Effect of Temperature Distribution on Interfacial Bonding Process between CFRTP Composite and Aluminum Alloy during Laser Direct Joining

Abstract

Laser direct joining enables non-destructive and lightweight joining of carbon fiber reinforced thermoplastic (CFRTP) composites and aluminum alloys. The interfacial bonding process determines the joint performance and is influenced by the time-varying temperature distribution. However, the interfacial bonding process occurs inside the joint, making it difficult to study the effect of temperature distribution. To resolve this issue, a novel online observation device for the interfacial bonding process between CFRTP composites and aluminum alloys is design, and the polymer melting, flowing, and bonding with metal during laser direct joining are observed. Further, temperature field simulation models for laser direct joining are established, and temperature distribution and gradient are calculated. The results show that the temperature distribution determines the melting of CFRTP composites, and bubbles generated by the thermal decomposition of the polymer hinder the melting. The temperature gradient is related to the movement of the molten matrix and fibers, and the movement towards the aluminum alloy induces cracking and delamination. Once the interface is filled with polymer, the motion changes to along the laser scanning direction and the joining defects are reduced. The study can provide a foundation for promoting interfacial bonding and reducing the defects of laser direct joining.

1. Introduction

Lightweight design and manufacturing are the development direction of aerospace, transportation, and other fields, which can realize performance improvement, energy saving, and emission reduction [1,2]. Carbon fiber reinforced polymer composites have the advantages of high specific strength, high specific stiffness, designable properties, and integral fabrication of components, making them the preferred materials for reducing weight and improving performance [3,4]. Depending on the polymer matrix, such composites can be divided into thermosetting and thermoplastic. Both types of composites need to be used in applications with metals such as aluminum and titanium alloys, thus weakening the negative effects of strong anisotropy, low interlaminar shear properties, and the high manufacturing costs of composites [5,6]. For instance, the fuselage of the Airbus A350XWB airliner is mainly made of carbon fiber reinforced polymer composites, while aluminum and titanium alloys are used for the reinforcement of certain high-load-bearing structures [7,8].

Compared with thermoset composites, carbon fiber reinforced thermoplastic (CFRTP) composites exhibit stronger impact resistance, as well as the advantages of heat molding, fusion joining, and recycling [9,10,11]. With the increasing maturity of raw material preparation and component molding, they are gradually used in aircraft flat tails, fuselages, etc., and face the same problem of joining with metals [12,13,14]. Mechanical joining methods such as bolting and riveting are commonly used for structures subjected to concentrated loads. For structures subjected to uniform or shear loads, fusion joining, which relies on interfacial bonding formed by molten thermoplastic polymers, is gaining importance. This approach avoids the performance degradation of CFRTP composites caused by material processing and reduces the use of adhesives and fasteners, thereby maximizing the effect of weight reduction [15,16]. Depending on the heat source, fusion joining of CFRTP composites with metals can be categorized as friction joining, ultrasonic joining, laser joining, induction joining, and resistance joining. Among them, laser direct joining is the approach where a high energy density laser beam is used to heat the metal, putting the polymer matrix in contact with metal to melt and bond with metal to form a joint, with the advantages of good adaptability to material types and geometrical structures and easy automation [17,18,19,20].

The joining temperature is the key to determining whether CFRTP composites and metals can be joined through fusion [21,22]. Scholars have investigated the effect of joining temperature on the laser direct joining process using simulations and experiments. Hussein et al. equated the laser energy input to a Gaussian surface heat source and simulated the temperature field when laser direct joining PMMA and stainless steel [23]. To improve the accuracy of the temperature simulation, the temperature-dependent thermal properties and the anisotropy of the CFRTP composite were further considered, which allowed the average error between the simulated and measured temperatures to be less than 10% [24,25]. Based on the temperature field simulation during laser direct joining, scholars investigated the relationship between temperature distribution and the melting and solidification process of CFRTP composites. Jiao et al. used the melting point of the polymer matrix as the melting boundary criterion in the temperature simulation, that is, 280 °C for the polyphenylene sulfide used, and found that the melting widths and depths obtained from the simulation agreed with the experiment [26]. Similarly, Zhang et al. used a melting temperature of 303 °C for a nylon 6T matrix as the melting boundary criterion, and the CFRTP composite solidification sequence during the laser direct joining was investigated [27]. Further, considering the effect of melt flow on temperature contributes to the accurate determination of the melting boundary, Ai et al. developed a simulation model of laser joining considering the melting and flow processes and found that thermal convection in the molten polymer increased the melting width, and increasing the laser power increased the rate of thermal convection [28].

The joining temperature is not only related to the melting and solidification at the interface, but also determines the defect formation during the laser direct joining. Tao et al. [29] measured the interfacial temperature between a CFRTP composite and titanium alloy using a K-type thermocouple, and Lambiase et al. [30] measured the surface temperature of an aluminum alloy during laser direct joining using an infrared thermographic camera. They both found insufficient melting of the CFRTP composite at low temperatures and bubbles from thermal decomposition at high temperatures. Based on these studies, it was proven that laser direct joining can be improved by the interfacial temperature control, which ranges between the melting and thermal decomposition temperatures of the polymer matrix [31]. Meanwhile, it is also important to increase the interfacial bonding area of the CFRTP composite to the metal. However, in optimized laser direct joining processes based on temperature control, the bonding area is related to the melting width and can be increased by increasing the temperature, but is accompanied by the risk of thermal decomposition of the polymer [32]. Considering that thermal convection in the molten polymer contributes to the increase in the melting width, to simultaneously promote interfacial bonding and reduce joining defects, it is also necessary to investigate the effect of temperature distribution on the polymer matrix melting, flow, and bonding with the metal.

To reveal the effect of temperature distribution on interfacial bonding between CFRTP composites and metals, this paper designs an observation device for the interfacial bonding process during laser direct joining, and the carbon fiber reinforced polyether-ether-ether-ketone thermoplastic composite (CF/PEEK) and 6061 aluminum alloy used in aviation are taken as the experimental materials, realizing the observation of the polymer melting, flowing, and bonding with metal during laser direct joining. For the interface with and without polymer layer filling, simulation models of the joining temperature field of considering the anisotropic heat transfer and temperature-varying material properties are established, respectively. The relationship between the joining temperature and the melting and flow of the thermoplastic polymer is analyzed, and the influence of the filled polymer on the interfacial bonding process is revealed, which can provide a foundation for promoting interfacial bonding and reducing the defects of laser direct joining.

2. Materials and Methods

2.1. Experimental Device and Setup for Interfacial Bonding Process Observation

To observe the interfacial bonding process during laser direct joining of CFRTP composite with aluminum alloy, an online observation experimental device as shown in Figure 1 was designed. Principle of observation of interfacial bonding process: Firstly, the polished end faces of the CFRTP composite and the aluminum alloy were pressed onto a refractory glass, thus equating them to the confined material interface during laser direct joining; subsequently, a laser was used to irradiate the aluminum alloy close to the refractory glass, and a camera with a microscope was used to observe the interfacial bonding process through the refractory glass. The observation device consisted of the fixture, refractory glass, microscope, camera, focus adjuster, height adjuster, and laser baffle. The camera with microscope had a spatial resolution of 1.23 μm/pixel and a frame rate of about 40 Hz, satisfying the observation requirements of the interfacial bonding process. To accurately observe the interfacial bonding process, thermal insulation was made at the contact area between the fixture and the materials.

The experimental materials were carbon fiber reinforced polyether-ether-ether-ketone (CF/PEEK) thermoplastic composite, 6061 aluminum alloy, and polyether-ether-ether-ketone (PEEK) polymer, which have been used in aviation and other fields. The lengths and widths were 100 mm × 25 mm, and the thicknesses were 2 mm, 3 mm, and 0.25 mm, respectively. The mechanical and thermal properties of the experimental materials are shown in

Table 1. Mechanical and thermal properties of the experimental materials [25].

Parameters	CF/PEEK	6061 Aluminum Alloy	PEEK Polymer
Density (kg/m3)	1580	2700	1300
Tensile strength (MPa)	980	320	100
Melting temperature (°C)	343	615–655	343
Decomposition temperature (°C)	520	-	520
Specific heat capacity J/(kg·K)
Thermal conductivity W/(m·K)

. The CFRTP composite consisted of 16 plies alternating at 0° and 90°, and its fiber volume fraction was 60%. As shown in Figure 2, the laser direct joining experimental device consisted of a laser head, motion platform, observation device, and acquisition computer. Since the quality of the end faces of the materials directly determines the effect of observation, 120#, 400#, 1000#, and 3000# sandpapers were used to polish the end faces of the CFRTP composite and aluminum alloy until the end faces were smooth and without obvious scratches. Once the materials were prefixed to the fixture, it was necessary to apply pressure on the refractory glass. After ensuring that the polished end faces of the materials were in close contact with the refractory glass, the materials were clamped under the same torque using a torque wrench, and the observing position was adjusted to the center of the interface using the height and focus adjusters. Finally, to prevent the laser scattering from affecting the observation effect, the laser baffle was attached to the fixture.

Because of the high fiber volume fraction of the CFRTP composite, insufficient polymer matrix may lead to poor interfacial bonding during laser direct joining, and filling the polymer layer at the interface between the CFRTP composite and the aluminum alloy is a common solution [33]. To this end, laser direct joining experiments with and without polymer layer filling were performed as shown in Figure 3. In the experiments, a fiber laser with a wavelength of 1064 nm was used to scan the aluminum alloy surface reciprocally six times. The laser beam was in defocused state; the defocusing distance was 80 mm, the corresponding spot diameter was 5 mm, the maximum laser power was 1000 W, and the maximum laser energy density was about 5 × 10³ W/cm². For both cases, the scanning speed of the laser was 10 mm/s, the single scanning length was 20 mm, and the center of the laser spot was 2.5 mm from the edge of the aluminum alloy. When there was no filled polymer layer, the laser powers were 400 W, 500 W, and 600 W, whereas the laser powers were 500 W, 600 W, and 700 W when the interface was filled with polymer layer, because the polymer has a high specific heat capacity and low thermal conductivity.

2.2. Simulation Modeling of Temperature Field during Laser Direct Joining

To simulate the temperature field during laser direct joining, the geometrical models shown in Figure 4 were established for the interface with and without polymer layer filling, respectively, and the material dimensions were the same as those in Section 2.1. To balance the computational accuracy and efficiency, two sizes of grids were chosen for meshing. Since the temperature gradient in the heated region Γ₁ irradiated by the laser was high, a fine hexahedron with a size of 0.625 mm was used for meshing parallel to the thickness direction. The other region, Γ₂, which had a low temperature gradient, was meshed with a coarse hexahedral of size 2.5 mm, and wedges were used for the transition between the two regions. In the material thickness direction, aluminum alloy was meshed using 0.5 mm, and the polymer layer was meshed using 0.25 mm. Considering the anisotropy and ply structure of the CFRTP composite, it was meshed using ply thickness of 0.125 mm.

After establishing the meshed geometric models, it is necessary to set the governing equations and boundary conditions of the laser direct joining, including the material properties, the laser heat source, and the cooling conditions. Specific details can be found in our previous study [25] and are briefly described here. The temperature field of the laser direct joining obeys the heat transfer equation in (1):

$$\rho C_p\frac{\partial T}{\partial t}=k_x\frac{{\partial }^{2}T}{\partial x^2}+k_y\frac{{\partial }^{2}T}{\partial y^2}+k_z\frac{{\partial }^{2}T}{\partial z^2}$$

(1)

where ρ is the density; C_p is the specific heat capacity; T is the temperature; t is the time; and k_x, k_y, and k_z are the thermal conductivity in the x, y, and z directions, respectively. To solve the above heat transfer equation, the initial temperature T₀ was set to be 20 °C, and the material properties were set based on Table 1. Additionally, based on the laser heat input and convective and radiative heat transfer on the outer surface of the material, the boundary conditions for the heated region Γ₁ and other regions Γ₂ were set as follows:

$$k\frac{\partial T}{\partial n}=h(T_s-T_0)+\epsilon \sigma (T_s^4-T_0^4)+q\mathrm{on}{\mathsf{\Gamma }}_1$$

(2)

$$k\frac{\partial T}{\partial n}=h(T_s-T_0)+\epsilon \sigma (T_s^4-T_0^4)\mathrm{on}{\mathsf{\Gamma }}_2$$

(3)

where n is the normal vector of the outer surface; q is the laser heat flow density; h is the convective heat transfer coefficient, which was set to 5 W/(m²·K); T_s is the surface temperature; ε is the emissivity; and σ is the Stefan Boltzmann constant, which was 5.67 × 10⁻⁸ W/(m²·K⁴)). The emissivity coefficients of CFRTP composite and aluminum alloy were roughly set because of the little effect of radiative heat transfer on the temperature during the laser direct joining [34,35]. Since the laser energy density was about 10³~10⁴ W/cm², and the aluminum alloy did not melt or had a small melting thickness, the Gaussian surface heat source was used to equalize the laser energy input. The expression is as follows:

$$q=\frac{3\eta P}{\pi R^2}\exp (-\frac{3[{(x-x_0-vt)}^2+{(y-y_0)}^2]}{R^2})$$

(4)

where η is the laser absorption rate of the aluminum alloy; P is the laser power; R is the effective radius of the laser, that is, the radius corresponding to when the laser intensity is attenuated to 5% of the maximum value; x₀ and y₀ are the initial irradiation position of the laser; and v is the laser scanning speed. Among them, the laser absorption rate of aluminum alloy η was set to 0.3 based on the inverse analysis method reported by Lambiase et al. [34], the effective radius R of the laser was set to 2.5 mm, and the corresponding defocusing distance of the laser was 80 mm.

For the heat transfer at the material interface, since the polymer matrix of the CFRTP composite experiences the heating, melting, and flowing processes, the contact thermal resistance only influences the temperature field before the melting. On the other hand, contact thermal resistance is affected by external pressure, which is required for laser direct joining of CFRTP composite with aluminum alloy; Wang et al. have shown that when the external pressure reached a certain level it could be regarded as an ideal contact heat transfer [36]. Therefore, the effect of contact thermal resistance was ignored during the temperature simulation of laser direct joining. The temperature field during laser direct joining was solved using ABAQUS, and transient thermal analysis was performed on a simulation platform with an i5-11500 CPU and 8 GB of RAM.

3. Results and Discussion

3.1. Consistency Verification of Observation and Simulation Results

To verify the relationship between the temperature distribution and the melting of the CFRTP composite, the observed interfacial bonding processes and the corresponding temperature distributions at the material end faces in the simulation were compared as shown in Figure 5 and Figure 6. In the observed results, the melting of the polymer matrix caused the thickness change in the CFRTP composite, and the melting depth of the CFRTP composite was difficult to accurately determine; therefore, the number of melted plies of CFRTP composite was used to compare the simulated and observed results. In the simulation results, the melting temperature (343 °C) of the PEEK polymer was used as the melting boundary criterion, and the corresponding melting area was indicated in gray. When the interface was not filled with polymer, Figure 5a shows the results at a laser power of 400 W. The position change of the fibers in the CFRTP composite and the cracking between the matrix and fibers were observed in the first ply, whereas the other plies were unchanged, indicating that melting only occurred in the first ply. Continuing to increase the laser power to 500 W and 600 W, the melting of the three and six plies, as shown in Figure 5b,c, respectively, was found in the observed results using a similar criterion. The observed number of melted plies was the same as the number of melted plies in the gray area in the simulation, indicating that there was a correlation between the number of melted plies and the temperature distribution. With the increase in laser power, the time corresponding to the maximum number of melted plies was delayed because the temperature of the aluminum alloy was higher at higher laser powers, and the CFRTP composite was still heated by a high-temperature aluminum alloy after the laser stopped irradiating the observation area. However, when the number of melted plies increased, the fibers in the direction of the parallel end faces were extruded, which affected the observation of the interfacial bonding process.

The observed results and corresponding temperature distributions after filling the polymer layer at the interface are shown in Figure 6. The molten polymer layer filled the gap between the end face of the material and the refractory glass, leading to a more pronounced color change compared with the melting. However, this did not mean that melting of the polymer matrix occurred, and it was still necessary to use the position change of the fibers or cracking between the matrix and fibers as the melting criterion. Due to the lower thermal conductivity of the polymer than the CFRTP composite, the number of melted plies was lower under the same laser power. When the laser power was 500 W, cracking was not observed in Figure 6a, but the observations at 11.58 s and 11.89 s showed that the position of the fibers in the first ply changed, indicating that the polymer layer was completely melted and the first ply was partially melted, which was consistent with the simulation. With similar criterion, it was found that the number of melted plies in the CFRTP composite increased to three plies when the laser power was 600 W, and it was accompanied by thermal decomposition in the molten polymer layer. When the laser power was 700 W, the number of melted plies in the CFRTP composite was still observed to be three plies, which was different from the simulation result of five plies, because bubbles in molten polymer induced by the thermal decomposition hindered the heat transfer. The above results proved that the temperature distribution during laser direct joining was correlated with the melting of the polymer matrix and the interfacial filled polymer.

To investigate the effect of bubbles formed by the thermal decomposition of molten polymer on the bonding process, the results of interfacial observations and temperature simulations were extracted for the laser powers of 600 W, shown in Figure 7, and 700 W shown in Figure 8, where the gray area in the temperature distribution was the area beyond the thermal decomposition temperature of the polymer (520 °C). When the laser power was 600 W, the surface temperature of the polymer layer exceeded the thermal decomposition temperature, and bubbles were observed on the corresponding surface of the polymer during the last laser scan. The temperature simulation results showed that the local temperature of the polymer surface exceeded the thermal decomposition temperature at 10.75 s, and the peak temperature of the polymer surface decreased as the laser scanned from the edge to the center. However, from the observations of the corresponding moments, it was found that the bubbles moved toward the laser scanning direction, resulting in the presence of bubbles in all the moments, which may be related to the thermal convection caused by the uneven temperature. Meanwhile, similar fiber position changes were found in the CFRTP composite, but the fiber position change was smaller than the bubbles in the molten polymer because of the constraints from the fiber axial direction. When the laser power was 700 W, as shown in Figure 8, the surface temperature of the thermoplastic polymer exceeded the thermal decomposition temperature at 9.00 s, and bubbles appeared not only on the surface of the molten polymer, but also in the contact area with the CFRTP composite. Since the thermal conductivity of the gas in bubbles was significantly lower than that of the molten polymer, the bubbles reduced the heat transfer from the polymer to the CFRTP composite, which on the one hand made it easier to generate high temperatures in the molten polymer and exacerbated thermal degradation, and on the other hand reduced the number of melted plies of the CFRTP composite, which resulted in a lower number of melted plies than that of the simulation results.

3.2. Effect of Temperature Distribution on Interfacial Bonding Process

To clarify the effect of temperature distribution on the interfacial bonding process, the end face temperatures at different moments when the laser power was 600 W were extracted, and the corresponding temperature gradients were calculated using MATLAB, with the results shown in

Table 2. Temperature distributions and gradients at the end face when the laser power was 600 W.

	Interface without Polymer Filling	Interface with Polymer Filling
Temperature distribution
Temperature gradient
x direction temperature gradient
y direction temperature gradient

. For the interface with and without polymer filling, the temperature distributions in the aluminum alloy were similar, that is, the temperature at the center of the laser irradiation was the highest, a trailing appeared along the laser scanning direction, and the temperatures were uniformly distributed in the thickness direction. In contrast, the temperatures in the CFRTP composite and filled polymer showed a significant decrease along the thickness direction, and the filled polymer resulted in more temperature decreases. The maximum temperature gradient in the aluminum alloy was about 45 °C/mm, whereas that of the polymer layer was about 450 °C/mm and larger than that of the unfilled polymer of about 420 °C/mm. For the x and y direction temperature gradients, the y direction temperature gradient was consistent with the temperature gradient, while the x direction temperature gradient was related to the laser scanning direction. The x direction temperature gradient in the aluminum alloy was negative in the laser scanned area and positive in the unscanned area, when the laser scanned in the x−direction, with maximum and minimum temperature gradients occurring on the sides of the laser irradiation position. The x direction temperature gradients in the CFRTP composite and filled polymers were similar, with reduced positive and negative peak gradients and a trailing near the contact area with the aluminum alloy, which was mainly related to the heating hysteresis caused by the low thermal conductivity.

When the interface was not filled with polymer, Figure 9 shows the interfacial bonding process between the CFRTP composite and the aluminum alloy at a laser power of 500 W. The laser scanning direction is from right to left, and the laser irradiated the surface of the aluminum alloy in the center of the observation area at 7.00 s. At this moment, a gap of about 60 μm existed at the interface between the CFRTP composite and the aluminum alloy on the left side, while the gap on the right side was filled by the CFRTP composite. As the laser scanned to the left, the fibers in the CFRTP composite moved towards the aluminum alloy driven by the molten polymer matrix and filled the gap in the observed area by 7.19 s. Meanwhile, this also produced cracking between the matrix and fibers and formed the delamination between the first and second plies of the CFRTP composite. Similarly, the melting of the polymer matrix in the second ply was observed at 7.24 s, and the movement of the second ply towards the aluminum alloy eliminated the cracking in the first ply, but this likewise produced such defects in the second ply. In addition to the movement of the plies in the CFRTP composite towards the aluminum alloy, movement of the molten polymer matrix together with the fibers towards the laser scanning direction was also found, based on the position change of the cracking in the observations of the material end face from 7.14 s to 7.43 s.

To determine the effects on the interfacial bonding process, the x and y direction temperature gradients at the corresponding moments shown in Figure 10 were drawn. The color meanings in Figure 10 are the same as those in Table 2, red arrows are used to indicate the temperature gradient directions, and red contours in the CFRTP composite indicate the boundaries of the melting area. In all moments, the y direction temperature gradients in the CFRTP composite were directed toward the aluminum alloy, while the direction of the x direction temperature gradients was related to the laser irradiation position. From 7.00 s to 7.14 s, the melting areas surrounded by the contours were basically the same as the areas of the CFRTP composite moving toward the aluminum alloy, indicating that the molten polymer drove the fibers to move toward the aluminum alloy under the large y direction temperature gradient. Meanwhile, since the second ply of the CFRTP composite was not melted, the first ply moved toward the aluminum alloy, causing delamination between it and the second ply. Once the melting of the polymer in the second ply occurred at 7.24 s, the molten polymer and fibers in the second ply moved toward the aluminum alloy, further proving that the temperature gradient in the melting area drove the CFRTP composite to move and the direction was related to the gradient direction. For the movement toward the laser scanning direction, although the temperature gradients from 7.00 s to 7.14 s were in the x+ direction, the observations showed that the polymer matrix drove the fibers to move in the x− direction, whereas temperature gradients toward the x− direction appeared in the observation area at 7.24 s. This was because the flow of the molten polymer in the CFRTP composite was influenced by the fibers; under the extrusion effect of the scanned area and the support effect of the unscanned area, the movement of the molten polymer and fibers in the x direction was affected by the movement which occurred away from the observation area.

The interfacial bonding processes when the interface was filled with the polymer layer are shown in Figure 11, with a laser power of 600 W. The polymer began to melt at 4.85 s in the observation area, corresponding to the laser scanning direction from left to right. Due to the better melt fluidity of the pure polymer, the gas in the gap between the polymer and the aluminum alloy could not be vented in time, and bubbles were found in the molten polymer near the aluminum alloy in the observation results. As the laser scanned to the right, the molten polymer gradually filled the gap between the polymer and the aluminum alloy, accompanied by the movement of bubbles along the laser scanning direction, while there was no significant change in the position in the thickness direction. Compared with laser direct joining without filling the polymer layer, the polymer layer could rapidly infiltrate the aluminum alloy surface after melting. Once the molten polymer filled the gap at the material interface, the fixed aluminum alloy limited the movement towards the aluminum alloy. This resulted in the molten polymer moving mainly along the laser scanning direction, avoiding cracking and delamination in the CFRTP composite caused by the movement towards the aluminum alloy.

To reveal the motion of the molten polymer layer during the interfacial bonding process, the observations and temperature gradients in a single laser scan were analyzed, that is, the laser direct joining processes from 6.10 s to 7.90 are shown in Figure 12. The laser scanning direction was from right to left, and the laser irradiation position was on the sides of the aluminum alloy at 6.10 s and 7.90 s, while the laser irradiation position was on the sides of the observation area at 6.90 s and 7.09 s. From 6.10 s to 6.90 s, when the laser scanned the right side of the outer observation area, the bubble in the molten polymer appeared to move to the x+ direction, whereas the bubble position was unchanged when the laser scanned the observation area from 6.90 s to 7.09 s. Furthermore, the bubble moved to the x− direction when the laser scanned the left side of the outer observation area from 7.09 s to 7.90 s. The x direction temperature gradients in the observation area followed the same trend as those of the bubble movement, that is, the x direction temperature gradient was positive when the bubble moved to the x+ direction, gradually decreased to zero when the bubble position was unchanged, and there was a negative x direction temperature gradient when the bubble moved to the x−direction at 7.09 s. The above results proved that the direction of the temperature gradient was the key to determine the flow of molten polymer, and the molten polymer flowed faster when the direction of the temperature gradient was the same as the laser scanning direction.

4. Conclusions

To investigate the interfacial bonding process between the CFRTP composite and aluminum alloy during laser direct joining, an online observation device for the interfacial bonding process was designed, and the temperature field simulation model for laser direct joining was established. Based on the observed and simulated results, the relationship between the temperature distribution and the melting and flow of the polymer was analyzed, as well as the effect of the filled polymer at the interface on the interfacial bonding process. The main conclusions are as follows:

(1) The melting and flow of the polymer matrix and filled polymer at the interface were observed, as well as the position change of the bubbles and fibers therein. As the laser power increased, the number of melted plies of the CFRTP composite increased when the interface was not filled with polymer. Once the interface was filled with polymer, the number of melted plies increased and then remained constant as the laser power increased, because bubbles caused by the thermal decomposition of the polymer layer at high laser powers hindered the heat transfer into the CFRTP composite.

(2) The number of melted plies in the CFRTP composite obtained with the simulation were the same as the observed results, when thermal decomposition of the polymer did not occur. Affected by the low thermal conductivity of the CFRTP composite and polymer, the largest temperature gradient occurred at the contact area with the aluminum alloy. Regardless of whether the interface was filled with a polymer layer or not, the temperature gradient along the material thickness direction was larger than that along the laser scanning direction, which was about 400 °C/mm and 40 °C/mm, respectively.

(3) When the interface was not filled with polymer, the molten matrix together with the fibers moved to the aluminum alloy under the temperature gradient along the material thickness direction. This resulted in filling the gap between the CFRTP composite and the aluminum alloy, coupled with cracking between the matrix and fibers and delamination in the CFRTP composite. For the movement along the laser scanning direction, less movement of the molten matrix and fibers was observed because of the lower temperature gradient and the support effect of the unscanned area.

(4) When the interface was filled with polymer, the gap between the polymer layer and the aluminum alloy was filled rapidly once the polymer started to melt, after which the aluminum alloy limited the flow of the molten polymer towards the aluminum alloy, thus avoiding cracking and delamination. However, because of the better fluidity of the molten polymer, the gas in the interfacial gap tended to remain in the molten polymer. The molten polymer and bubbles moved in the laser scanning direction, and the motion was in the positive direction of the temperature gradient.

In the future, optimization and modification of the interface-filled polymer layer and development of the matched laser direct joining process are still needed. Those studies are expected to promote the flow of the molten polymer and bonding with aluminum alloy based on the suppression effect of the polymer layer on joining defects. The promoted interfacial bonding and reduced joining defects will contribute to the improved performance and application of laser direct joining between CFRTP composites and aluminum alloys.