included in Table 5. The values found for the WIF relate

34

ACCEPTED MANUSCRIPT directly with the morphology of the interface: above 5.1 the bond line is flat and below 4.2 the bond line is wavy. There is an exception, a Cu-Al weld [15] that has a 5.4 coefficient and the interface is not reported as being smooth. However, on analysing this Cu-Al weld [15] the waves seem to have a very low amplitude

PT

and were reported only with a much larger amplification than is usually

𝜌𝑏𝑎𝑠𝑒 𝑝𝑙𝑎𝑡𝑒

×

𝑇𝑓𝑙𝑦𝑒𝑟 𝑇𝑏𝑎𝑠𝑒 𝑝𝑙𝑎𝑡𝑒

= 𝜌𝑅 × 𝑇𝑅 (𝐸𝑞. 04)

SC

𝜌𝑓𝑙𝑦𝑒𝑟

NU

𝑊𝐼𝐹 =

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necessary to detect the waves was applied.

In Figure 17 the WIF values are plotted together for the welds found in the

MA

literature and presented on a logarithmic scale. It is possible to see that there is a clear separation between the welds that presented a flat interface (higher WIF)

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and the welds with a wavy interface (lower WIF).

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It is clear that the analysis in Figure 17 has some limitations since these studies were performed by different researchers and detailed information on the

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welding conditions and parameters for each experiment is not available.

AC

Nevertheless, it suggests an interesting approach that can predict the interface morphology and also helps to better understand why the bond zone is sometimes wavy and in others cases smooth. The wave interface factor (WIF) does not directly indicate whether the interface will be wavy or flat, it indicates the possibility of obtaining a wavy bond line. That is, above a certain value, and even with the appropriate welding parameters, the weld will not be wavy and below

35

ACCEPTED MANUSCRIPT that value the weld may or may not be wavy, depending on the welding parameters.

Table 5 - Physical properties [18] and interface morphology of several dissimilar

Ta-Al [2] Nb-Al [43] Cu-Al [1,2,23] Cu-Al [3] Ta-Cu [44] Ti-Al [2] Al-Mg [45] SS-Ti [2] Zr-Fe [46] SS-Fe [47,48] Cu-SS [49,50] Ag-Fe [44] Cu-Fe [51] Ti-Cu [52] Ti-SS [17,53] Ti-Fe [54] Pb-Cu [55] Pb-Fe [55] Al-Cu [15] Al-Ta [44]

16.6 8.57 8.93 8.93 16.6 4.51 2.7 8.0 6.5 8.0 8.93 10.49 8.93 4.51 4.51 4.51 11.3 11.3 2.7 2.7

WIF

Interface morphology

27.91 11.87 5.44 5.44 5.13 4.22 1.58 1.49 0.99 0.93 0.87 0.83 0.80 0.78 0.67 0.62 0.38 0.31 0.18 0.04

Flat Flat Flat VLAW Flat Wavy Wavy Wavy Wavy Wavy Wavy* Wavy Wavy Wavy Wavy* Wavy Wavy Wavy Wavy Wavy

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D

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NU

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Density (g/cm³)

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Melting Density Melting temp. (°C) temp. ratio B.Plate ratio “ρR” Flyer B.Plate “TR” 2.7 6.1 2996 660 4.5 2.7 3.2 2468 660 3.7 2.7 3.3 1085 660 1.6 2.7 3.3 1085 660 1.6 8.93 1.9 2996 1085 2.8 2.7 1.7 1668 660 2.5 1.74 1.6 660 650 1.0 4.51 1.8 1400 1668 0.8 7.87 0.8 1852 1538 1.2 7.87 1.0 1400 1538 0.9 8.0 1.1 1085 1400 0.8 7.87 1.3 962 1538 0.6 7.87 1.1 1085 1538 0.7 8.93 0.5 1668 1085 1.5 8.0 0.6 1668 1400 1.2 7.87 0.6 1668 1538 1.1 8.93 1.3 327 1085 0.3 7.87 1.4 327 1538 0.2 8.93 0.3 660 1085 0.6 16.6 0.2 660 2996 0.2

Combination Flyer-base plate

Flyer

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explosive welds.

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VLAW - very low amplitude waves; SS - stainless steel. * Flat and wavy interfaces were found in these studies. Nonetheless, they analyse the explosive ratio and flat interfaces were obtained for a low ratio.

36

ACCEPTED MANUSCRIPT

10

VLAW

Wavy

Transition

Flat interface Wavy interface

1 0.1

Ta-Al [2]

Nb-Al [43]

Cu-Al [1,2,23]

Cu-Al [3]

Ta-Cu [44]

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Ti-Al [2]

Al-Mg [45]

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SS-Ti [2]

Zr-Fe [46]

SS-Fe [47,48]

Cu-SS [49,50]

Ag-Fe [44]

Cu-Fe [51]

Ti-Cu [52]

Ti-SS [17,53]

Ti-Fe [54]

Pb-Cu [55]

Pb-Fe [55]

Al-Cu [15]

0.01 Al-Ta [44]

Weld interface factor (WIF)

Flat 100

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Figure 17 - Distribution of the ρR x TR factor on a logarithmic scale.

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4. CONCLUSIONS

The effect of the physical and mechanical properties of the flyer material

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on the weldability window and interface morphology of explosive welds was

•

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studied in this research. The following conclusions could be reached: The morphology of the weld bond interface is strongly influenced by

For the aluminium welds, the use of a flyer with a hardness and yield

AC

•

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the physical and mechanical properties of the flyer plate and base plate.

strength lower than the base plate results in asymmetrical waves, bigger in amplitude and smaller in wavelength than similar welds.

•

The copper-aluminium welds did not display a wavy bond zone even though they were inside the weldability window, at the right of the Cowan equation line. That means that the weldability window should be improved for dissimilar welding.

37

ACCEPTED MANUSCRIPT •

It was found that when the flyer is denser than the base plate, a strong density difference can prevent the formation of a wavy interface. The same happens if the melting temperature of the flyer is substantially higher than the base plate.

•

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Combining the ratios of these two physical properties for the flyer and base plate, which results in the weld interface factor (WIF), a threshold

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above which a flat weld interfaces is always obtained was found. •

SC

The SEM and hardness measurements suggest that mainly the CuAl2

NU

(θ) intermetallic phase forms in the bond line of Cu-Al welds.

MA

It is suggested that the density and melting temperature should be analysed more profoundly in order to understand the real influence of these properties on

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interface morphology. Once proved and scaled, this influence should be

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dissimilar welds.

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considered and new equations for the weldability window may be developed for

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ACKNOWLEDGEMENTS

The authors are indebted to the Portuguese Foundation for the Science and Technology

(FCT)

for

the

financial

support

through

the

Project

UID/EMS/00285/2013. The author G.H.S.F.L. Carvalho is supported by the Brazilian National Council for Scientific and Technological Development (CNPq).

38

ACCEPTED MANUSCRIPT REFERENCES

[1]

A. Loureiro, R. Mendes, J.B. Ribeiro, R.M. Leal, I. Galvão, Effect of explosive mixture on quality of explosive welds of copper to

S.M. Aceves, F. Espinosa-Loza, J.W. Elmer, R. Huber, Comparison

RI

[2]

PT

aluminium, Materials and Design 95 (2016) 256–267.

SC

of Cu, Ti and Ta interlayer explosively fabricated aluminum to stainless steel transition joints for cryogenic pressurized hydrogen

NU

storage, International Journal of Hydrogen Energy 40 (2015) 1490–

[3]

MA

1503.

H. Amani, M. Soltanieh, Intermetallic Phase Formation in

D

Explosively Welded Al/Cu Bimetals, Metallurgical and Materials

PT E

Transactions B: Process Metallurgy and Materials Processing Science 47 (2016) 2524–2534.

V.D. Linse, P.I. Temple, Explosion Welding, in: AWS Welding

CE

[4]

AC

Handbook – Volume 2 - Welding Processes, 8th ed., American Welding Society, USA, 1991: pp. 765–781. [5]

T.Z. Blazynski, Explosive Welding, Forming and Compaction, 1st ed., Applied Science Publishers Ltd, Essex, England, 1987.

[6]

W.S. de Rosset, Analysis of Explosive Bonding Parameters, Materials and Manufacturing Processes 21 (2006) 634–638.

[7]

J.B. Ribeiro, R. Mendes, A. Loureiro, Review of the weldability 39

ACCEPTED MANUSCRIPT window concept and equations for explosive welding, Journal of Physics: Conference Series 500 (2014) 1–6. [8]

B. Zlobin, V. Sil’Vestrov, A. Shtertser, A. Plastinin, V. Kiselev, Enhancement of Explosive Welding Possibilities by the Use of

PT

Emulsion Explosive, Archives of Metallurgy and Materials 59

J.M. Walsh, R.G. Shreffler, F.J. Willig, Limiting conditions for jet

SC

[9]

RI

(2014).

formation in high velocity collisions, Journal of Applied Physics 24

NU

(1953) 349–359.

MA

[10] G.R. Cowan, O.R. Bergmann, A.H. Holtzman, Mechanism of bond zone wave formation in explosion-clad metals, Metallurgical

D

Transactions 2 (1971) 3145–3155.

PT E

[11] A.A. Deribas, I.D. Zakharenko, Surface effects with oblique collisions between metallic plates, Combustion, Explosion, and

CE

Shock Waves 10 (1974) 358–367.

AC

[12] S.H. Carpenter, R.H. Wittman, Explosion Welding, Annual Review of Materials Science 5 (1975) 177–199. [13] B.S. Zlobin, Explosion welding of steel with aluminum, Combustion, Explosion and Shock Waves 38 (2002) 374–377. [14] J.H. Han, J.P. Ahn, M.C. Shin, Effect of interlayer thickness on shear deformation behavior of AA5083 aluminum alloy/SS41 steel plates manufactured by explosive welding, Journal of Materials Science 38 40

ACCEPTED MANUSCRIPT (2003) 13–18. [15] B. Gülenç, Investigation of interface properties and weldability of aluminum and copper plates by explosive welding method, Materials and Design 29 (2008) 275–278.

PT

[16] M.M. Hoseini Athar, B. Tolaminejad, Weldability window and the

RI

effect of interface morphology on the properties of Al/Cu/Al

SC

laminated composites fabricated by explosive welding, Materials and Design 86 (2015) 516–525.

NU

[17] N. Kahraman, B. Gülenç, F. Findik, Joining of titanium/stainless steel

MA

by explosive welding and effect on interface, Journal of Materials Processing Technology 169 (2005) 127–133.

D

[18] ASM Handbook, Volume 2 - Properties and selection: Nonferrous

PT E

alloys and special-purpose materials, 10th ed., ASM International, Materials Park, Ohio, USA, 1990.

CE

[19] S.J. Matthews, D.G. Howden, N.F. Bratkovich, J.F. King, Dissimilar

AC

Metals, in: AWS - Welding Handbook - Volume 4 - Metals and Their Weldability, American Welding Society, USA, 1982: pp. 513–547. [20] M.E. Drits, N.R. Bochvar, L.S. Guzei, E. V. Lysova, E.M. Padezhnova, L.L. Rokhlin, N.I. Turkina, Cu-Al, in: Binary and Multicomponent Copper-Base Systems [in Russian], Nauka, Moscow, Soviet Union, 1979: p. 8. [21] J.Z. Ashani, S.M. Bagheri, Explosive scarf welding of aluminum to 41

ACCEPTED MANUSCRIPT copper plates and their interface properties, Materialwissenschaft Und Werkstofftechnik 40 (2009) 690–698. [22] S. Berski, Z. Stradomski, H. Dyja, Quality of bimetal Al-Cu joint after explosive cladding, Journal of Achievements in Materials and

PT

Manufacturing Engineering 22 (2007) 73–76.

RI

[23] D. Rydz, M. Krakowiak, T. Bajor, G. Stradomski, The Analysis of

SC

the Impact Relative Rolling Reduction at the Area Connector Bimetallic Plate After the Rolling Process, Solid State Phenomena

NU

199 (2013) 514–517.

MA

[24] J.E. Kennedy, Gurney energy of explosives: Estimation of the velocity and impulse imparted to driven metal, New Mexico, USA,

D

1970.

PT E

[25] P.W. Cooper, Explosives Engineering, Wiley-VCH, New York, USA, 1996.

CE

[26] J.E. Kennedy, Explosive output for driving metal, in: Proceedings of

AC

the 12th Annual Symposium of the New Mexico Section of the ASME, New Mexico, USA, 1972. [27] D.R. Hardesty, J.E. Kennedy, Thermochemical estimation of explosive energy output, Combustion and Flame 28 (1977) 45–59. [28] R. Mendes, J. Ribeiro, I. Plaksin, J. Campos, Non ideal detonation of emulsion explosives mixed with metal particles, AIP Conference Proceedings 1426 (2012) 267–270. 42

ACCEPTED MANUSCRIPT [29] R. Mendes, J. Ribeiro, I. Plaksin, J. Campos, B. Tavares, Differences between the detonation behavior of emulsion explosives sensitized with glass or with polymeric micro-balloons, Journal of Physics: Conference Series 500 (2014) 2–8.

PT

[30] P. Kah, E. Hiltunen, J. Martikainen, Investigation of hot cracking in

RI

the welding of aluminium alloys (6005 & 6082), in: 63rd Annual

Istanbul, Turkey, 2010: pp. 373–380.

SC

Assembly of the International Institute of Welding, IIW 2010,

NU

[31] G. Mathers, The welding of Aluminium and its alloys, 1st ed.,

MA

Woodhead Publishing Limited, Cambridge, England, 2002. [32] D. Jaramillo, A. Szecket, O.T. Inal, On the transition from a waveless

D

to a wavy interface in explosive welding, Materials Science and

PT E

Engineering 91 (1987) 217–222. [33] G.R. Cowan, A.H. Holtzman, Flow configurations in colliding plates:

CE

Explosive bonding, Journal of Applied Physics 34 (1963) 928–939.

AC

[34] J.L. Murray, The aluminium-copper system, International Metals Reviews 30 (1985) 211–234. [35] C.-Y. Chen, W.-S. Hwang, Effect of Annealing on the Interfacial Structure of Aluminum-Copper Joints, Materials Transactions 48 (2007) 1938–1947. [36] J. Ouyang, E. Yarrapareddy, R. Kovacevic, Microstructural evolution in the friction stir welded 6061 aluminum alloy (T6-temper condition) 43

ACCEPTED MANUSCRIPT to copper, Journal of Materials Processing Technology 172 (2006) 110–122. [37] S. Mróz, P. Szota, A. Stefanik, S. Wąsek, G. Stradomski, Analysis of Al-Cu bimetallic bars properties after explosive welding and rolling

PT

in modified passes, Archives of Metallurgy and Materials 60 (2015)

RI

427–432.

SC

[38] Y. Guo, G. Liu, H. Jin, Z. Shi, G. Qiao, Intermetallic phase formation in diffusion-bonded Cu/Al laminates, Journal of Materials Science 46

NU

(2011) 2467–2473.

MA

[39] Y. Wei, J. Li, J. Xiong, F. Zhang, Investigation of interdiffusion and intermetallic compounds in Al–Cu joint produced by continuous

D

drive friction welding, Engineering Science and Technology, an

PT E

International Journal 19 (2016) 90–95. [40] H. Dyja, S. Mróz, Z. Stradomski, Properties of joint in the bimetallic

CE

rods Cu-Al and Cu-Steel after explosive cladding and the process of

AC

rolling, Metalurgija 42 (2003) 185–191. [41] A. Stern, V. Shribman, A. Ben-Artzy, M. Aizenshtein, Interface Phenomena and Bonding Mechanism in Magnetic Pulse Welding, Journal of Materials Engineering and Performance 23 (2014) 3449– 3458. [42] A. Stern, M. Aizenshtein, Bonding zone formation in magnetic pulse welds, Science and Technology of Welding and Joining 7 (2002) 44

ACCEPTED MANUSCRIPT 339–342. [43] T. Palmer, J. Elmer, D. Brasher, D. Butler, R. Riddle, Development of an Explosive Welding Process for Producing High-Strength Welds between Niobium and 606-T651 Aluminum, Welding Journal 85

PT

(2006) 252–263.

RI

[44] B.A. Greenberg, M.A. Ivanov, V. V. Rybin, O.A. Elkina, O. V.

SC

Antonova, A.M. Patselov, A. V. Inozemtsev, A. V. Plotnikov, A.Y. Volkova, Y.P. Besshaposhnikov, The problem of intermixing of

NU

metals possessing no mutual solubility upon explosion welding (Cu-

MA

Ta, Fe-Ag, Al-Ta), Materials Characterization 75 (2013) 51–62. [45] N. Zhang, W. Wang, X. Cao, J. Wu, The effect of annealing on the

D

interface microstructure and mechanical characteristics of

PT E

AZ31B/AA6061 composite plates fabricated by explosive welding, Materials and Design 65 (2015) 1100–1109.

CE

[46] A. Nobili, J. Banker, Recent developments in Zirconium-Steel

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Explosion Clad, in: Proceedings 1999 Reactive Metals in Corrosive Applications Conference, Wah Chang, Albany USA, 1999: pp. 83– 88.

[47] R. Mendes, J.B. Ribeiro, A. Loureiro, Effect of explosive characteristics on the explosive welding of stainless steel to carbon steel in cylindrical configuration, Materials and Design 51 (2013) 182–192. 45

ACCEPTED MANUSCRIPT [48] R. Kacar, M. Acarer, An investigation on the explosive cladding of 316L stainless steel-din-P355GH steel, Journal of Materials Processing Technology 152 (2004) 91–96. [49] A. Durgutlu, B. Gülenç, F. Findik, Examination of copper/stainless

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steel joints formed by explosive welding, Materials and Design 26

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(2005) 497–507.

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[50] A. Durgutlu, H. Okuyucu, B. Gülenç, Investigation of effect of the stand-off distance on interface characteristics of explosively welded

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copper and stainless steel, Materials and Design 29 (2008) 1480–

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1484.

[51] Z. Livne, A. Munitz, Characterization of explosively bonded iron and

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copper plates, Journal of Materials Science 22 (1987) 1495–1500.

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[52] N. Kahraman, B. Gülenç, Microstructural and mechanical properties of Cu-Ti plates bonded through explosive welding process, Journal of

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Materials Processing Technology 169 (2005) 67–71.

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[53] S.A.A. Akbari Mousavi, P. Farhadi Sartangi, Experimental investigation of explosive welding of cp-titanium/AISI 304 stainless steel, Materials and Design 30 (2009) 459–468. [54] Q. ling Chu, M. Zhang, J. hong Li, Q. Jin, Q. yang Fan, W. wei Xie, H. Luo, Z. yue Bi, Experimental investigation of explosion-welded CP-Ti/Q345 bimetallic sheet filled with Cu/V based flux-cored wire, Materials and Design 67 (2015) 606–614. 46

ACCEPTED MANUSCRIPT [55] R. Prümmer, The use of Lead in explosive welding, Propellants and

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Explosives 1 (1976) 103–107.

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ACCEPTED MANUSCRIPT FIGURES LIST

Figure 1 - Weldability window, reproduced from Blazynski [5] with Springer’s permission.

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Figure 2 - Binary phase diagram of the Cu-Al system (Adapted from Drits

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et al. [20]).

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Figure 3 - Schematic representation of the welding configuration. Figure 4 - Location of microhardness measurements.

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Figure 5 - Weldability window for each weld conducted. SWAA

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represented in green, DWAA in red and DWCA in blue. Figure 6 - Microhardness profiles.

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Figure 7 - Interface of welds SWAA (a) and DWAA (b). Optical

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microscopy - Weck’s reagent.

Figure 8 - Detail of a wave in the SWAA series. Optical microscopy -

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Weck’s reagent (a) and the Graff-Sargent etch (b).

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Figure 9 - Microhardness indentations (HV0.025) on melted regions of two different waves (a and b) in the SWAA series. Optical microscopy - GraffSargent etch.

Figure 10 - Detail of a wave in the DWAA series. Optical microscopy Weck’s reagent (a) and Graff-Sargent etch (b).

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ACCEPTED MANUSCRIPT Figure 11 - Microhardness indentations (HV0.025) in melted regions of two different waves (a and b) in the DWAA series. Optical microscopy - GraffSargent etch. Figure 12 - Comparison of the wave geometries between the similar (solid

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blue line) and dissimilar (dashed red line) aluminium welds series.

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Figure 13 - Characteristic morphology (a) and detail (b) of the copper-

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aluminium interface. The arrows indicate the cracks in the new phase formed in the welding process. Optical microscopy - Weck’s reagent.

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Figure 14 - Microhardness indentations (HV0.025) in the intermetallic phase

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of two different locations on the interface (a and b). Cu-flyer (above) and Al-base plate (below). Optical microscopy - unetched.

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Figure 15 - Microstructure of the DWCA weld interface, revealed by SEM,

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indicating the location of EDS analyses related to microhardness measurements - unetched.

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Figure 16 - EDS analysis presenting the map of copper in green (a), the

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map of aluminium in red (b) and both maps overlapped (c). Figure 17 - Distribution of the ρR x TR factor on a logarithmic scale.

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ACCEPTED MANUSCRIPT

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Graphical abstract

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