The preparation of TC4/7075 composites for ultra-thin flyer plates is a significant challenge in the field of explosive welding. A weldability window for multi-layer metal explosive welding and a configuration for ultra-thin flyer plate explosive welding were established in this study. TC4/7075 composite materials were successfully prepared with a flyer plate thickness of only 0.3 mm. An analysis was conducted on the material bonding ability, element diffusion, crystal evolution, and microscopic morphology during the explosive welding process of TC4/7075, utilizing the weldability window, molecular dynamics algorithm, and electron backscattered diffraction testing. The results show that various dislocations are present at the interface, predominantly 1/6 ⟨112⟩ dislocations. Element diffusion primarily occurs during the unloading stage at high temperature and zero external pressure; the interface has the best bonding ability when titanium exhibits FCC lattice structure. The prepared composite material demonstrates high intra-grain and grain boundary stresses; the rolling texture is observed on the aluminum side while an equiaxed twin structure forms on the titanium side due to interactions between stacking faults, Stair-rod dislocations, and Hirth immovable dislocations.

Explosive welding (EXW) is a special manufacturing technology for layered or tubular composite materials that utilizes explosions to cause high-speed collisions between metals and achieve high-strength metallurgical bonding.1 Currently, it has been able to achieve large-scale and high-strength metallurgical bonding of dozens of dissimilar metals.1–4 The composite materials obtained have found extensive applications in the fields of aerospace, marine vessels, chemical engineering, and medical equipment.5–7 The EXW of ultra-thin flyer plates is a special technique that involves achieving a thickness of flyer plate less than 0.5 mm. This process allows for the full utilization of the surface properties of flyer plate material, such as conductivity, wear resistance, and corrosion resistance, while maintaining the excellent mechanical properties of the baseplate. Therefore, EXW of ultra-thin flyer plates can also be considered a unique metal coating process. Different from traditional explosive welding, the combination of ultra-thin flyer plate and baseplate is very difficult, the “ultra-thin” characteristics of the flyer plate make it susceptible to tearing and breaking when subjected to the collision force generated by detonation, and the explosive heat generated by explosive will also melt or burning loss the flyer plate. In addition, in the ultra-thin state, the material is far from meeting the rigidity requirements of bearing explosives without deformation, and the bending of the flyer plate leads to inconsistent clearance between the flyer plate and the baseplate, further increasing the difficulty of bonding.8,9

As titanium alloy is lightweight, corrosion resistant, and has a complex smelting process, it possesses the characteristics of high strength and durability.10 Aluminum alloy is also an important lightweight metal material with low density, mature smelting technology, and excellent thermal conductivity.11 Ti/Al laminated composite metal can complement the advantages of the two and has important research value in the field of lightweight equipment manufacturing. Fang et al.,12 Xia et al.,13 and Paul et al.14 studied the influence of explosive welding parameters on the interface microstructure from the perspective of interface bonding morphology. Chulist et al.15 and Lazurenko et al.16 further studied the interface microstructure and found that the typical rolling texture of aluminum after explosive welding and the phenomenon of titanium solid solution in the FCC-Al. Yan et al.17 and Wu et al.18 also used numerical simulation methods to study phenomena such as interface jet flow and waveform growth in the explosive welding process of Ti/Al. Mahmood et al.,19 Xunzhong et al.,20 and Foadian et al.21 detected the formation of various Ti/Al intermetallic compounds such as TiAl, TiAl2, TiAl3, and Ti2Al5 during explosive welding of Ti/Al and studied the growth mechanism of interfacial compounds by heat treatment. In addition to the preparation of single-layer Ti/Al composites, Paul et al.14 and Lazurenko et al.16 also carried out the preparation and testing of multi-layer Ti/Al composites. Scholars have also conducted extensive research on the application properties of Ti/Al composites. Kahraman et al.22 conducted tensile strength test and corrosion resistance test on the prepared TC4/Al composites and found that with the increase in explosive load, the corrosion resistance of the composite gradually decreased. Liang23 and Pei,24 respectively, explored the causes of crack failure of Ti/Al composites and came to the conclusion that stress concentration caused by brittle intermetallic compounds was the main cause of cracks. Kotyk and Boroński studied the influence of explosive welding parameters on the labor properties of Ti/Al composites,25 and Ege and Inal tested the heat resistance of Ti/Al composites.26 With the rapid development of Ti/Al composite preparation technology, scholars have higher and higher expectations on the mechanical properties of Ti/Al alloy EXW composite. The high-quality combination of Ti alloy TC4, which possesses the best mechanical properties among titanium alloys, and aluminum alloy 7075, which possesses the best mechanical properties among aluminum alloys, has become a new hot topic in the study of EXW for titanium and aluminum. However, compared with other Ti/Al composites, the preparation of TC4/7075 is extremely difficult. The abundant magnesium and zinc elements in 7075 are prone to ablative defects at high temperatures, resulting in stress concentration and reducing the strength of the material.27 The high strength and hardness of 7075 also make its welding window narrow and difficult to achieve high-quality composite.28 

The EXW of dissimilar metals is essentially the result of the interaction between atoms of interfacial materials, and the atomic behavior and structure of interfacial materials are the key factors affecting the bonding quality of materials. In order to achieve high-quality bonding of refractory metals, molecular dynamics (MD) algorithm has been applied in the welding field.29 Hao et al.,30 Yan et al.,31 Evstafiev et al.,32 and Wang et al.33 used molecular dynamics algorithms to study welding processes such as expansion welding, laser welding, and ultrasonic welding. Chen et al.34 introduced molecular dynamics algorithms to calculate the EXW process and observed the diffusion behavior of interface atoms during the EXW process. Zhang et al.35 and Ma et al.36 further used molecular dynamics algorithms to study the form of element diffusion of specific material combinations and the main factors affecting element diffusion. Other scholars have also studied the behavior of grain growth37 and atomic strain38 during EXW and concluded that pressure welding and diffusion welding are the main mechanisms of metal bonding in EXW.39 The application of molecular dynamics algorithms provides an important theoretical reference for studying EXW processes and improving EXW conditions. However, the existing molecular dynamics algorithms have limited research on the lattice evolution process of homogeneous and heterogeneous materials such as Ti, and insufficient research on the mechanical properties of materials after bonding so that the molecular dynamics calculation results cannot directly correspond to the bond quality of materials.

In order to avoid force damage and thermal melting of ultra-thin flyer plate, a novel configuration for explosive welding of ultrathin flyer plate is proposed, which involves an additional force transition layer and insulation layer. Furthermore, to enhance the bonding strength between the flyer plate TC4 and baseplate 7075, a weldability window for EXW of multi-layer metal is established, providing a theoretical foundation for the selection of interlayer materials and the use of interlayer. In addition, the paper employs the molecular dynamics algorithm to calculate the EXW behavior of Ti/Al and successfully prepares TC4/7075 composite materials with a flyer plate thickness of only 0.3 mm, based on the multi-layer metal EXW weldability window and the new configuration of ultra-thin flyer plate explosive welding. Based on molecular dynamics algorithm and scanning electron microscope (SEM), energy dispersive spectrometry (EDS), and electron backscattered diffraction (EBSD) characterization techniques, the element diffusion, lattice evolution, dislocation growth and annihilation, and grain morphology of Ti/Al explosive welding were studied, which provides a reference for the manufacture of other metal materials that are difficult to weld.

The arrangement of parallel explosive welding is shown in Fig. 1. Under the instantaneous high pressure generated by explosive explosion, the flyer plate has plastic deformation and crashes into the baseplate at a certain speed and tilt angle. Under the comprehensive action of high temperature and high pressure generated by the collision, the flyer plate and the baseplate are combined.

FIG. 1.

Schematic diagram of parallel explosive welding.

FIG. 1.

Schematic diagram of parallel explosive welding.

Close modal

According to the principle of explosive welding bonding, a theoretical model shown in Fig. 2 is proposed, which includes three key parameters: collision velocity Vp, collision angle β, and welding velocity Vc.

FIG. 2.

Theoretical model of parallel explosive welding.

FIG. 2.

Theoretical model of parallel explosive welding.

Close modal
According to the geometric relationship shown in Fig. 2, the relationship between collision velocity Vp, collision angle β, and welding velocity Vc can be deduced [Eq. (1)],
V c = 2 V p sin ( β / 2 ) .
(1)
In parallel explosive welding, the welding speed Vc of the flyer plate and the baseplate is equal to the explosive detonation speed Vd. With the increasing demand for higher performance of composite materials, the realization of high strength bonding between more matrix materials with excellent mechanical properties and application value has become the primary problem faced by explosive welding technology. When analyzing the bonding ability of materials, the concept of weldability window based on welding speed and collision angle was proposed.40 

1. Upper limit of weldability window

With the increase in explosive thickness or intensity, the explosive welding system gets more energy input, and the flyer plate also has a more obvious tilt deformation and a greater collision speed. High collision speed leads to higher pressure and temperature generation; severe plastic deformation occurs at the contact interface between the flyer plate and the baseplate; large-scale melting, material burning, and high temperature oxidation occur at the interface; and then defects such as continuous melting zone, ingot structure, and vortex are generated, which affect the bonding quality and the performance of the composite material. Carpenter et al.41 analyzed the form of energy flow in the explosive welding process and found that the chemical energy generated by explosives was partially absorbed by the flyer plate in the form of kinetic energy, and the kinetic energy obtained by the flyer plate was converted into heat energy and plastic deformation energy of metals during the collision process. In order to achieve the optimal bonding between the flyer plate and the baseplate, the welding interface after the collision point must be fully cooled so that the welding interface is not pulled apart by the reflected tensile waves to form a continuous non-bonding area. This determines the upper limit of the weldability window for explosive welding [Eq. (2)],
sin ( β 2 ) = ( T m C 0 ) 1 / 2 2 N V c 2 ( k C P C 0 ρ f h ) 1 / 4 ,
(2)
where Tm is the melting point of the lower melting point metal in the flyer and baseplate materials, k is the thermal conductivity, Cp is the constant pressure heat capacity, C0 is the volume sound velocity, ρf is the density of the flyer plate, h is the thickness of the flyer plate, and N is a constant which is generally taken as 0.11. Due to the fact that the input energy of the system is the kinetic energy obtained by the flyer plate, except for the lowest melting point Tm of the reaction interface material melting state, all other parameters used in Eq. (2) come from the flyer plate.

2. Lower limit of weldability window

Deribas and Zakharenko42 pointed out that the metal jet formed by the collision between the flyer plate and the baseplate is a necessary condition for achieving high-quality bonding. Based on the fluid dynamics theory, they gave the lower limit boundary of the weldability window [Eq. (3)],
V c = 1.14 β H v ρ a v e ,
(3)
where ρave is the arithmetic average of the density of the explosive welded baseplate and the flyer plate, because both interfaces need to be softened and cleaned by the jet stream, Hv represents the hardness value of the harder material in both the flyer plate and the baseplate. Zakharenko and Zlobin43 conducted further research based on Eq. (3) and found that in explosive welding of dissimilar metals with significant differences in hardness, when the hardness of materials with higher hardness changes by a factor of 5, the collision angle value within the weldability window changes by 25%. In contrast, the material with less hardness in baseplate and flyer plate is the main reason for affecting the bonding quality. As long as the softer material is subjected to severe plastic deformation, enough metal jets can be generated. The jet washes the impurities and stains on the metal surface, which causes the surface oxide film of the welded metal to break and inhibit its reproduction. In addition to the cleaning effect of jets on the welding surface, the original cleanliness, flatness, and welding vacuum of the flyer plate and baseplate materials also affect the welding quality.40 According to the above considerations, Eq. (3) is further improved into the form shown in Eq. (4),
V c = k 1 β H v ρ ,
(4)
where ρ is the density of the softer material in the flyer plate material and the baseplate material, k1 is a constant related to the cleanliness, flatness, and welding vacuum of the material to be welded surface, and its value ranges from 0.6 to 1.2. When the two surfaces to be welded reach high cleanliness and high flatness, k1 is 0.6. When the surface to be welded is completely uncleaned, k1 is 1.2. In fact, the effect of k1 on the weldable window is huge, and a slight numerical change can cause a significant change in the lower limit of the weldable window.

3. Left limit of weldability window

The left limit of the weldability window in explosive welding is associated with the formation of the waveform interface. Cowan et al.44 elucidated the formation of the interface waveform based on the Karman vortex street principle of fluid flowing through a vortex generated by a fixed obstacle. Through an analysis of the typical wavy interface in explosive welding, Cowan predicted the minimum collision velocity required for the formation of the interface waveform in the process of explosive welding. This led to the establishment of a quantitative relationship between collision velocity and collision angle [Eq. (5)],
R e = ( ρ f + ρ b ) V c 2 2 ( H f + H b ) ,
(5)
where Re is the Reynolds number, the value is between 8 and 13; Cowan gives the best Reynolds number is 10.6; ρf and ρb are the density of the flyer plate and the baseplate, respectively. Hb and Hf are the Vickers hardness of the baseplate and the flyer plate, respectively.

4. Right limit of weldability window

Cowan and Holtzman45 further analyzed the interface state during explosive welding using the non-viscous fluid theory. He noted that when the welding speed is lower than the velocity of shock wave transmission, jets are generated and combined after material collision. However, when the welding speed exceeds the volume velocity of the material, an expansion wave is produced, preventing the formation of jets and the realization of combination at the welding interface. Therefore, the collision velocity Vc needs to meet Eq. (6),
V c < min ( C f , C b ) ,
(6)
where Cf and Cb are the volume sound velocity of the flyer plate and the baseplate, respectively. The welding speed should be lower than the smaller value of the volume sound velocity of the flyer plate and the baseplate. This conclusion was also reached by Walsh et al.46 The volume sound velocity of metal materials can be calculated using Eq. (7),
C 0 = ( G ρ ) 1 / 2 ,
(7)
where G is Young's modulus of the metal.

Zamani and Liaghat40 standardized the range of collision angles for explosive welding. In order to achieve an appropriate jet state without excessive melting of the interface, the collision angle is typically maintained between 2° and 31°. In conclusion, the traditional parallel explosive welding weldability window can be established.

The parameters related to the weldability window of typical Ti/Al metal materials are presented in Table I. Considering a flyer plate thickness of 3 mm, a baseplate thickness of 5 mm, and k1 as 0.85 (the value under normal circumstances), the weldability window values for various Ti/Al explosive welding scenarios can be calculated as depicted in Fig. 3 (with titanium alloy serving as the flyer plate and aluminum alloy as the baseplate).

FIG. 3.

Weldability window of explosive welding of several Ti/Al alloys, (a) TA2-1060 weldability window, (b) TA2-5083 weldability window, (c) TA2/7075 weldability window, (d) TC4/1060 weldability window, (e) TC4/5083 weldability window, and (f) TC4/7075 weldability window.

FIG. 3.

Weldability window of explosive welding of several Ti/Al alloys, (a) TA2-1060 weldability window, (b) TA2-5083 weldability window, (c) TA2/7075 weldability window, (d) TC4/1060 weldability window, (e) TC4/5083 weldability window, and (f) TC4/7075 weldability window.

Close modal
TABLE I.

Parameters related to the welding window of Ti/Al explosive welding.

Density (g/cm3)Hardness (Hv)Tensile strength (MPa)Melting point (°C)Young's modulus (GPa)Volumetric sound velocity (m/s)Thermal conductivity (W/m°C)Specific Heat capacity (J/kg °C)
TA2 4.51 250 441 1660 100 4709 15 550 
TC4 4.51 316 1012 1670 110 4938 7.955 612 
Al1060 2.68 32 120 660 69 5074 237 880 
Al5083 2.80 87 280 638 72 5070 156 947 
Al7075 2.81 155 560 635 71 5026 173 960 
Density (g/cm3)Hardness (Hv)Tensile strength (MPa)Melting point (°C)Young's modulus (GPa)Volumetric sound velocity (m/s)Thermal conductivity (W/m°C)Specific Heat capacity (J/kg °C)
TA2 4.51 250 441 1660 100 4709 15 550 
TC4 4.51 316 1012 1670 110 4938 7.955 612 
Al1060 2.68 32 120 660 69 5074 237 880 
Al5083 2.80 87 280 638 72 5070 156 947 
Al7075 2.81 155 560 635 71 5026 173 960 

The weldability window of aluminum alloys 1060, 5083, and 7075 gradually narrows when the flyer plate material remains constant, as illustrated in Fig. 3. Further examination of the weldability window depicted in Fig. 3 reveals that the lower limit depends on the softer materials present in both the flyer plate and baseplate during explosive welding. Under the condition that the hardness of the aluminum alloy baseplate is lower than that of the titanium alloy flyer plate, the hardness of the aluminum alloy baseplate predominantly influences the lower limit of the weldability window. As the hardness of the aluminum alloy increases, the challenge of generating a metal jet at the interface also increases, necessitating an improvement in the collision speed of the flyer plate to achieve the objective of cleaning the welding surface. In the explosive welding of Ti/Al with TA2 as the flyer plate, the lower limit of the weldability window for aluminum alloy 7075 is notably higher than that of 1060 and 5083. The tensile strength of 7075 exceeds that of the flyer plate TA2, the heightened tensile strength of the aluminum alloy makes it more challenging to surpass the material's dynamic yield limit, necessitating a higher collision velocity to ensure the generation of an interface jet.

When comparing the weldability windows of TC4 and TA2 flyer plates with the same baseplate, it is evident that the weldability window of Ti/Al explosive welding with TC4 as the flyer plate is narrower. The exceptionally high tensile strength of TC4 directly contributes to an increase in the minimum welding speed required for explosive welding, resulting in a higher left limit of the weldability window compared to that with TA2 as the flyer plate. Additionally, the upper limit of the weldability window for TC4/7075 is notably lower than that of TA2/7075. The narrow weldability window poses challenges in achieving a high-quality combination of TC4/7075.

The calculation of the weldability window of explosive welding shows that the excellent mechanical performance of TC4/7075 leads to the extremely narrow weldability window of TC4/7075, and it is difficult to achieve high-quality combination. In order to solve this problem, an interlayer can be introduced to solve the problem that TC4/7075 is difficult to weld. Figure 4 shows the schematic diagram of explosive welding experiment with interlayer. Different from the traditional explosive welding arrangement, a layer of metal plate is added between the baseplate and the flyer plate. The ultimate goal of welding is still to prepare the layered composite metal based on the flyer plate and the baseplate, but a force transition layer is added between the baseplate and the flyer plate to improve the difficult situation of welding the flyer plate and the baseplate. The interlayer arrangement can make the flyer plate and baseplate materials with great differences in physical and chemical properties find a transition so that the differences between the flyer plate and the interlayer, the interlayer and the baseplate in the melting point, thermal conductivity, thermal deformation coefficient, and other physical properties are smaller than the differences between the flyer plate and the baseplate related physical and chemical properties. When explosive welding results in severe deformation due to high temperature and high pressure, the interface deformation after the introduction of interlayer is more coordinated, thus reducing the phenomenon of interface over melting and non-bonding caused by huge deformation differences. According to the literature,18 numerical simulations and energy distribution analysis demonstrate that the introduction of the interlayer effectively transforms the original primary collision molding into secondary collision molding. The detonation energy generated by explosives is distributed at the flyer plate–interlayer interface and the interlayer–interlayer interface, respectively. The study also confirms that the introduction of the interlayer can efficiently control the maximum energy on a single interface and enhance energy utilization efficiency throughout the explosive welding process. While this research offers a reasonable theoretical explanation for the enhancement principle of interlayer strength in explosive welding, it does not provide a specific quantitative basis for the selection of interlayer materials. Currently, the selection of interlayer materials in explosive welding primarily relies on empirical knowledge and repeated experiments, leading to significant contingency and uncertainty.

FIG. 4.

Schematic diagram of explosive welding with interlayer.

FIG. 4.

Schematic diagram of explosive welding with interlayer.

Close modal

The difficulty of welding composite materials can be directly reflected through the weldability window, so as long as the explosive welding weldability window after the use of interlayer can be established, the selection of interlayer can be quantitatively analyzed. As pointed out above, the use of interlayer is actually a one-time welding molding into multiple explosive welding molding. By ensuring a high-quality combination between the flyer plate and interlayer, as well as the interlayer and baseplate, it is possible to guarantee the overall quality of the composite material. The overlap area of flyer plate–interlayer and interlayer–baseplate weldability window is the overall weldability window after using interlayer. Based on the analysis of the weldability window characteristics of explosive welding, it can be observed that the determination of interlayer mainly considers the following aspects: (1) The interlayer material significantly impacts the overall mechanical properties, including strength, stiffness, density, and other physical and chemical characteristics of composite materials. The primary objective of incorporating an interlayer is to achieve composite materials with enhanced bonding quality while preserving their application value. (2) It is advisable to opt for metal materials belonging to the same family as the baseplate or flyer plate, as metals within the same group typically exhibit similar ductility and physical and chemical properties. This facilitates better coordination of stress and deformation during explosive welding preparation and subsequent application, thereby mitigating the generation of internal stress and microscopic defects resulting from non-uniform deformation that could potentially impact material performance. (3) From the perspective of weldability window, it is advisable for the interlayer to possess a relatively lower hardness. A reduced hardness facilitates improved jetting capability at interfaces where interlayer material washes over welding surfaces, thereby augmenting the upper limits within the weldability window. (4) The tensile strength of interlayer materials can be appropriately reduced as it primarily affects the left limit of explosive welding windows; having a lower tensile strength than both the baseplate and flyer plate ensures that this lower limit does not shift toward higher positions. (5) The volume sound velocity of the interlayer material should not be excessively low, as it may cause a leftward shift in the right limit of the weldability window.

Based on the aforementioned justifications, aluminum alloy 1060, which belongs to the same family as the baseplate, was chosen as the interlayer to optimize the explosive welding conditions of TC4/7075. The establishment of a weldability window for multi-layer metal explosive welding requires a paradigm shift in understanding the concepts of “flyer plate” and “baseplate.” To calculate the weldability window for the “flyer plate–interlayer” system, it is relatively straightforward to consider the interlayer as an explosive welding baseplate at this stage and incorporate relevant parameters from both the flyer plate and interlayer into Eqs. (2)–(6).

The construction of the weldability window for the “interlayer–baseplate” system is relatively intricate. The current flyer plate refers to a novel composition of both flyer plate and interlayer. Therefore, when establishing the weldability window for the interlayer baseplate, it is crucial to select relevant parameters judiciously. In Eq. (2) for calculating the upper limit of the weldability window, Tm should encompass the value of the lower melting point within the interlayer–baseplate structure. Additionally, h and ρ are pertinent parameters that reflect energy absorption by the flyer plate; thus, h should be substituted with a sum of thickness values comprising both flyer plate and interlayer layers. Furthermore, density denotes an arithmetic average density derived from both flyer plate and interlayer components as illustrated in Eq. (8),
ρ f i = ρ f h f + ρ i h i h f + h i .
(8)

The other parameters in Eq. (2) reflect the heat conduction ability of the interface, so they are replaced by the relevant parameters of the interlayer material. When calculating the lower limit of the weldability window, HV takes the hardness value of the lower hardness in the interlayer and the baseplate, and the density ρ is the arithmetic average density of the flyer plate and the interlayer. The right limit of the weldability window is determined by selecting the smaller value of volume sound velocity in the interlayer and baseplate.

Figure 5 illustrates the weldability window of TC4/1060/7075. Region II represents the weldability window of flyer plate–interlayer after interlayer usage, while region III is the weldability window of interlayer–baseplate. The overlapping area between regions II and III indicates the weldability window of TC4/1060/7075 explosive welding with interlayer, which is similar in size to region II. Region I displays the weldability window of TC4/7075 without interlayers. As shown in Fig. 5, using an interlayer significantly expands the weldability window for TC4/7075 explosive welding compared to not using one. This theoretical proof demonstrates that pure aluminum 1060 can enhance the weldability of TC4/7075 explosive welding.

FIG. 5.

Weldability window of TC4/1060/7075.

FIG. 5.

Weldability window of TC4/1060/7075.

Close modal

As previously mentioned, the preparation of ultra-thin flyer plate explosive welding poses several challenges: (1) The thickness of the flyer plate restricts its ability to acquire sufficient kinetic energy from the explosive, leading to ineffective collision with the baseplate and inadequate metal jet formation at the interface, impacting the lower limit of the weldability window. (2) The ultra-thin flyer plate's low material strength makes it susceptible to being easily torn apart by explosive forces. (3) With the exception of materials with high melting points, most ultra-thin flyer plate materials are prone to melting into clumps or vaporization due to the heat generated by the explosion. (4) Ultra-thin flyer plates are prone to deformation, affecting their flatness during explosive welding field experiments. To address these challenges, a new configuration for the explosive welding experiment of TC4/7075 ultra-thin flyer plate can be implemented, as depicted in Fig. 6. Pure aluminum 1060 continues to be used as the interlayer to enhance the welding conditions between the flyer plate and the baseplate. Instead of the traditional explosive welding setting, an insulation layer for absorbing and isolating the explosion heat and a force transmission layer for conducting the detonation pressure to guide the deformation of the flyer plate are successively added to the top of the flyer plate. The force transmission layer typically employs a metal plate, and the bending deformation of the metal plate drives the deformation of the flyer plate, while also providing preliminary heat insulation. The insulation layer utilizes a thin, insulated, easily burned paper board to quickly dissipate residual explosive heat passing through the force transmission layer. In the experiment setting, the force transmission layer, the insulation layer, and the flyer plate are closely fitted to prevent excessive air inclusion and ensure the maximum flatness of the flyer plate, thereby improving bonding quality.

FIG. 6.

Explosive welding configuration of ultra-thin flyer plate.

FIG. 6.

Explosive welding configuration of ultra-thin flyer plate.

Close modal

In explosive welding, the force transmission layer solely participates in detonation force transmission and does not bond with either the flyer plate or baseplate. Consequently, the choice of material for the force transmission layer is more flexible, enabling the selection of superior materials based on weldability windows. Revised sentence: In accordance with the weldability window calculation method for interlayer structures, attention should also be given to parameter selection when calculating the weldability window of explosive welding for ultra-thin flyer plates. The “flyer plate–interlayer” weldability window calculation involves a “new flyer plate” consisting of a force transmission layer, insulation layer, and flyer plate; while the “interlayer–baseplate” weldability window calculation considers a “new interlayer” composed of a force transmission layer, residual insulation layer, flyer plate, and interlayer. Table II presents relevant parameters for low carbon steel Q195 and pure copper T1 during the weldability window calculation.

TABLE II.

Related parameters of material weldable window.

Density (g/cm3)Hardness (Hv)Tensile strength (MPa)Melting point (°C)Young’s modulus (GPa)Volumetric sound velocity (m/s)Thermal conductivity (W/m°C)Specific heat capacity (J/kg°C)
Q195 7.85 120 400 1350 198 5022 46.8 502 
T1 8.94 110 325 1083 200 4729 380 390 
Density (g/cm3)Hardness (Hv)Tensile strength (MPa)Melting point (°C)Young’s modulus (GPa)Volumetric sound velocity (m/s)Thermal conductivity (W/m°C)Specific heat capacity (J/kg°C)
Q195 7.85 120 400 1350 198 5022 46.8 502 
T1 8.94 110 325 1083 200 4729 380 390 

Figure 7 shows the weldability windows of 0.3 mm TC4 + 1 mm 1060 + 5 mm 7075 explosive welding using pure aluminum 1060, low carbon steel Q195, and pure copper T1 as the force transmission layer, respectively. The coincidence area between region I and region II is the weldability window for explosive welding of the ultra-thin flyer plate with the new configuration. As can be seen from Figs. 7(a) and 7(b), since the densities of iron and copper are relatively close, the weldability window area with the two as the force transmission layer is basically the same. The density of the two is significantly higher than that of aluminum plate, so the upper and lower limits of the ultra-thin TC4/7075 weldability window with Q195 and T1 as the force transmission layer are significantly lower than the upper and lower limits of the weldability window with pure aluminum as the force transmission layer. Literature47 points out that when welding parameters are taken around the lower limit of the weldability window, the welding interface is more likely to generate a flat or wavy shaped interface with strong bonding force and low melting defects. Therefore, in the ultra-thin TC4/7075 explosive welding experiment, the choice of Q195 or T1 is slightly better than the choice of pure aluminum plate 1060.

FIG. 7.

Weldability window of explosive welding configuration of ultra-thin flyer plate, (a) 1060 force transmission layer, (b) Q195 force transmission layer, and (c) T1 force transmission layer.

FIG. 7.

Weldability window of explosive welding configuration of ultra-thin flyer plate, (a) 1060 force transmission layer, (b) Q195 force transmission layer, and (c) T1 force transmission layer.

Close modal

The weldability window shown in Fig. 5 has indicated that during TC4/1060/7075 explosive welding, the interface that is difficult to bond is mainly the interface between Ti flyer plate and Al interlayer. Therefore, the molecular dynamics algorithm can be used to analyze the evolution law of complex phase state, crystal structure at the interface of Ti/Al explosive welding and further determine the bonding mechanism of Ti/Al explosive welding. The open source software LAMMPS was used for molecular dynamics calculation.48 The initial molecular dynamics model is shown in Fig. 8. Titanium is a close-packed hexagonal (HCP) structure, aluminum is a face-centered cubic (FCC) structure, the system size is 129.6 × 116.85 × 295 Å (x × y × z), and the system has a total of 523 460 atoms. The x and y directions are periodic boundaries, and the z direction boundary conditions change with the simulation stage. In order to adjust the movement of the atomic region and eliminate the influence of stress waves, the three thicknesses of atomic layers at the z direction boundary of the Ti/Al atomic region were set as the transition region. The EAM metal potential function developed by Zope and Mishin49 was used to describe the interaction force between Ti/Al atoms, and the collision crystal planes were set as Al (001) and Ti (0001) planes. The molecular dynamics simulation of explosive welding is divided into three stages:34 the first stage is the loading stage, in which the Al atom transition region is fixed, and the Ti atom region is crashed into the Al atom region at a certain speed, and the Ti atom region is fixed when the system volume is compressed to the minimum, and then relaxation by 2 ns. In order to achieve the purpose of comparative analysis, three initial collision velocities of 1750, 2250, and 2750 m/s are set, respectively. The second stage is the unloading stage, in which the system is maintained at the temperature at the end of the loading stage and relaxes 2 ns under 0 external pressure. The third stage is the cooling stage, where the external pressure of the system is maintained at 0, and the system temperature is lowered to room temperature at a rate of 1012 °C/s. Relaxation is carried out at room temperature for 2 ns until the system is completely cooled. The molecular dynamics simulation of explosive welding was completed in the above three stages. In order to further compare the bonding strength of composite materials under different parameters, after the system was fully cooled, the aluminum plate transition region was fixed, and the composite test block was stretched along the z axis at a speed of 0.3 m/s, and the stress–strain curve during the stretching process was output in the form of metal tensile experiment. At this time, the system is a shrinkable boundary in the x, y, and z directions. The common neighbor analysis (CNA) algorithm was used to observe the crystal type change of the system, the radial distribution function (RDF) was used to analyze the phase state of the system, and the mean square displacement (MSD) was used to calculate the diffusion capacity of different atoms. The atomic coordinate information and atomic distribution images were integrated to analyze the atomic distribution and diffusion capacity of Ti/Al in the explosive welding process.

FIG. 8.

Initial molecular dynamics model: (a) the atomic type of the system and (b) the distribution of lattice type of the system.

FIG. 8.

Initial molecular dynamics model: (a) the atomic type of the system and (b) the distribution of lattice type of the system.

Close modal

The explosive welding experiment was set up according to the form shown in Fig. 6, Q195 was selected as the force transmission layer, and the composition and specifications of the materials used in the experiment were shown in Table III. Soft sandy land is selected as the foundation, and non-woven fabric is placed above to buffer the impact force of the composite board and avoid cracking of the composite board. A pure copper sheet folded into a “V” shape is placed between the flyer plate–interlayer and interlayer–baseplate used to control the gap.

TABLE III.

Chemical composition and specifications of experiment materials.

MaterialsChemical composition (wt. %)Length × width × thickness (mm)
Ti6Al4V Al Fe Ti 100 × 100 × 0.3 
0.05 0.08 0.01 0.3 0.2 Residual 
1060 Fe Ti Si Mn Zn Mg Cu Al 100 × 100 × 1 
0.35 0.03 0.25 0.03 0.05 0.03 0.05 Residual 
7075 Zn Mg Cu Si Fe Mn Cr Ti Al 100 × 100 × 5 
2.9 2.0 0.3 0.5 0.3 0.28 0.2 Residual 
Q195 Mn Si Fe … … 100 × 100 × 1 
0.1 0.5 0.3 0.04 0.03 Residual … … 
MaterialsChemical composition (wt. %)Length × width × thickness (mm)
Ti6Al4V Al Fe Ti 100 × 100 × 0.3 
0.05 0.08 0.01 0.3 0.2 Residual 
1060 Fe Ti Si Mn Zn Mg Cu Al 100 × 100 × 1 
0.35 0.03 0.25 0.03 0.05 0.03 0.05 Residual 
7075 Zn Mg Cu Si Fe Mn Cr Ti Al 100 × 100 × 5 
2.9 2.0 0.3 0.5 0.3 0.28 0.2 Residual 
Q195 Mn Si Fe … … 100 × 100 × 1 
0.1 0.5 0.3 0.04 0.03 Residual … … 

Quartz sand and wood chips were added to the pure ammonium nitrate explosive to control the explosive detonation speed of 2200 m/s and the density of 0.8 g/cm3. After placing the flyer plate, insulation layer, and force transmission layer, 15 mm thickness of explosives are laid so that the experiment is located in the weldability window determined above. Because the doped quartz sand and wood chips reduce the initiation ability of the explosive, pure explosive can be appropriately laid around the detonator to detonate in the actual experiment. After the experiment, the test sample was prepared at a suitable location. Struers Tegramin-25 was used to abrade the sample, Struers CitoPress was used to insert the sample, Zeiss-Axio Imager 2 was used to observe the sample, and the microscopic morphology at the bonding interface was analyzed. The element scanning of the composite plate prepared by the test was carried out to analyze the element diffusion at the interface between flyer plate and interlayer. The EBSD test was performed on the sample to observe the microstructure and atomic arrangement at the interface.

1. Element diffusion

At the relaxation stage, the FCC lattice and HCP lattice stabilized at 51.8% and 46.8%, respectively, indicating that the molecular dynamics model was reasonably constructed and the system reached a stable state. The atomic distribution at the Ti/Al interface at the end of each stage of the molecular dynamics simulation is shown in Fig. 9. As depicted in Fig. 9(a), when the collision velocity is 1750 m/s, there is no significant diffusion of interface atoms at the end of the three stages. When the collision velocity is 2250 m/s [see Fig. 9(b)], no obvious atomic diffusion can be observed at the interface at the end of the loading stage. At the end of the unloading phase, a substantial number of Ti atoms diffuse into the Al atomic region. Moreover, at the end of the cooling stage, the diffusion of Ti atoms into the Al atomic region further increases. When the collision speed is 2750 m/s [see Fig. 9(c)], the loading stage ends at the same speed as the other two collisions, and no obvious element diffusion can be observed at the interface. At the end of the unloading stage, in addition to a large amount of Ti atoms diffusing to the Al atomic region, there is also a small amount of Al atoms diffusing to the Ti atomic region. The degree of mutual diffusion between titanium and aluminum further increases at the end of the cooling stage. Element diffusion is a typical behavior of interfacial atoms in explosive welding. On the one hand, element diffusion can generate a continuous intermetallic compound layer with a certain strength at the interface, and the continuous and stable intermetallic compound layer is one of the important sources of the bonding strength between the flyer plate and the baseplate. On the other hand, for some metals, some intermetallic compounds generated by element diffusion exhibit obvious brittleness, and discontinuous brittle intermetallic compounds become microscopic defects at the bonding interface, which is prone to cause stress concentration and reduce the mechanical properties during the use of materials. According to the analysis of Fig. 9, the atomic diffusion thickness at the interface gradually increases with the increase in collision velocity at the end of explosive welding. Combined with Figs. 9(b) and 9(c), it can be observed that element diffusion mainly occurs in the unloading stage at high temperature and zero external pressure, and the main form of atomic diffusion in the explosive welding process of Ti/Al is the diffusion of titanium atoms to the Al atomic region.

FIG. 9.

Distribution of atoms in three stages with different collision velocities, (a) 1750, (b) 2250, and (c) 2750 m/s.

FIG. 9.

Distribution of atoms in three stages with different collision velocities, (a) 1750, (b) 2250, and (c) 2750 m/s.

Close modal

2. Crystal structure and dislocation evolution

Figures 10–12 show the CNA results of the system at the end of each stage under three collision velocities. In the loading stage (see Fig. 10), the volume of the system is severely compressed, and the temperature and pressure of the system rise rapidly. Under the three collision speeds, the lattice structure of the Ti/Al atomic region changes significantly. The lattice structure change of Al atoms is relatively simple. With the increase in the collision speed of Ti atomic region, the volume of the system is gradually compressed violently, and more and more FCC-Al is transformed into a higher-density body-centered cubic (BCC) structure. In contrast, the lattice transformation of the Ti atomic region is more complex. When the collision velocity is 1750 m/s, the atomic region of titanium is α-Ti (HCP) and β-Ti (BCC) co-existing. When the collision velocity is 2250 m/s, α-Ti at room temperature basically disappears, and the titanium transforms into a high-temperature β-Ti and FCC phase coexistence state, and the body-centered cubic structure becomes the main structure of titanium. When the collision velocity continues to increase to 2750 m/s, most titanium is transformed into a FCC structure. FCC-Ti is a metastable phase, which is often found after severe plastic deformation of titanium and its alloys, so the appearance of FCC-Ti can also prove that serious plastic deformation occurs in the atomic region of titanium. Although the temperature of the system increases sharply in the loading stage, the movement of atoms is strongly limited under the collision pressure, so titanium and aluminum all appear as metal crystal states with an ordered crystal lattice under the three collision speeds in the loading stage. Due to the compression of atomic distance, the movement of atoms to produce effective displacement is limited (under the influence of collision, atoms vibrate violently without deviating too much from the position, which is physically manifested as the temperature of the system increases), so element diffusion can hardly be observed in the loading stage, which verifies the analysis of element diffusion in the loading stage mentioned above.

FIG. 10.

CNA and DXA results of the system after loading. In the CNA results, the green area is FCC lattice structure, the blue area is BCC lattice structure, and the red area is HCP lattice structure. In the DXA results, the blue line represents a 1/2⟨110⟩(Perfect) dislocation, the green line represents a 1/6⟨112⟩(Shockley) dislocation, and the purple line represents a 1/6⟨110⟩(stair-rod) dislocation. The yellow line represents a 1/3⟨110⟩(Hirth) dislocation, while the blue line represents a 1/3⟨111⟩(Frank) dislocation.

FIG. 10.

CNA and DXA results of the system after loading. In the CNA results, the green area is FCC lattice structure, the blue area is BCC lattice structure, and the red area is HCP lattice structure. In the DXA results, the blue line represents a 1/2⟨110⟩(Perfect) dislocation, the green line represents a 1/6⟨112⟩(Shockley) dislocation, and the purple line represents a 1/6⟨110⟩(stair-rod) dislocation. The yellow line represents a 1/3⟨110⟩(Hirth) dislocation, while the blue line represents a 1/3⟨111⟩(Frank) dislocation.

Close modal
FIG. 11.

System CNA and DXA results at the end of the unloading stage.

FIG. 11.

System CNA and DXA results at the end of the unloading stage.

Close modal
FIG. 12.

CNA and DXA results of the system after cooling.

FIG. 12.

CNA and DXA results of the system after cooling.

Close modal

At the end of the loading stage, a large number of various dislocations, primarily consisting of 1/6 ⟨112⟩ (Shockley) dislocations, are generated due to lattice distortion. Shockley dislocations serve as boundaries between areas with and without stacking faults and are formed through local slip in the crystal. At a collision speed of 1750 m/s, multiple parallel stacking faults with a tilt angle of 45° are generated in the Al atomic region, and a limited slip Hirth dislocation is generated at the intersection of stacking faults, further limiting the stress conduction in the Ti atomic region. When the collision velocities are 2250 and 2750 m/s, dislocations are mainly generated in the face centered cubic titanium region, with multiple dislocations intersecting and numerous in number but limited in length.

The unloading stage, as depicted in Fig. 11, corresponds to the brief period after the detonation wave leaves during explosive welding experiments. During this stage, the collision pressure dissipates, and the system volume expands. The release of atomic motion restrictions and an increase in atomic spacing within the system provide the necessary conditions for element diffusion during this stage. Additionally, the phase state of the atoms undergoes changes during this stage. When the collision velocity is 1750 m/s, the temperature increase caused by collisions is still lower than the melting point of titanium or aluminum. At this point, both titanium and Al atomic regions revert to their original crystal lattice types and remain in a metallic crystalline state. Therefore, no significant element diffusion can be observed in this state. Therefore, no significant element diffusion can be observed in this state. As the collision velocity reaches 2250 m/s, the system temperature exceeds that of aluminum's melting point but remains below titanium's melting point. In this scenario, aluminum transitions into a disordered molten state with significantly increased atomic spacing. Titanium atoms with considerable mobility quickly fill in the gaps between Al atoms, resulting in elemental diffusion from titanium to aluminum regions at the interface. The spacing between Ti atoms returns to normal, transforming into a predominantly FCC structure as a metallic crystalline state, while Al atoms are unable to effectively diffuse toward the titanium side. When the collision speed is 2250 m/s, the system temperature is higher than the melting point of titanium. At this stage, both titanium and aluminum are in a molten state, and bidirectional diffusion of titanium/aluminum atoms occurs. However, due to the larger atomic spacing caused by thermal collision of aluminum, the diffusion of titanium atoms toward the aluminum atomic region is more significant.

At the conclusion of the unloading stage, a substantial number of Frank dislocations appeared in the system at a collision velocity of 1750 m/s, and Frank dislocation loops also formed in the Al atomic region. When combined with CNA results, it can be determined that the aggregation and collapse of vacancies in regions with lattice disorder are the primary causes for generating Frank dislocations. Additionally, many pre-existing Shockley dislocations in the Al atomic region were annihilated during dislocation slip, leading to a transformation from HCP structure to FCC structure at corresponding positions and the elimination of stacking faults. At a collision velocity of 2250 m/s, numerous stair-rod dislocations are present at the interfaces between the FCC-Ti and HCP structures. During the loading stage, a significant amount of Shockley dislocations react to form pressure stair-rod dislocations, as shown in Eq. (9). In comparison to the previous stage, some dislocations are annihilated, resulting in an overall decrease in dislocation density, while others proliferate into long-range dislocations,
1 6 [ 12 1 ¯ ] + 1 6 [ 2 ¯ 1 ¯ 1 ] 1 6 [ 1 ¯ 10 ] .
(9)

At the cooling stage, as depicted in Fig. 12, the system temperature returns to room temperature. At a collision velocity of 1750 m/s, there is no significant change in the lattice structure of the system compared to the previous stage, and both titanium and aluminum return to their pre-collision lattice types. At a collision velocity of 2250 m/s, aluminum cools down and reverts back to the FCC crystal state. Due to the melting–cooling process, there are fewer other phases in the Al atomic region at this time, making it more pure in terms of the crystalline phase. Meanwhile, titanium maintains its FCC structure, with an additional HCP phase present at a 45° angle within its atomic region. According to research,50 the presence of a FCC structure in titanium can significantly improve its strength and mechanical properties without reducing its plasticity. Furthermore, if the titanium at the interface is in FCC structure, the bonding strength at the interface can be further improved. However, the phase transition mechanism of FCC-Ti generation is very complex. Ren et al.51 pointed out that the lattice transition from HCP to FCC is caused by the dislocation slip of Shockley partial dislocation under size limitation. The formation of FCC-Ti was also observed when titanium was subjected to GPa high pressure and a high strain rate of 106–107/s, generated on the sample surface by laser shock strengthening equipment at the Northwest China Institute of Non-ferrous Metals. This observation is very similar to the 10 GPa high pressure at the interface and the high strain rate environment of the interface metal during explosive welding.

When the collision velocity is 2750 m/s, titanium changes from the melting state of the previous stage to a metal crystal state dominated by a co-existence of HCP structure and FCC structure.

After the cooling of the system, a small number of Shockley dislocation rings appeared in the Al atomic region when the collision velocity was 1750 m/s. Due to cooling, the contraction of atomic spacing generates a large number of vacancy defects, which further leads to lattice distortion. Lattice distortion breaks the atomic bonds causing collapse; the vacancies between the twin boundaries come together to form the Shockley dislocation loop and the Stair-rod dislocation tetrahedron [Figs. 13(a) and 13(b)]. When the collision velocity is 2250 m/s, some Shockley dislocations are further overcome in the form of Eq. (10) and transformed into Hirth dislocations. At a collision velocity of 2750 m/s, the combined structure formed by a large number of intersecting stacking faults and stationary dislocations such as Stair rod and Hirth makes dislocation slip difficult to occur. Severe twinning deformation occurred in the Ti atomic region, forming twin interfaces at the intersection of stacking faults,
1 6 [ 2 ¯ 1 ¯ 1 ] + 1 6 [ 211 ] 1 3 [ 110 ] .
(10)
FIG. 13.

Dislocations and atomic images at the local position of the end of cooling, (a) the local dislocation at the collision speed of 1750 m/s, (b) the local atom distribution at the collision velocity of 1750 m/s, (c) the local dislocation at the collision speed of 2250 m/s, and (d) the local dislocation at the collision speed of 2250 m/s.

FIG. 13.

Dislocations and atomic images at the local position of the end of cooling, (a) the local dislocation at the collision speed of 1750 m/s, (b) the local atom distribution at the collision velocity of 1750 m/s, (c) the local dislocation at the collision speed of 2250 m/s, and (d) the local dislocation at the collision speed of 2250 m/s.

Close modal

Figure 14 shows the grain segmentation results of the system cooling process at the collision velocity of 2750 m/s. With the decrease in temperature, the grain structure of Ti atom region gradually precipitates. When the temperature of the system drops to 600 °C, the Ti atomic region ends the lattice disorder state, and the coarse equiaxed grains are precipitated by cooling. As the temperature continues to drop, the grains in the Ti atomic region are further refined, indicating that there are still a large number of crystal nuclei in the Ti atomic region at this stage. Compared with the rate of grain growth, the nucleation rate of titanium crystal nuclei has a greater effect on the grain growth. With the decrease in temperature, the grain structure of the Ti atomic region away from the interface is basically stable, and the Ti atomic region near the interface continues to precipitate fine grains and finally shows that the grain size at the interface is smaller than that at the boundary when it drops to room temperature. This is because the plastic deformation near the interface is more intense, and there are more defects in the crystal structure of titanium, which generates more crystal nuclei. In addition, a small amount of Al atoms diffused into the Ti atomic region at the interface can also lead to the formation of partial heterogeneous crystal nuclei on the titanium side. Manna et al.52 pointed out that when metals undergo severe plastic deformation, a large number of twin structures will be formed to relieve stress if there is not enough slip system. When the slip and twin deformation are still unable to withstand the stress, titanium will be transformed from HCP to FCC crystal, which is also the reason why there are many FCC structures in addition to HCP structures in the Ti atomic region at the end of the cooling stage. The formation of fine grains results in the material undergoing fine crystal strengthening, which aligns with the strengthening effect observed in FCC-Ti. Fine crystal strengthening has the potential to enhance the strength of titanium while maintaining the material's plasticity, thereby indirectly improving the strength of the Ti/Al interface bonding region. Upon reducing the system to room temperature, it becomes evident that the grain size decreases as the titanium side approaches the interface.

FIG. 14.

Results of grain segmentation during the cooling stage.

FIG. 14.

Results of grain segmentation during the cooling stage.

Close modal

The molecular dynamics results indicate that element diffusion primarily occurs during the unloading stage. To further analyze the bonding mechanism of Ti/Al explosive welding, a phase analysis of the unloading stage, during which element diffusion occurs, was performed using the radial distribution function. The bonding mechanism of Ti/Al composites was analyzed based on the phase state of the materials. The radial distribution function of the material is presented in Fig. 15. At a collision velocity of 1750 m/s, both titanium and aluminum materials exhibit long-range ordered solid metal structures. At a collision velocity of 2250 m/s, titanium maintains its solid crystal structure, while aluminum changes to a long-range disordered liquid state. At a collision velocity of 2750 m/s, the radial distribution function of Ti/Al atoms indicates a long-range disordered liquid state.

FIG. 15.

Calculation results of radial distribution function under three collision velocities.

FIG. 15.

Calculation results of radial distribution function under three collision velocities.

Close modal

At the unloading stage, Ti/Al showed three phase combinations of solid + solid, solid + liquid, and liquid + liquid, respectively. After the end of the cooling stage, the Ti atomic region showed three crystal structures, namely, HCP crystal structure, FCC crystal structure, and “HCP + FCC crystal structure.” The material characteristics of the three collision velocities are obviously different. After the system is cooled, the tensile test directly along the thickness direction is the most effective way to verify the bonding force. At the end of the cooling stage, the tensile stress–strain curves of the system under three collision velocities are shown in Fig. 16.

FIG. 16.

Three velocity molecular dynamics tensile test curves.

FIG. 16.

Three velocity molecular dynamics tensile test curves.

Close modal

The stress–strain curves of the system under three different collision velocities follow the typical trend of stress–strain curves in metal tensile tests. The stress–strain results show that there is no clear distinction between the yield stage and strengthening stage in Ti/Al composite materials, as the material directly reaches its maximum stress value from the yield stage. At collision velocities of 1750 and 2250 m/s, the tensile strength of the Ti/Al system is essentially the same. However, at a collision velocity of 2750 m/s, the tensile strength of the system is approximately 30% lower than in both cases mentioned earlier. Further comparison between a collision velocity of 2250 m/s and a collision velocity of 1750 m/s reveals that the former has a slightly longer elastic stage than the latter. After reaching their respective maximum stress values, they show stronger and more stable tensile capabilities compared to the latter.

The volumetric strain cloud maps for the three systems are presented in Fig. 17 at the conclusion of the calculation. Both samples at collision velocities of 1750 and 2750 m/s display evident necking phenomena, which is a typical behavior of metal materials in tensile tests. At a collision velocity of 1750 m/s, the strain mainly occurs on the aluminum side, while at a collision velocity of 2750 m/s, the strain primarily takes place in the diffusion layer composed of Ti/Al atomic mixture. At a collision velocity of 2250 m/s, the entire system exhibits a volume strain with an inclined angle of 45°, with more intense strains observed on the titanium side. The 45° strain is primarily caused by shear stress. As mentioned earlier, materials develop slip systems to transmit stress during plastic deformation. Under shear stress, crystal planes undergo relative displacement along slip systems resulting in shear strains. Shear strains facilitate the production of FCC-Ti. The combination of stress–strain curves and volumetric strain cloud maps reveals that a collision velocity of 2250 m/s is more likely to generate shear strains. The formation of FCC-Ti at this collision velocity enhances material ductility and tensile strength. When tested at a stretching rate of 0.3 m/s, this combined Ti/Al composite material demonstrates higher bonding strength under such collision conditions.

FIG. 17.

Volumetric strain cloud map of the system at different collision velocities, (a) 1750, (b) 2250, and (c) 2750 m/s.

FIG. 17.

Volumetric strain cloud map of the system at different collision velocities, (a) 1750, (b) 2250, and (c) 2750 m/s.

Close modal

1. Interface microstructure

The microstructure analysis of the composite materials prepared by the test can accurately read the bonding condition of the materials. The samples were prepared near and far from the initiation point respectively (see Fig. 18), and the flyer plate–interlayer interface and the interlayer–baseplate interface were tested. As shown in Fig. 18, the boundary between the flyer plate Ti-6Al-6 V and the interlayer is very obvious, while the boundary between the 1060 interlayer and the 7075 baseplate is not easy to distinguish.

FIG. 18.

Samples at different positions, (a) location of initiation point and (b) far from the point of initiation.

FIG. 18.

Samples at different positions, (a) location of initiation point and (b) far from the point of initiation.

Close modal

The interface combination near the initiation point is shown in Fig. 19. It can be seen from Figs. 19(a) and 19(b) that the TC4/1060 interface and 1060/7075 interface near the initiation point are typical straight interfaces. At a further magnification, Fig. 19(c) shows that the TC4/1060 interface has ingot microstructure caused by excessive melting, while the 1060/7075 interface shows a high-quality bonding morphology.

FIG. 19.

SEM results of the location of the initiation point, (a) TC4/1060 interface, (b) 1060/7075 interface, and (c) and (d) are the (a) and (b) local magnification results, respectively.

FIG. 19.

SEM results of the location of the initiation point, (a) TC4/1060 interface, (b) 1060/7075 interface, and (c) and (d) are the (a) and (b) local magnification results, respectively.

Close modal

The SEM results of the interface far from the detonation point are shown in Fig. 20. Different from near the initiation point and far away from the initiation point, the TC4/1060 interface presents a waveform shape with a wavelength of 600 μm and a wave height of 80 μm. Along with the waveform morphology, a small number of splash melt block defects can be observed at the interface. As can be seen from Figs. 20(b) and 20(d), far from the initiation point, the 1060/7075 interface still has a flat bonding morphology with excellent bonding quality. It is worth mentioning that in the study of explosive welding of TA2/1060/5083 in literature,18 it was also found that the flyer plate–interlayer morphology was waveform, and the interface between interlayer and baseplate was flat morphology. There are obvious differences between the TC4/1060 interface shown in Figs. 19 and 20. This is because, in the actual explosive welding experiment, the explosive requires a certain distance to improve the detonation speed and make the detonation speed reach a stable state, resulting in the detonation speed of the explosive near the initiation point being lower than the actual detonation speed of the explosive far from the initiation end. Further comparison shows that the bonding quality of 1060/7075 interface is better than that of TC4/1060 interface, whether near or far from the initiation point. In the weldability window construction conducted above, the Ti/Al window is obviously narrower than the aluminum/aluminum window, and the test results verify the calculated results of the window.

FIG. 20.

SEM results far from the initiation point, (a) TC4/1060 interface, (b) 1060/7075 interface, and (c) and (d) are the (a) and (b) local magnification results, respectively.

FIG. 20.

SEM results far from the initiation point, (a) TC4/1060 interface, (b) 1060/7075 interface, and (c) and (d) are the (a) and (b) local magnification results, respectively.

Close modal

2. Interface element analysis

Figure 21 shows the line scanning and surface scanning results of elements at the interface of TC4/1060 away from the detonation location. Figure 21(b) shows that there is obvious element diffusion at the flyer plate–interlayer interface, with a diffusion thickness of 8 μm. In the element surface scanning analysis of the sample, the presence of another element cannot be clearly observed at the Ti/Al interface due to the limitation of the test accuracy. However, an obvious coexistence of Ti/Al elements can be seen in the vortex area identified in the figure, indicating that there may be a certain amount of Ti/Al intermetallic compounds.

FIG. 21.

Sample element scanning results, (a) scan position, (b) line scan results, (c) results of the aluminum side of the face scan, and (d) results of the titanium side of the face scan.

FIG. 21.

Sample element scanning results, (a) scan position, (b) line scan results, (c) results of the aluminum side of the face scan, and (d) results of the titanium side of the face scan.

Close modal

For composite materials, the weakest link in bonding strength is crucial in determining the overall bonding quality. SEM results indicate that the bonding condition of the TC4/1060 interface is significantly inferior to that of the 1060/7075 interface. Consequently, EBSD technology is employed to conduct further testing and analysis of the TC4/1060 interface. The test results, as depicted in Fig. 22, illustrate the elements near and away from the initiation point. In the EBSD test, blue represents the detected titanium region, while red represents the detected aluminum region. The findings reveal that the titanium element has evidently diffused to the aluminum side, and the depth of titanium element diffusion in the sample away from the detonation point is greater, indicating a more pronounced temperature rise and plastic deformation resulting from the collision at this location.

FIG. 22.

EBSD element phase results of samples, (a) location of initiation point and (b) far from the point of initiation.

FIG. 22.

EBSD element phase results of samples, (a) location of initiation point and (b) far from the point of initiation.

Close modal

3. Texture analysis

Figure 23 shows the test results of EBSD inverse pole figure of the sample. Microstructure is a critical factor that reflects the degree of material deformation and determines its performance. Following explosive welding, the titanium side exhibits a typical twin grain structure, with smaller grain sizes observed closer to the interface. This observation aligns closely with the results of the molecular dynamics simulation at a collision velocity of 2750 m/s, validating the accuracy of the molecular dynamics settings for explosive welding. On the other hand, the aluminum side displays elongated grains parallel to the direction of detonation. A comparison of the microscopic morphology near and far from the initiation point reveals that the grains at the interface, farther from the initiation point, are finer, indicating a more intense collision at a greater distance from the initiation point.

FIG. 23.

EBSD inverse pole figure of the sample, (a) the titanium side of the initiation point position, (b) the aluminum side of the point of initiation, (c) the titanium side away from the initiation point position, and (d) the aluminum side away from the point of initiation.

FIG. 23.

EBSD inverse pole figure of the sample, (a) the titanium side of the initiation point position, (b) the aluminum side of the point of initiation, (c) the titanium side away from the initiation point position, and (d) the aluminum side away from the point of initiation.

Close modal

Further statistical analysis of the grain orientation in the material allows for texture analysis. The textures of Ti and Al can be determined by combining the polar and inverse polar maps of Ti and Al, as shown in Figs. 23 and 24. In aluminum, there is a noticeable change in texture at two locations: near the initiation point, aluminum mainly exhibits {111}⟨101⟩ texture, while further away from the initiation point, it transitions to predominantly {101}⟨111⟩ and {111}⟨112⟩ textures with a small amount of {011}⟨211⟩ brass texture. On the other hand, there is no significant change in texture observed in titanium flyer plate at both locations—near and far from the initiation point—where they exhibit {2−1−10}⟨21−1−3⟩ and {4−2−2−3}⟨10−1−2⟩ textures, respectively. The difference between FCC and HCP structures in terms of difficulty for dislocation slip and grain boundary sliding is considered the main reason for these variations (Fig. 25).

FIG. 24.

The result of the polar figure, (a) the titanium side of the initiation point position, (b) the titanium side away from the initiation point position, (c) the aluminum side of the point of initiation, and (d) the aluminum side away from the point of initiation.

FIG. 24.

The result of the polar figure, (a) the titanium side of the initiation point position, (b) the titanium side away from the initiation point position, (c) the aluminum side of the point of initiation, and (d) the aluminum side away from the point of initiation.

Close modal
FIG. 25.

The result of the inverse polar figure, (a) the titanium side of the initiation point position, (b) the titanium side away from the initiation point position, (c) the aluminum side of the point of initiation, and (d) the aluminum side away from the point of initiation.

FIG. 25.

The result of the inverse polar figure, (a) the titanium side of the initiation point position, (b) the titanium side away from the initiation point position, (c) the aluminum side of the point of initiation, and (d) the aluminum side away from the point of initiation.

Close modal

Figure 26 shows the KAM results of the sample. The average orientation difference at different locations was calculated by applying a maximum azimuth difference of 5° to adjacent grains. It can be obviously observed from the figure that the sample has high internal grain stress and grain boundary stress, which indicates that after the huge collision pressure in the explosive welding process, there is still a large amount of internal stress in the stably cooled material. Further comparison shows that the internal grain stress and grain boundary stress on the titanium side are significantly higher than that on the aluminum side, because the slip system of the HCP crystal structure of titanium is much less than that of the FCC structure of aluminum, so the stress conduction ability is lower than that of aluminum, and more stress is limited to the grain and the internal material. It is worth mentioning that the difference in the grain structure of Ti/Al is also due to the above reasons.

FIG. 26.

KAM test results of sample, (a) location of initiation point and (b) far from the point of initiation.

FIG. 26.

KAM test results of sample, (a) location of initiation point and (b) far from the point of initiation.

Close modal

The paper establishes a weldability window for multi-layered metal explosive welding and designs an experimental configuration for ultra-thin flyer plate explosive welding. The molecular dynamics algorithm is employed to calculate the diffusion of elements, lattice changes, and dislocation evolution during Ti/Al explosive welding. Based on theoretical calculations and numerical simulation results, a TC4/7075 explosive welding composite material with a flyer plate thickness of only 0.3 mm is fabricated. Microscopic testing and characterization are conducted on the prepared composite material, leading to the following conclusions:

  • In TC4/7075 explosive welding, incorporating a pure aluminum 1060 interlayer can alleviate the challenges in preparing TC4/7075 by expanding the weldability window. Additionally, the utilization of a thin steel plate as a force transmission layer and a thin paperboard as an insulation layer in the ultra-thin flyer explosive welding configuration enables the high-strength bonding of ultra-thin flyer TC4/7075 composite materials.

  • During the preparation of TC4/7075 composite materials by explosive welding, element diffusion at the interface primarily manifested as titanium spreading toward the aluminum plate. The diffusion of elements predominantly occurred during the unloading stage under high temperature and zero external pressure (low external pressure). Under the high temperature and ultrahigh pressure generated by explosive welding, three rare crystal structures were observed in the system: BCC-Al, BCC-Ti, and FCC-Ti. Molecular dynamics results indicated that at a collision velocity of 2250 m/s, the interface cooled to exhibit an FCC structure for titanium, which demonstrated superior tensile mechanical properties.

  • At the initiation point and away from the initiation point, the interface of TC4/1060 exhibits a flat morphology and waveform morphology, respectively, accompanied by a small number of ingot structure and melt block defects. In contrast, the 1060/7075 interface is an ideal straight interface without defects, and these two interface characteristics align with the calculation results of the weldability window. Due to the HCP crystal having fewer slip systems than FCC crystals, under the action of explosive pressure, fine equiaxed twins are formed on the titanium side, while elongated grains are generated along the explosive direction on the aluminum side.

This work was supported by the National Natural Science Foundation of China (Grant No. 51541112) and the Natural Science Foundation of Jiangsu Province (Grant No. BK20211232). The authors would like to express their gratitude to Anhui Honglei New Materials Group for providing materials and experimental sites for this study. They also appreciate the technical support provided by Nanjing University for testing and analysis.

The authors have no conflicts to disclose.

  W. X. and C. Y. contribute equally to this work.

Wu Xiaoming: Conceptualization (equal); Methodology (equal); Writing – original draft (equal). Shi Changgen: Writing – review & editing (equal).

The data that support the findings of this study are available from the corresponding author upon reasonable request.

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