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Shaped laser beam profiles for heat conduction welding of aluminium-copper alloys
Optics and Lasers in Engineering 115 (2019) 179–189
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Shaped laser beam profiles for heat conduction welding of aluminium-copper alloys Michael Rasch a,b,c,∗, Clemens Roider a, Stefanie Kohl a,c, Johannes Strauß a,c, Niklas Maurer a, Konstantin Y. Nagulin d, Michael Schmidt a,b,c a
Institute of Photonic Technologies, Friedrich-Alexander University Erlangen-Nürnberg, Erlangen, Germany Collaborative Research Center CRC 814 – Additive Manufacturing, Friedrich-Alexander University Erlangen-Nürnberg, Erlangen, Germany c Erlangen Graduate School in Advanced Optical Technologies (SAOT), Erlangen, Germany d Kazan National Research Technical University, Kazan, Russia b
a r t i c l e
i n f o
Keywords: Heat conduction mode laser beam welding Surface roughness Beam shaping Diffractive optical elements Aluminium alloys Numerical process simulation
a b s t r a c t Heat conduction welding is often used where weld seams with a high surface quality and the exact retention of the chemical composition are needed. This study investigates the influence of intensity profiles in laser beam welding with laser powers up to 3.2 kW. The resulting process windows for heat conduction mode welding, the melt pool shape, the process stability and the dynamics, and the produced surface roughness are analysed. Therefore, three different intensity profiles are created with diffractive optical elements and the process dynamics are monitored with two high-speed cameras. Before analysing etched cross sections with respect to microstructural properties, the surface roughness is measured on all produced samples with a laser scanning microscope. With beam shaping, we found that a high peak intensity does not necessarily lead to an instable melt pool and that the distribution itself plays the major role. The use of shaped beam profiles leads to a more stable process and an enlarged heat conduction process regime compared to a defocussed multimode spot of the same size. Applying shaped laser beam profiles surface roughness close to laser polished surfaces can be achieved while also simultaneously enlarging the weld cross section area.
1. Introduction High strength aluminium alloys are commonly used in aerospace and automobile applications [1]. Due to certain material properties the connection of two parts not using force or form-fit but welding technologies is challenging. Especially EN AW-2xxx alloys suffer from a high risk of hot cracking due to their huge solidification interval, with EN AW-2024 being one of the alloys with the highest hot crack susceptibility suffering from a huge solidification interval [2]. One possibility to deal with this problem is the use of non-melting technologies like friction stir welding (FSW). FSW is advantageous because the process is carried out below the melting temperature. In 1999 crack free welding seams were successfully shown [3]. However, these welding seams are rather large and suffer from a rough surface and therefore, the corrosion resistance is classified as poor. Another drawback is the limitation to the clamping. Recent research is focussed on improving the weld surface quality by laser polishing. Kalita could show that re-melting using a diode laser with laser powers between 700 W and 800 W and a traverse speed of 1.66 mm/s increases the resistance
∗
of EN AW-2024 FSW-welds to pit growth without producing any cracks [4]. Avoiding hot cracks using melting technologies is even more challenging. While the material is molten recrystallization is taking place. Dendritic grain morphology must be avoided because of its well-known high tendency for hot cracks and crack propagation [5]. In electron beam welding a weld seam for both regimes – heat conduction welding and deep welding – consisting of equiaxial grains due to the low temperature gradient but high solidification rate can easily be produced [6]. In laser beam welding most attention is paid to the keyhole welding process due to the higher achievable aspect ratio. The first common solution was to add a filler material like EN AW-4xxx to produce an alloy with an higher silicon content in the weld seam with a lower tendency for hot cracking [5]. Similarly, adding grain refiners like TiB2 leads to fine equiaxial grain structures without any cracks [7]. Successful welding without any additives was achieved using low traverse speed of 40 mm/s and a high laser power of 2.75 kW focussed on a spot with a diameter of 0.45 mm to 0.60 mm [8]. The weld seam consisted mostly of equiaxial dendritic grains. It seems that the high melt dynamics that occur in the keyhole regime and a relatively low temperature gradient
Corresponding author. E-mail address: [email protected] (M. Rasch).
https://doi.org/10.1016/j.optlaseng.2018.11.025 Received 9 July 2018; Received in revised form 11 November 2018; Accepted 26 November 2018 0143-8166/© 2018 Elsevier Ltd. All rights reserved.
M. Rasch, C. Roider and S. Kohl et al.
Optics and Lasers in Engineering 115 (2019) 179–189
due to the low traverse speed lead to strong constitutional undercooling and the preferred grain morphology is formed. Higher traverse speeds of up to 120 mm/s could be reached using an additional heat source reducing the effective thermal stress on the weld seam due to a lower temperature gradient [9]. In this case the grain structure is fully dendritic. The downsides of keyhole welding are rough surfaces, a change in the chemical composition due to a selectively higher evaporation rate of certain alloying constituents and vapour pores [10]. Heat conduction mode welding (HCMW) is known to be the most stable process regime in laser welding, but only little research has been carried out on this process regime (and in most cases HCMW is not even mentioned explicitly). A compressive review with different aluminium alloys was published by Quintino and Assunção pointing out the clearly above mentioned advantages [11]. Sánchez-Amaya et al. are to the authors’ knowledge the only ones investigating 2xxx alloys in HCMW. They used a diode laser with a maximum power of 2.75 kW and a spot size of 2.2 × 1.7 mm2 to create weld seams on six different high strength aluminium alloys, one of them being EN AW-2024. They found that a minimum power of 2 kW and maximum speed of 16 mm/s is needed to create a stable melt pool and crack free weld seams. The alloying constituents proved to play a major role in weldability and hot crack susceptibility as well [12]. In 1999, Zhao et al. predicted that the control of the melt pool geometry will be an important tool for influencing the evaporation of alloying elements and hot crack susceptibility [13]. The first experimental implementation in keyhole welding of AA 6016 was done using a twin laser system equipped with a CO2 laser as primary source and a pulsed Nd: YAG laser as secondary source. It could be shown that at a rather high welding speed of 60 mm/s and CO2 -laser power of 1.7 kW, the use of the pulsed Nd: YAG laser with a pulse energy of 3 J at a repetition frequency of 150 Hz and a pulse duration of 2 ms could effectively prevent the formation of cracks. It was found that the pulsed nature of the Nd: YAG laser and the significant decrease of the thermal gradient promotes the columnar to equiaxial transition and therefore the formation of equiaxial-globular grains [14]. Hansen et al. showed that with multispot keyhole welding of EN DC01 and EN 1.4301 using diffractive optical elements weld seams with an almost rectangular cross section can be obtained [15]. Sundqvist et al. made some studies focussing on the numerical and analytical optimization of laser spot welding with shaped laser beams [16]. Funck et al. did experimental work showing in laser spot welding with a ring profile that mainly the outer diameter affects the convection in the melt pool [17]. As keyhole welding is suffering from poor surface quality this process cannot be used when the surface of weld seams is of interest. In contrast HCMW yields high quality weld surfaces but at the moment suffers from slow speeds. Therefore, the goal of the presented work is to clearly identify the process windows for welding with different shaped laser beams to achieve higher welding speeds. Furthermore, the influence of the beam shape on melt pool dynamics and shape and on the obtained surface quality is determined. Comparing experimental observations with temperature and evaporation predictions from simulations yields further insight and enhances process understanding when using shaped beam profiles.
Fiber plug Fiber end Concav lens: -75 mm JGS1 Convex lens: 75 mm BK7 Area for DOE convex lens: 100 mm BK7 20 mm Fig. 1. Customised welding optics.
of the DOE in the optical path, a collimated laser beam is preferred. Using the knife edge method and fitting with the Gaussian error-function, the beam diameter with dismounted focussing lens is determined at a distance of 100 mm and 300 mm from the lower edge of the optics for verifying the collimation (see Fig. 1). For further calculation a collimated Gaussian beam with a diameter of 18.6 mm is assumed. The diameter was chosen in order to keep the peak intensity of the beam as low as possible on the 25.4 mm DOE, simultaneously avoiding any significant amount of laser power hitting the DOE mount. Another important fact is the caustic of the unshaped laser beam with assembled focussing lens for determining the Rayleigh range, the focus profile and the position of the focal plane. Therefore, the 3D beam shape is determined with a PRIMES MicroSpotMonitor. A calibration procedure was applied for quantifying the intensity distributions. The beam waist was found to be w = 57 μm with an almost perfect flat top shape as the fibre end is imaged onto the sample with a slight demagnification (see Fig. 2). The final experimental setup is shown in Fig. 3. The welding optics is mounted at a fixed position on a welding desk. Additionally, two highspeed cameras (Phantom v1210, Vision Research) are used in order to monitor the melt pool. While camera 1 is monitoring the melting process in welding direction, camera 2 is positioned perpendicular to camera 1 resulting in sharp views on different regions of the melt pool. For illumination, a ns-pulsed diode laser (Cavilux HF, 𝜆 = 808 nm, Pavg = 500 W) is used. A sandblasted EN AW-2024 sheet metal (56.5 × 18.7 × 3.8 mm3 , SA = 1.2 μm) is placed inside a small process chamber, which is moved by an industrial robot (KR30, Kuka) underneath the stationary laser beam. To ensure a constant movement below the welding optics, an acceleration and a deceleration section before and after the welding of 500 mm and a section with constant velocity of 100 mm below the welding optics are implemented. The process chamber is equipped with an easily exchangeable 1 mm thick cover glass (Nexterion D263T, Schott; transmittance including Fresnel reflection losses is 91.85% [18]) 11 mm above the sheet metal. To prevent pollution of the glass and oxidation of the weld seam, the processing chamber is constantly flushed with argon.
2. Methods 2.1. Experimental setup
2.2. Determination of the surface roughness, metallurgical preparation and analysis of the melt pool length
For the following experiments a disk laser with a maximum output power of 8 kW (TruDisk 8001, Trumpf, 𝜆 = 1030 nm, 150 μm fibre, NA = 0.067) and a customised optics with a 100 mm focussing lens have been used (see Fig. 1). As the process is accompanied by an extensive heat build-up the fiber plug is water cooled, while the lenses as well as the diffractive optical element (DOE) are air cooled. For the design of the DOEs it is important to know the diameter and intensity distribution of the laser beam at the DOE position. To reduce axial alignment errors
The arithmetical mean height of the surface SA of produced specimens is analysed with a confocal laser scanning microscope (LEXT OLS4000, Olympus) according to DIN EN ISO 25,178–2 [19] with a cut-off wavelength 𝜆c of 800 μm in a rectangle of 394 × 564 μm2 at the weld centre. Afterwards the samples are cut into three sections. While two sections are metallographically prepared for further analyses, one section is used to analyse the top view. The samples are ground with up 180
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Fig. 2. a) Beam Caustic b) beam profile in the focal plane c) defocussed beam profile d) crosssection in the focal plane e) cross-section through defocussed beam (w ∼ 450 μm).
Fig. 3. Experimental setup. Green Box: Miniature process chamber with an inert gas flow.
to 4000 grit and subsequently polished using a final grain size of 1 𝜇m. To determine the presence of hot cracks, unetched metallographically prepared cross-sections are investigated by light microscopy. For characterizing the dimensions of the weld seams and the grain morphology, etched and electrolytically polished specimens are analyzed. The length of the melt pool lw is determined by analysis of the high speed videos. lw is defined from the last interaction of the laser beam with the melt pool to complete solidification at the symmetry axis.
The corresponding DOEs were calculated using a weighted Gerchberg-Saxton algorithm [21] for POINTS and LINE and the classical Gerchberg-Saxton algorithm [22] for RING. The fabrication procedure follows the on presented by Alexeev et al. [23] with a pixel pitch of 10 μm and a layer with an optical thickness corresponding to 𝜆/2 at 1030 nm. As with this method only binary DOEs can be produced, the desired BPs were also chosen to be point-symmetric. Therefore, the theoretically highest possible diffraction efficiency is 81% with the remaining 19% being distributed in higher diffraction orders. The first row in Fig. 4 shows simulated predictions of the BPs, while the following rows show experimentally measured intensity distributions, integrated intensity, and cross-section views. Note the “ghost spots” visible for LINE at the left/right edge of the image which are caused by the periodic structure [24]. Also a weak zeroth order is visible in the cross-section of RING due to an imperfect DOE fabrication. Imperfection will also lead to stronger higher diffraction orders and stray light, therefore reducing the diffraction efficiency.
2.3. Chosen beam profiles In this study three different beam profiles (BP), which are depicted in Fig. 4, and a defocussed spot (DEF) as reference (see Fig. 2c and e) with the same constant “spot” diameter of around 900 μm, that is about nine times the minimal unmodified spot diameter, are investigated. Note that DEF is not a Gaussian profile but looks like a tipi shape resulting in nearly linear flanks and a pointy maximum. The intention of the first BP, called LINE, is to keep the intensity as well as the integrated intensity along the welding direction constant (see Fig. 4h and k). Recent publications have shown that line like beam profiles can lead to evaporation induced “bow wave” like melt flows in front of the interaction zone [20]. To reduce these effects the second BP (POINTS) was chosen. Instead of having a straight line, the beam profile consists of five spots with a distance of 200 μm between each other. The peak intensity of POINTS is roughly twice as high as in LINE but the space between the spots might create a deformation free region where the melt can flow unhindered (see Fig. 4l). The third BP is a RING, which has a radius of 400 μm and a width of 110 μm, corresponding to the unmodified spot diameter. RING has a higher integrated intensity along the welding direction (Iint ) on the left and the right (see Fig. 4g). That is meant to take the larger heat dissipation at the edges of the melt pool into consideration. Therefore, the temperature should be more equally distributed over the melt pool.
2.4. Numerical simulations Multiphysical process simulations were carried out to complement the experimental results and analyse quantities that are experimentally not accessible. The thermo-fluiddynamic numerical model is based on the finite volume method (FVM) open source framework OpenFOAM® (Open Field Operation and Manipulation) [25]. Within this programming environment, the physics of laser-material interaction were implemented in the form of partial differential equations (PDEs) to model the laser welding process. Detailed information on the physics of numerical model and its verification can be taken from [26]. Since this particular paper addresses HCMW, multiple reflections could be neglected and were not considered in the calculations to save computational time. Although evaporation is not expected to be a dominant effect in HCMW, it still is an important process characteristic since the evaporation driven 181
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Optics and Lasers in Engineering 115 (2019) 179–189
Fig. 4. a-c) calculated beam profiles; d-f) measured beam profiles; g-i) integrated intensity in traverse direction of the measured profiles; j-l) cross-section through the measured profiles; each row was normalised to the highest value for adequately displaying the relative strengths of the different profiles. White arrows mark the welding direction.
recoil pressure defines the border of the process regime towards laser beam deep-penetration welding. Therefore surface tension, evaporation and the resulting recoil pressure are included in the model to be able to reconstruct possible indentations in the melt pool. The included evaporation model calculates the temperature-dependent evaporation pressure as given in [27]: ( ( )) 𝐻𝐵 𝑀 1 1 𝑝𝑣𝑎𝑝 (𝑇 ) = 0.54 ⋅ 𝑃𝑎 ⋅ exp − . (1) 𝑅 𝑇𝐵 𝑇
between the two super-Gaussian functions, the beam radii and the order of the super-Gaussian functions were calculated and the calculated parameters were used as input parameters in the simulation model. Intensities in higher orders and “ghost spots” were neglected in the simulations. As expected, the fitting of the beam profile of the defocussed spot turned out to be more challenging, since its profile was super-Gaussian in its focus. In contrast to Gaussian beams, super-Gaussian beams do not keep their beam profile constant during defocus. For this beam, only one super-Gaussian could be fitted with an effective exponent of n = 1.56, resulting in a rather pointy “sub” -Gaussian function.
In Eq. 1 HB and TB denote the enthalpy and temperature of evaporation, respectively, M denotes the material’s molar mass, R the gas constant, and Pa the ambient pressure. The factor of about 0.54 originates from the consideration of recondensation within the Knudsen layer. It holds for the pressure and evaporation conditions of laser material processing and is widely used for laser welding simulations [28]. All optical, fluid dynamical and thermo-physical material properties of the alloy used are included in the model considering their temperature-dependence. The thermo-physical properties are extracted from the CalPhaD-based (Calculation of Phase Diagrams) material database JMatPro [29] based on the chemical composition of the alloy [30]. The optical properties and their temperature dependence are calculated dependent based on the extended Drude approach presented in [31]. The dimensions of the simulation domain are chosen so large that the influence of the boundary conditions can be neglected. For reasons of simplicity, the temperature field at the boundaries is set to room temperature (298 K). Since this process model is based on the fundamental equations of laser-matter interaction, it can be applied to basically any laser-based material processing technique. Therefore, the model can be applied on evaporation driven processes like deep penetration welding [26, 32] or remote cutting [33] and non-evaporation driven such as brazing [34] or laser beam melting in powder bed [35]. In order to implement the different beam profiles in the simulation model, the measured intensity profiles of the beam profiles were used. The individual spots of the beam profiles of the first and zeroth order were approximated by a combination of two radially symmetric superGaussian functions. By an optimisation algorithm, the distribution factor
3. Results and discussion 3.1. Analyses of the DOE´s BP and performance For further experimental investigations the laser induced damage threshold of the DOEs is evaluated. By increasing the laser power in steps of 250 W the limit is determined to be 4.75 kW, which corresponds to a peak intensity of 3.5 kW/cm2 , agreeing well with results shown by Alexeev et al. [23]. The BPs of the DOEs were experimentally evaluated at different laser powers to investigate if any changes can be observed. As no significant changes could be observed, the diffraction efficiency was evaluated at a power of 1.5 kW and used for predicting the actually used power for welding in the experiments. The results are presented in Fig. 4 and Table 1. The losses for the defocussed beam can be attributed to reflection on the fused silica substrate used for the DOEs and the cover glass, as both do not have an anti-reflection-coating. The diffraction efficiency of the DOEs including the zeroth order, i. e. the power used for welding, was similar for all three designs at about 78% with about 1% in the zeroth order. 3.2. Process window for HCMW Bead-on-plate welds with a length of 40 mm and parameters shown in Table 2 have been produced. At first a combination of high feed rate and low laser power is chosen. Subsequently, the speed is reduced until 182
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Fig. 5. a) HCMW. Highly stable melt pool with RING Profile: Pw = 2.7 kW, v = 25 mm/s. The dashed ellipse indicated the size of the melt pool which is partly hidden by an oxide layer[vis1] b) Early TMW. The first turbulences in the melt pool occur. DEF with Pw = 2.6 kW, v = 400 mm/s.[vis2] c) Late TMW. The keyhole is about to form but collapses right afterwards. DEF with Pw = 2.6 kW, v = 200 mm/s. [vis3] d) A stable keyhole is formed. LINE with Pw = 2.0 kW, v = 100 mm/s. [vis4].
Table 1 Laser power in the working spot (Pw ) and peak intensity (IPeak ) of the used BP with respect to the laser input power (PL ). Profile
DEF RING POINTS LINE
Parameter
Value
PL Pw IPeak Pw IPeak Pw IPeak Pw IPeak
1 0.86 19.8 0.67 2.94 0.66 14.3 0.68 9.37
Since a significant part of the melt pool is still covered by the oxide layer its dimensions can only be evaluated through surface vibrations in the high-speed videos. The resulting actual melt pool is marked with a dashed red line in Fig. 5a). With increasing laser intensity, deformations on the melt pool surface caused by small areas of strong evaporations form. Those deformations can be seen by bypassing reflection on the melt pool (see Fig. 5 b). This early stage of TMW is taken as boarder between TMW and HCMW. When the intensity is high enough, the deformation of the melt pool caused by the evaporation pressure enlarges and a keyhole is formed. In the early stages the keyhole (indicated by a green arrow Fig. 5c) is instable and collapses right after formation resulting in a highly oscillating melt pool surface. As previously shown by Cho et al. in these cases the melt pool surface temperature and flow direction are oscillating with high frequency [38]. Therefore, this represents the most unstable welding regime. As soon as the laser intensity becomes higher, evaporation becomes stronger and with the resulting higher evaporation pressure the keyhole becomes more stable and remains open (see Fig. 5 d). The transition between TMW and HCMW, estimated by analysing of the high speed videos, is shown in Fig. 7. Within the limits of the accuracy of the measurements it can be said that DEF and POINTS share almost the same transition with POINTS having a slightly lower one. RING reached HCMW in all conducted experiments. Only at the highest used power and the lowest velocity, slightly increased movements of the reflections could be observed (see Fig. 5). LINE showed the lowest threshold for the transition between TMW and HCMW. Below a peak intensity of 1.9∙106 W/cm2 (see Fig. 7, purple dashed line) and above 25 mm/s all tested profiles are in HCMW. Differences are also visible in the high-speed images of POINTS and LINE shown in Fig. 8. For POINTS floating narrow oxide stripes inbetween the individual spots can still be observed (see Fig. 8b, yellow arrows), while in LINE a continuous oxide-free melt pool surface forms. For LINE also a dynamically reflecting area around the laser interaction zone develops. This might indicate a higher evaporation rate due to higher temperatures. The assumption of higher temperatures for LINE is supported by the larger dark area in front of the laser interaction zone indicating larger amount of melt below the oxide layer. The remaining floating narrow oxide stripes in-between the spots of POINTS and their absence for LINE can be reproduced and explained by numerical simulations. Since the employed process model does not take oxidation per se into account, due to the very small thickness of the aluminium oxide layer of only 25 – 50 Å it can be assumed that the layer will be molten as soon as its melting temperature is reached [39]. Fig. 9 shows lateral temperature profiles for both BPs. Right at the laser position (red line) the evaporation temperature is reached for the whole laser-interaction zone of LINE. For POINTS, the evaporation temperature is reached at the center of the individual spots, but in-between the spots the temperatures drop below the melting temperature of aluminium oxide Al2 O3 (Tmelt,Al2 O3 ) of 2339 K [39]. The lateral temperature profile of LINE 50 μm behind the point of maximum laser irradiation (yellow line), where only a negligible amount of radiation reaches the surface, shows that the temperature has dropped homogenously to around Tmelt,Al2 O3 . For POINTS the peak structure of the temperature
Unit 2 1.71 39.6 1.35 5.88 1.31 28.5 1.36 18.7
3 2.57 59.4 2.02 8.82 1.97 42.8 2.04 28.1
4 3.42 79.2 2.69 11.8 2.62 57.1 2.71 37.5
4.5 3.85 89.1 3.03 13.2 2.95 64.2 3.05 42.2
kW kW ∙105 W/cm2 kW ∙105 W/cm2 kW ∙105 W/cm2 kW ∙105 W/cm2
Table 2 Processing Parameters. Parameter
Value
Unit
Laser input power PL Traverse speed v
1 / 2 / 3 / 4/ 4.5 25 / 50 / 75 / 100 / 200 / 400 / 800 / 1200
kW mm/s
Fig. 6. Cross-section at the front section of the interaction zone of the velocity field of the melt pool with RING Profile. As can be seen, only very low velocities appear within the melt pool. The slightly higher velocities at the melt pool surface indicate that some, but still very limited evaporation takes place. Process parameters: Pw = 2.69 kW, v = 50 mm/s.
the melt pool becomes unstable. After that the next higher laser power with maximum speed is examined. With that procedure the process window for HCMW is determined. High-speed process monitoring offers the opportunity to conclude in which welding regime the process lies. Compared to the analysis of cross-sections, the captured high-speed videos show the melt pool dynamics over the complete weld seam [36, 37]. Starting instabilities can be observed and a clear line between transition mode (TMW) (see Fig. 5c) and HCMW (see Fig. 5a) welding can be drawn. Regarding the melt pool stability, the most stabile regime is the HCMW regime. As can be taken from the very stable melt pool surface in the highspeed videos (Fig. 5a), only a negligible amount of evaporation occurs in HCMW. This is confirmed by the simulative results, which also indicate that evaporative mass flux and evaporation-based recoil pressure are negligible. Simulations also show that barely any melt flow occurs, as can be taken from Fig. 6. 183
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Fig. 7. TMW / HCMW Transition. Comparison of different BPs. The purple dashed line in b) indicates the peak intensity for which all profiles were in HCMW.RING is not shown as this BP was always in HCMW.
Fig. 8. Comparison between POINTS (a and b) with Pw = 1.31 kW and LINE (c and d) with Pw = 1.36 kW; both at v = 50 mm/s. Images correspond to the views of both cameras (a and c camera 1, b and d camera 2) at the same point in time.
Fig. 9. Simulated lateral temperature profiles of a) POINTS and b) LINE at the laser position (dx = 0 μm, red), 50 μm behind the laser position (dx = 50 μm, yellow) and 150 μm behind the laser position (dx = 150 μm, green) with Pw = 1.3 kW; both at v = 50 mm/s. The blue line indicates the melting and the purple line the evaporation temperature of Al2 O3 .
Table 3 Average temperatures of the melt pool volume, -surface and average evaporation pressure of the melt pool surface in the numerical simulations for POINTS and LINE.
Melt pool volume temperature Melt pool surface temperature Melt pool surface evaporation pressure
POINTS
LINE
Unit
1131 ± 9 1626 ± 19 45,770 ± 603
1132 ± 9 1646 ± 23 46,016 ± 382
K K Pa
multiple cross-sections and the top view have been metallographically prepared. In all samples that could be produced with 25 mm/s in HCMW, that is about 56% faster than the current state of the art for this alloy [12], no hot cracks have been found using optical microscopy (see Fig. 10) and secondary electron imaging (data not shown). The crosssection consists mainly of equiaxial dendritic grains.
3.4. Melt-pool dimensions The cross-sections of the weld seams in HCMW have been analysed on two cross-sections and the depth DW and the area AW of the weld seam have been determined. The results are summarised in Fig. 11 and the full data can be found in the appendix (Fig. 18-21). As RING stays in HCMW for higher laser powers, deeper weld seams can be achieved. With our process parameters an up to three times larger weld depth compared to DEF could be achieved, resulting in a dramatically larger weld area. Comparing the obtained weld areas versus Pw at fixed traverse speeds up to 400 mm/s a linear correlation regardless of the used BP is found (see Fig. 11a). Although RING remains in HCMW for all investigated process parameters, simulations indicate that the absorbed laser power is significantly higher (about 35% compared to LINE and POINTS) than for the other BPs when the oxide layer is completely molten. Therefore, the predicted melt pool volume is larger. This is supported by the area of the weld seam in the cross-section views being the largest for RING.
profile prevails leaving the oxide stripes intact. Only for temperature profiles far below Tmelt,Al2 O3 the temperature profile has smeared such that the point structure has dissolved. Simulations support also the earlier onset of evaporation and TMW. Fig. 11 shows that for LINE higher maximum and mean temperatures are reached. Table 3 lists the average temperatures of the melt pool volume and surface for both profiles. It shows that although the average temperature of the melt pool volume are comparable, the averaged surface temperature of LINE is higher, resulting in a higher averaged evaporation pressure of the surface for LINE and therefore a more pronounced overall evaporation for LINE. 3.3. Hot crack tendency When working with aluminium copper alloys special attention has to be paid to hot cracking. For determining the existence of hot cracks, 184
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Fig. 10. No visible hot cracks in HCMW; a) and c) crosssection and polished top view of RING with Pw = 3.0 kW and v = 25 mm/s; b) and d) cross-section and top view of Points with Pw = 0.7 kW and v = 25 mm/s.
Fig. 11. Weld area and depth at a fixed v of a) and b) 50 mm/s, and c) and d) 800 mm/s.
At higher traverse speed the melt pool cross-section areas differ depending on the used BP (see Fig. 11c and d). The high-speed videos provide an explanation for this behaviour: The specific point energy (SPE ∝ Iint ∙w2 ∙/v) [40], which corresponds to the accumulated local energy from the laterally shaped intensity, is too low to melt the oxide layer. The layer remains partly (see Fig. 13b – d) or completely unmolten (see Fig. 13a) on the hidden melt pool (dashed red lines), causing a lower absorption [41] of laser irradiation compared to the liquid aluminium alloy [42]. At POINTS and DEF only the centers of the spots have a high enough SPE to melt the oxide layer. For LINE the SPE is too low to melt the layer but still a huge bow wave most likely increasing the absorption by changing the angle of incidence and enlarging the area in the laser beam irradiated zone can be observed. This might be the reason for LINE having the highest weld area compared to the other intensity profiles with comparable Pw . This behaviour is also visible when comparing the welt pool length (see Fig. 12). LINE has a longer weld pool compared to POINTS for all investigated traverse speeds. This is also seen in simulations, where the temperature drops slower with increasing distance to the interaction zone (see Fig. 9). The decrease of the length for POINTS at high speeds, which is not found for LINE, further indicates that less energy is absorbed when the oxide layer is not completely molten. Additionally the overall melt pool size and geometry, especially the shape of the re-solidifying front, varies significantly with the BP at constant process parameters. While the BP translates directly to the shape of the front of the melt pool, the rear end cannot be directly related to the BP due to heat conduction.
Fig. 12. Comparison of the weld pool length lW of LINE and POINTS at constant Pw = 1.3 kW.
Inspecting the cross-sections of the welds created by the different intensity profiles, a change in the form is visible as well. At larger melt pool sizes (see Fig. 14), the welds look similar for the shaped profiles as the initial intensity profile integrated in traverse direction and therefore the temperature field becomes less distinct due to heat conduction inside the melt pool. In contrast, with decreasing melt pool size (see Fig. 15) the original intensity profile becomes more visible. DEF creates a half circle like weld seam. The three other profiles produce rather flat sections. The weld cross-section of POINTS has a rectangular shape with quite sharp edges and the individual spots shining through at smallest melt pools (see Fig. 15b red arrows). The cross-section of RING comes with slightly 185
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Fig. 13. Fully and partly hidden melts pools below oxide layer at high traverse speeds (800 mm/s); a) RING Pw = 1.35 kW [vis5]; b) POINTS Pw = 1.31 kW [vis6]; c) DEF Pw = 1.71 kW [vis7]; d) LINE Pw = 1.35 kW [vis8].
Fig. 14. Etched cross-section at v = 50 mm/s a) RING Pw = 0.67 kW b) POINTS Pw = 0.66 kW c) DEF Pw = 0.86 kW d) LINE Pw = 0.68 kW.
Fig. 15. Etched cross-section at v = 800 mm/s a) RING Pw = 1.35 kW b) POINTS Pw = 1.31 kW c) DEF Pw = 1.71 kW d) LINE Pw = 1.35 kW.
Fig. 16. Unpolished bright-field top-view images of a) RING with Pw = 1.35 kW, v = 25 mm/s and a clearly visible grain structure b) POINTS with Pw = 1.31 kW, v = 50 mm/s and oxide layer leftovers c) LINE with Pw = 1.35 kW, v = 50 mm/s d) DEF Pw = 1.71 kW, v = 100 mm/s.
rounder edges and is slightly deeper at both sides (Fig. 15a). LINE has flat tapered flanks induced by the experimental BP. 3.5. Surface roughness The lowest SA was obtained with RING. The low melt pool dynamics and therefore, a high melt pool stability enables a surface finish where even grains become visible (see Fig. 16 a). POINTS suffers from its passage channels. The oxide layer in HCMW remains on top of the melt pool and reduces the surface quality (see Fig. 16b). LINE and DEF need higher traverse speeds to stay in HCMW. Consequently, the dynamics inside the melt pool are considerably higher and the surface of the weld is permeated with solidified waves (see Fig. 16c and d). The effect of the traverse speed and laser power on the resulting surface roughness is subsequently exemplarily discussed on the RING profile because of its large process window for HCMW (see Fig. 17). First
Fig. 17. Achieved surface roughness with RING depending on Pw and v.
of all, a high enough SPE is needed to break the oxide layer which is not the case for Pw = 0.67 kW at all traverse speeds and Pw = 1.35 kW and 186
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traverse speed higher than 200 mm/s. At the lowest applied laser power Pw = 0.67 kW an interesting phenomenon can be observed. Starting at 25 mm/s a melt pool in size equal to RING is created leading to a good surface quality most likely by reformation of the sand blasted surface by surface tension of the melt pool below the oxide layer. At 50 mm/s vibrations going throw the melt pool can be observed causing a local SA peek. Presumably the liquid is partly solidified in the central area of RING and remolten at the second interaction with the laser beam. That sequence of melting, solidification and remelting might be the reason for the observed behaviour. At higher speeds the centre for the RING stays solid and the surface roughness is getting better again. This explanation should be valid for the peak at Pw = 1.35 kW and v = 400 mm/s as well. As soon as the oxide layer is broken the surface roughness is positively affected. SA with the curvature of the weld seam having the highest impact down to 0.2 μm at v = 25 mm/s, 50 mm/s and Pw = 2.02 kW can be obtained. Those values are normally only achieved by laser polishing [43]. With increasing traverse speed the surface quality is decreasing again due to the increasing melt pool dynamics.
and length, can be influenced, where the front of the melt pool mimics the BP directly, while the re-solidification front cannot be directly deduced. Numerical simulations predict that the coupling efficiency of the laser power can be enhanced by beam shaping which is in agreement with experimental results showing different cross-sectional areas of the weld seam for different profiles at the same processing parameters. Simulations also show that although the mean temperature of the melt pool volume is comparable, the mean surface temperature and resulting evaporation pressure can differ between the BPs. Therefore, this work shows a huge potential of beam shaping for optimising laser welding processes. Beam shaping promises to improve this process due to larger attainable scanning speeds with large crosssection areas and melt pools with low process dynamics. To achieve the goal of beam profiles tailored to the individual requirements of each welding process further studies investigating more elaborate beam profiles and their influence on process dynamics and the quality of the weld are needed.
4. Conclusion
Acknowledgements
Beam shaping by diffractive optical elements enables the possibility to influence melt pool size, shape, and dynamics. Therefore, the process window for HCMW can be significantly increased by using shaped intensity profiles and surface qualities comparable to laser polishing can be obtained. In the presented study, a ring-profile had the largest process window and yielded the best surface qualities. In contrast to welding with a Gaussian spot, for beam shaping a high peak intensity does not necessarily lead to an instable melt pool, but the distribution itself plays the major role as could be shown by the different process windows for LINE and POINTS. The size of the melt pool, i. e. the depth
The authors want to thank the German Research Foundation (DFG) for funding the Collaborative Research Center 814 (CRC 814) – Additive Manufacturing, sub-project A5. The authors gratefully acknowledge the support provided by the Erlangen Graduate School in Advanced Optical Technologies. The work was performed with the support of the Ministry of Education and Science of the Russian Federation, decree N220, state Contract No. 14.Z50.31.0023 Appendix
Fig. 18. Weld dimensions of RING. Area AW and depth DW versus v for Pw .
Fig. 19. Weld dimensions of POINTS. AW and DW versus v for different Pw .
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Fig. 20. Weld dimensions of LINE. AW and DW versus v for different Pw .
Fig. 21. Weld dimensions of DEF. AW and DW versus v for different Pw .
Supplementary materials
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Optics and Lasers in Engineering journal homepage: www.elsevier.com/locate/optlaseng
Shaped laser beam profiles for heat conduction welding of aluminium-copper alloys Michael Rasch a,b,c,∗, Clemens Roider a, Stefanie Kohl a,c, Johannes Strauß a,c, Niklas Maurer a, Konstantin Y. Nagulin d, Michael Schmidt a,b,c a
Institute of Photonic Technologies, Friedrich-Alexander University Erlangen-Nürnberg, Erlangen, Germany Collaborative Research Center CRC 814 – Additive Manufacturing, Friedrich-Alexander University Erlangen-Nürnberg, Erlangen, Germany c Erlangen Graduate School in Advanced Optical Technologies (SAOT), Erlangen, Germany d Kazan National Research Technical University, Kazan, Russia b
a r t i c l e
i n f o
Keywords: Heat conduction mode laser beam welding Surface roughness Beam shaping Diffractive optical elements Aluminium alloys Numerical process simulation
a b s t r a c t Heat conduction welding is often used where weld seams with a high surface quality and the exact retention of the chemical composition are needed. This study investigates the influence of intensity profiles in laser beam welding with laser powers up to 3.2 kW. The resulting process windows for heat conduction mode welding, the melt pool shape, the process stability and the dynamics, and the produced surface roughness are analysed. Therefore, three different intensity profiles are created with diffractive optical elements and the process dynamics are monitored with two high-speed cameras. Before analysing etched cross sections with respect to microstructural properties, the surface roughness is measured on all produced samples with a laser scanning microscope. With beam shaping, we found that a high peak intensity does not necessarily lead to an instable melt pool and that the distribution itself plays the major role. The use of shaped beam profiles leads to a more stable process and an enlarged heat conduction process regime compared to a defocussed multimode spot of the same size. Applying shaped laser beam profiles surface roughness close to laser polished surfaces can be achieved while also simultaneously enlarging the weld cross section area.
1. Introduction High strength aluminium alloys are commonly used in aerospace and automobile applications [1]. Due to certain material properties the connection of two parts not using force or form-fit but welding technologies is challenging. Especially EN AW-2xxx alloys suffer from a high risk of hot cracking due to their huge solidification interval, with EN AW-2024 being one of the alloys with the highest hot crack susceptibility suffering from a huge solidification interval [2]. One possibility to deal with this problem is the use of non-melting technologies like friction stir welding (FSW). FSW is advantageous because the process is carried out below the melting temperature. In 1999 crack free welding seams were successfully shown [3]. However, these welding seams are rather large and suffer from a rough surface and therefore, the corrosion resistance is classified as poor. Another drawback is the limitation to the clamping. Recent research is focussed on improving the weld surface quality by laser polishing. Kalita could show that re-melting using a diode laser with laser powers between 700 W and 800 W and a traverse speed of 1.66 mm/s increases the resistance
∗
of EN AW-2024 FSW-welds to pit growth without producing any cracks [4]. Avoiding hot cracks using melting technologies is even more challenging. While the material is molten recrystallization is taking place. Dendritic grain morphology must be avoided because of its well-known high tendency for hot cracks and crack propagation [5]. In electron beam welding a weld seam for both regimes – heat conduction welding and deep welding – consisting of equiaxial grains due to the low temperature gradient but high solidification rate can easily be produced [6]. In laser beam welding most attention is paid to the keyhole welding process due to the higher achievable aspect ratio. The first common solution was to add a filler material like EN AW-4xxx to produce an alloy with an higher silicon content in the weld seam with a lower tendency for hot cracking [5]. Similarly, adding grain refiners like TiB2 leads to fine equiaxial grain structures without any cracks [7]. Successful welding without any additives was achieved using low traverse speed of 40 mm/s and a high laser power of 2.75 kW focussed on a spot with a diameter of 0.45 mm to 0.60 mm [8]. The weld seam consisted mostly of equiaxial dendritic grains. It seems that the high melt dynamics that occur in the keyhole regime and a relatively low temperature gradient
Corresponding author. E-mail address: [email protected] (M. Rasch).
https://doi.org/10.1016/j.optlaseng.2018.11.025 Received 9 July 2018; Received in revised form 11 November 2018; Accepted 26 November 2018 0143-8166/© 2018 Elsevier Ltd. All rights reserved.
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due to the low traverse speed lead to strong constitutional undercooling and the preferred grain morphology is formed. Higher traverse speeds of up to 120 mm/s could be reached using an additional heat source reducing the effective thermal stress on the weld seam due to a lower temperature gradient [9]. In this case the grain structure is fully dendritic. The downsides of keyhole welding are rough surfaces, a change in the chemical composition due to a selectively higher evaporation rate of certain alloying constituents and vapour pores [10]. Heat conduction mode welding (HCMW) is known to be the most stable process regime in laser welding, but only little research has been carried out on this process regime (and in most cases HCMW is not even mentioned explicitly). A compressive review with different aluminium alloys was published by Quintino and Assunção pointing out the clearly above mentioned advantages [11]. Sánchez-Amaya et al. are to the authors’ knowledge the only ones investigating 2xxx alloys in HCMW. They used a diode laser with a maximum power of 2.75 kW and a spot size of 2.2 × 1.7 mm2 to create weld seams on six different high strength aluminium alloys, one of them being EN AW-2024. They found that a minimum power of 2 kW and maximum speed of 16 mm/s is needed to create a stable melt pool and crack free weld seams. The alloying constituents proved to play a major role in weldability and hot crack susceptibility as well [12]. In 1999, Zhao et al. predicted that the control of the melt pool geometry will be an important tool for influencing the evaporation of alloying elements and hot crack susceptibility [13]. The first experimental implementation in keyhole welding of AA 6016 was done using a twin laser system equipped with a CO2 laser as primary source and a pulsed Nd: YAG laser as secondary source. It could be shown that at a rather high welding speed of 60 mm/s and CO2 -laser power of 1.7 kW, the use of the pulsed Nd: YAG laser with a pulse energy of 3 J at a repetition frequency of 150 Hz and a pulse duration of 2 ms could effectively prevent the formation of cracks. It was found that the pulsed nature of the Nd: YAG laser and the significant decrease of the thermal gradient promotes the columnar to equiaxial transition and therefore the formation of equiaxial-globular grains [14]. Hansen et al. showed that with multispot keyhole welding of EN DC01 and EN 1.4301 using diffractive optical elements weld seams with an almost rectangular cross section can be obtained [15]. Sundqvist et al. made some studies focussing on the numerical and analytical optimization of laser spot welding with shaped laser beams [16]. Funck et al. did experimental work showing in laser spot welding with a ring profile that mainly the outer diameter affects the convection in the melt pool [17]. As keyhole welding is suffering from poor surface quality this process cannot be used when the surface of weld seams is of interest. In contrast HCMW yields high quality weld surfaces but at the moment suffers from slow speeds. Therefore, the goal of the presented work is to clearly identify the process windows for welding with different shaped laser beams to achieve higher welding speeds. Furthermore, the influence of the beam shape on melt pool dynamics and shape and on the obtained surface quality is determined. Comparing experimental observations with temperature and evaporation predictions from simulations yields further insight and enhances process understanding when using shaped beam profiles.
Fiber plug Fiber end Concav lens: -75 mm JGS1 Convex lens: 75 mm BK7 Area for DOE convex lens: 100 mm BK7 20 mm Fig. 1. Customised welding optics.
of the DOE in the optical path, a collimated laser beam is preferred. Using the knife edge method and fitting with the Gaussian error-function, the beam diameter with dismounted focussing lens is determined at a distance of 100 mm and 300 mm from the lower edge of the optics for verifying the collimation (see Fig. 1). For further calculation a collimated Gaussian beam with a diameter of 18.6 mm is assumed. The diameter was chosen in order to keep the peak intensity of the beam as low as possible on the 25.4 mm DOE, simultaneously avoiding any significant amount of laser power hitting the DOE mount. Another important fact is the caustic of the unshaped laser beam with assembled focussing lens for determining the Rayleigh range, the focus profile and the position of the focal plane. Therefore, the 3D beam shape is determined with a PRIMES MicroSpotMonitor. A calibration procedure was applied for quantifying the intensity distributions. The beam waist was found to be w = 57 μm with an almost perfect flat top shape as the fibre end is imaged onto the sample with a slight demagnification (see Fig. 2). The final experimental setup is shown in Fig. 3. The welding optics is mounted at a fixed position on a welding desk. Additionally, two highspeed cameras (Phantom v1210, Vision Research) are used in order to monitor the melt pool. While camera 1 is monitoring the melting process in welding direction, camera 2 is positioned perpendicular to camera 1 resulting in sharp views on different regions of the melt pool. For illumination, a ns-pulsed diode laser (Cavilux HF, 𝜆 = 808 nm, Pavg = 500 W) is used. A sandblasted EN AW-2024 sheet metal (56.5 × 18.7 × 3.8 mm3 , SA = 1.2 μm) is placed inside a small process chamber, which is moved by an industrial robot (KR30, Kuka) underneath the stationary laser beam. To ensure a constant movement below the welding optics, an acceleration and a deceleration section before and after the welding of 500 mm and a section with constant velocity of 100 mm below the welding optics are implemented. The process chamber is equipped with an easily exchangeable 1 mm thick cover glass (Nexterion D263T, Schott; transmittance including Fresnel reflection losses is 91.85% [18]) 11 mm above the sheet metal. To prevent pollution of the glass and oxidation of the weld seam, the processing chamber is constantly flushed with argon.
2. Methods 2.1. Experimental setup
2.2. Determination of the surface roughness, metallurgical preparation and analysis of the melt pool length
For the following experiments a disk laser with a maximum output power of 8 kW (TruDisk 8001, Trumpf, 𝜆 = 1030 nm, 150 μm fibre, NA = 0.067) and a customised optics with a 100 mm focussing lens have been used (see Fig. 1). As the process is accompanied by an extensive heat build-up the fiber plug is water cooled, while the lenses as well as the diffractive optical element (DOE) are air cooled. For the design of the DOEs it is important to know the diameter and intensity distribution of the laser beam at the DOE position. To reduce axial alignment errors
The arithmetical mean height of the surface SA of produced specimens is analysed with a confocal laser scanning microscope (LEXT OLS4000, Olympus) according to DIN EN ISO 25,178–2 [19] with a cut-off wavelength 𝜆c of 800 μm in a rectangle of 394 × 564 μm2 at the weld centre. Afterwards the samples are cut into three sections. While two sections are metallographically prepared for further analyses, one section is used to analyse the top view. The samples are ground with up 180
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Fig. 2. a) Beam Caustic b) beam profile in the focal plane c) defocussed beam profile d) crosssection in the focal plane e) cross-section through defocussed beam (w ∼ 450 μm).
Fig. 3. Experimental setup. Green Box: Miniature process chamber with an inert gas flow.
to 4000 grit and subsequently polished using a final grain size of 1 𝜇m. To determine the presence of hot cracks, unetched metallographically prepared cross-sections are investigated by light microscopy. For characterizing the dimensions of the weld seams and the grain morphology, etched and electrolytically polished specimens are analyzed. The length of the melt pool lw is determined by analysis of the high speed videos. lw is defined from the last interaction of the laser beam with the melt pool to complete solidification at the symmetry axis.
The corresponding DOEs were calculated using a weighted Gerchberg-Saxton algorithm [21] for POINTS and LINE and the classical Gerchberg-Saxton algorithm [22] for RING. The fabrication procedure follows the on presented by Alexeev et al. [23] with a pixel pitch of 10 μm and a layer with an optical thickness corresponding to 𝜆/2 at 1030 nm. As with this method only binary DOEs can be produced, the desired BPs were also chosen to be point-symmetric. Therefore, the theoretically highest possible diffraction efficiency is 81% with the remaining 19% being distributed in higher diffraction orders. The first row in Fig. 4 shows simulated predictions of the BPs, while the following rows show experimentally measured intensity distributions, integrated intensity, and cross-section views. Note the “ghost spots” visible for LINE at the left/right edge of the image which are caused by the periodic structure [24]. Also a weak zeroth order is visible in the cross-section of RING due to an imperfect DOE fabrication. Imperfection will also lead to stronger higher diffraction orders and stray light, therefore reducing the diffraction efficiency.
2.3. Chosen beam profiles In this study three different beam profiles (BP), which are depicted in Fig. 4, and a defocussed spot (DEF) as reference (see Fig. 2c and e) with the same constant “spot” diameter of around 900 μm, that is about nine times the minimal unmodified spot diameter, are investigated. Note that DEF is not a Gaussian profile but looks like a tipi shape resulting in nearly linear flanks and a pointy maximum. The intention of the first BP, called LINE, is to keep the intensity as well as the integrated intensity along the welding direction constant (see Fig. 4h and k). Recent publications have shown that line like beam profiles can lead to evaporation induced “bow wave” like melt flows in front of the interaction zone [20]. To reduce these effects the second BP (POINTS) was chosen. Instead of having a straight line, the beam profile consists of five spots with a distance of 200 μm between each other. The peak intensity of POINTS is roughly twice as high as in LINE but the space between the spots might create a deformation free region where the melt can flow unhindered (see Fig. 4l). The third BP is a RING, which has a radius of 400 μm and a width of 110 μm, corresponding to the unmodified spot diameter. RING has a higher integrated intensity along the welding direction (Iint ) on the left and the right (see Fig. 4g). That is meant to take the larger heat dissipation at the edges of the melt pool into consideration. Therefore, the temperature should be more equally distributed over the melt pool.
2.4. Numerical simulations Multiphysical process simulations were carried out to complement the experimental results and analyse quantities that are experimentally not accessible. The thermo-fluiddynamic numerical model is based on the finite volume method (FVM) open source framework OpenFOAM® (Open Field Operation and Manipulation) [25]. Within this programming environment, the physics of laser-material interaction were implemented in the form of partial differential equations (PDEs) to model the laser welding process. Detailed information on the physics of numerical model and its verification can be taken from [26]. Since this particular paper addresses HCMW, multiple reflections could be neglected and were not considered in the calculations to save computational time. Although evaporation is not expected to be a dominant effect in HCMW, it still is an important process characteristic since the evaporation driven 181
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Fig. 4. a-c) calculated beam profiles; d-f) measured beam profiles; g-i) integrated intensity in traverse direction of the measured profiles; j-l) cross-section through the measured profiles; each row was normalised to the highest value for adequately displaying the relative strengths of the different profiles. White arrows mark the welding direction.
recoil pressure defines the border of the process regime towards laser beam deep-penetration welding. Therefore surface tension, evaporation and the resulting recoil pressure are included in the model to be able to reconstruct possible indentations in the melt pool. The included evaporation model calculates the temperature-dependent evaporation pressure as given in [27]: ( ( )) 𝐻𝐵 𝑀 1 1 𝑝𝑣𝑎𝑝 (𝑇 ) = 0.54 ⋅ 𝑃𝑎 ⋅ exp − . (1) 𝑅 𝑇𝐵 𝑇
between the two super-Gaussian functions, the beam radii and the order of the super-Gaussian functions were calculated and the calculated parameters were used as input parameters in the simulation model. Intensities in higher orders and “ghost spots” were neglected in the simulations. As expected, the fitting of the beam profile of the defocussed spot turned out to be more challenging, since its profile was super-Gaussian in its focus. In contrast to Gaussian beams, super-Gaussian beams do not keep their beam profile constant during defocus. For this beam, only one super-Gaussian could be fitted with an effective exponent of n = 1.56, resulting in a rather pointy “sub” -Gaussian function.
In Eq. 1 HB and TB denote the enthalpy and temperature of evaporation, respectively, M denotes the material’s molar mass, R the gas constant, and Pa the ambient pressure. The factor of about 0.54 originates from the consideration of recondensation within the Knudsen layer. It holds for the pressure and evaporation conditions of laser material processing and is widely used for laser welding simulations [28]. All optical, fluid dynamical and thermo-physical material properties of the alloy used are included in the model considering their temperature-dependence. The thermo-physical properties are extracted from the CalPhaD-based (Calculation of Phase Diagrams) material database JMatPro [29] based on the chemical composition of the alloy [30]. The optical properties and their temperature dependence are calculated dependent based on the extended Drude approach presented in [31]. The dimensions of the simulation domain are chosen so large that the influence of the boundary conditions can be neglected. For reasons of simplicity, the temperature field at the boundaries is set to room temperature (298 K). Since this process model is based on the fundamental equations of laser-matter interaction, it can be applied to basically any laser-based material processing technique. Therefore, the model can be applied on evaporation driven processes like deep penetration welding [26, 32] or remote cutting [33] and non-evaporation driven such as brazing [34] or laser beam melting in powder bed [35]. In order to implement the different beam profiles in the simulation model, the measured intensity profiles of the beam profiles were used. The individual spots of the beam profiles of the first and zeroth order were approximated by a combination of two radially symmetric superGaussian functions. By an optimisation algorithm, the distribution factor
3. Results and discussion 3.1. Analyses of the DOE´s BP and performance For further experimental investigations the laser induced damage threshold of the DOEs is evaluated. By increasing the laser power in steps of 250 W the limit is determined to be 4.75 kW, which corresponds to a peak intensity of 3.5 kW/cm2 , agreeing well with results shown by Alexeev et al. [23]. The BPs of the DOEs were experimentally evaluated at different laser powers to investigate if any changes can be observed. As no significant changes could be observed, the diffraction efficiency was evaluated at a power of 1.5 kW and used for predicting the actually used power for welding in the experiments. The results are presented in Fig. 4 and Table 1. The losses for the defocussed beam can be attributed to reflection on the fused silica substrate used for the DOEs and the cover glass, as both do not have an anti-reflection-coating. The diffraction efficiency of the DOEs including the zeroth order, i. e. the power used for welding, was similar for all three designs at about 78% with about 1% in the zeroth order. 3.2. Process window for HCMW Bead-on-plate welds with a length of 40 mm and parameters shown in Table 2 have been produced. At first a combination of high feed rate and low laser power is chosen. Subsequently, the speed is reduced until 182
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Fig. 5. a) HCMW. Highly stable melt pool with RING Profile: Pw = 2.7 kW, v = 25 mm/s. The dashed ellipse indicated the size of the melt pool which is partly hidden by an oxide layer[vis1] b) Early TMW. The first turbulences in the melt pool occur. DEF with Pw = 2.6 kW, v = 400 mm/s.[vis2] c) Late TMW. The keyhole is about to form but collapses right afterwards. DEF with Pw = 2.6 kW, v = 200 mm/s. [vis3] d) A stable keyhole is formed. LINE with Pw = 2.0 kW, v = 100 mm/s. [vis4].
Table 1 Laser power in the working spot (Pw ) and peak intensity (IPeak ) of the used BP with respect to the laser input power (PL ). Profile
DEF RING POINTS LINE
Parameter
Value
PL Pw IPeak Pw IPeak Pw IPeak Pw IPeak
1 0.86 19.8 0.67 2.94 0.66 14.3 0.68 9.37
Since a significant part of the melt pool is still covered by the oxide layer its dimensions can only be evaluated through surface vibrations in the high-speed videos. The resulting actual melt pool is marked with a dashed red line in Fig. 5a). With increasing laser intensity, deformations on the melt pool surface caused by small areas of strong evaporations form. Those deformations can be seen by bypassing reflection on the melt pool (see Fig. 5 b). This early stage of TMW is taken as boarder between TMW and HCMW. When the intensity is high enough, the deformation of the melt pool caused by the evaporation pressure enlarges and a keyhole is formed. In the early stages the keyhole (indicated by a green arrow Fig. 5c) is instable and collapses right after formation resulting in a highly oscillating melt pool surface. As previously shown by Cho et al. in these cases the melt pool surface temperature and flow direction are oscillating with high frequency [38]. Therefore, this represents the most unstable welding regime. As soon as the laser intensity becomes higher, evaporation becomes stronger and with the resulting higher evaporation pressure the keyhole becomes more stable and remains open (see Fig. 5 d). The transition between TMW and HCMW, estimated by analysing of the high speed videos, is shown in Fig. 7. Within the limits of the accuracy of the measurements it can be said that DEF and POINTS share almost the same transition with POINTS having a slightly lower one. RING reached HCMW in all conducted experiments. Only at the highest used power and the lowest velocity, slightly increased movements of the reflections could be observed (see Fig. 5). LINE showed the lowest threshold for the transition between TMW and HCMW. Below a peak intensity of 1.9∙106 W/cm2 (see Fig. 7, purple dashed line) and above 25 mm/s all tested profiles are in HCMW. Differences are also visible in the high-speed images of POINTS and LINE shown in Fig. 8. For POINTS floating narrow oxide stripes inbetween the individual spots can still be observed (see Fig. 8b, yellow arrows), while in LINE a continuous oxide-free melt pool surface forms. For LINE also a dynamically reflecting area around the laser interaction zone develops. This might indicate a higher evaporation rate due to higher temperatures. The assumption of higher temperatures for LINE is supported by the larger dark area in front of the laser interaction zone indicating larger amount of melt below the oxide layer. The remaining floating narrow oxide stripes in-between the spots of POINTS and their absence for LINE can be reproduced and explained by numerical simulations. Since the employed process model does not take oxidation per se into account, due to the very small thickness of the aluminium oxide layer of only 25 – 50 Å it can be assumed that the layer will be molten as soon as its melting temperature is reached [39]. Fig. 9 shows lateral temperature profiles for both BPs. Right at the laser position (red line) the evaporation temperature is reached for the whole laser-interaction zone of LINE. For POINTS, the evaporation temperature is reached at the center of the individual spots, but in-between the spots the temperatures drop below the melting temperature of aluminium oxide Al2 O3 (Tmelt,Al2 O3 ) of 2339 K [39]. The lateral temperature profile of LINE 50 μm behind the point of maximum laser irradiation (yellow line), where only a negligible amount of radiation reaches the surface, shows that the temperature has dropped homogenously to around Tmelt,Al2 O3 . For POINTS the peak structure of the temperature
Unit 2 1.71 39.6 1.35 5.88 1.31 28.5 1.36 18.7
3 2.57 59.4 2.02 8.82 1.97 42.8 2.04 28.1
4 3.42 79.2 2.69 11.8 2.62 57.1 2.71 37.5
4.5 3.85 89.1 3.03 13.2 2.95 64.2 3.05 42.2
kW kW ∙105 W/cm2 kW ∙105 W/cm2 kW ∙105 W/cm2 kW ∙105 W/cm2
Table 2 Processing Parameters. Parameter
Value
Unit
Laser input power PL Traverse speed v
1 / 2 / 3 / 4/ 4.5 25 / 50 / 75 / 100 / 200 / 400 / 800 / 1200
kW mm/s
Fig. 6. Cross-section at the front section of the interaction zone of the velocity field of the melt pool with RING Profile. As can be seen, only very low velocities appear within the melt pool. The slightly higher velocities at the melt pool surface indicate that some, but still very limited evaporation takes place. Process parameters: Pw = 2.69 kW, v = 50 mm/s.
the melt pool becomes unstable. After that the next higher laser power with maximum speed is examined. With that procedure the process window for HCMW is determined. High-speed process monitoring offers the opportunity to conclude in which welding regime the process lies. Compared to the analysis of cross-sections, the captured high-speed videos show the melt pool dynamics over the complete weld seam [36, 37]. Starting instabilities can be observed and a clear line between transition mode (TMW) (see Fig. 5c) and HCMW (see Fig. 5a) welding can be drawn. Regarding the melt pool stability, the most stabile regime is the HCMW regime. As can be taken from the very stable melt pool surface in the highspeed videos (Fig. 5a), only a negligible amount of evaporation occurs in HCMW. This is confirmed by the simulative results, which also indicate that evaporative mass flux and evaporation-based recoil pressure are negligible. Simulations also show that barely any melt flow occurs, as can be taken from Fig. 6. 183
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Fig. 7. TMW / HCMW Transition. Comparison of different BPs. The purple dashed line in b) indicates the peak intensity for which all profiles were in HCMW.RING is not shown as this BP was always in HCMW.
Fig. 8. Comparison between POINTS (a and b) with Pw = 1.31 kW and LINE (c and d) with Pw = 1.36 kW; both at v = 50 mm/s. Images correspond to the views of both cameras (a and c camera 1, b and d camera 2) at the same point in time.
Fig. 9. Simulated lateral temperature profiles of a) POINTS and b) LINE at the laser position (dx = 0 μm, red), 50 μm behind the laser position (dx = 50 μm, yellow) and 150 μm behind the laser position (dx = 150 μm, green) with Pw = 1.3 kW; both at v = 50 mm/s. The blue line indicates the melting and the purple line the evaporation temperature of Al2 O3 .
Table 3 Average temperatures of the melt pool volume, -surface and average evaporation pressure of the melt pool surface in the numerical simulations for POINTS and LINE.
Melt pool volume temperature Melt pool surface temperature Melt pool surface evaporation pressure
POINTS
LINE
Unit
1131 ± 9 1626 ± 19 45,770 ± 603
1132 ± 9 1646 ± 23 46,016 ± 382
K K Pa
multiple cross-sections and the top view have been metallographically prepared. In all samples that could be produced with 25 mm/s in HCMW, that is about 56% faster than the current state of the art for this alloy [12], no hot cracks have been found using optical microscopy (see Fig. 10) and secondary electron imaging (data not shown). The crosssection consists mainly of equiaxial dendritic grains.
3.4. Melt-pool dimensions The cross-sections of the weld seams in HCMW have been analysed on two cross-sections and the depth DW and the area AW of the weld seam have been determined. The results are summarised in Fig. 11 and the full data can be found in the appendix (Fig. 18-21). As RING stays in HCMW for higher laser powers, deeper weld seams can be achieved. With our process parameters an up to three times larger weld depth compared to DEF could be achieved, resulting in a dramatically larger weld area. Comparing the obtained weld areas versus Pw at fixed traverse speeds up to 400 mm/s a linear correlation regardless of the used BP is found (see Fig. 11a). Although RING remains in HCMW for all investigated process parameters, simulations indicate that the absorbed laser power is significantly higher (about 35% compared to LINE and POINTS) than for the other BPs when the oxide layer is completely molten. Therefore, the predicted melt pool volume is larger. This is supported by the area of the weld seam in the cross-section views being the largest for RING.
profile prevails leaving the oxide stripes intact. Only for temperature profiles far below Tmelt,Al2 O3 the temperature profile has smeared such that the point structure has dissolved. Simulations support also the earlier onset of evaporation and TMW. Fig. 11 shows that for LINE higher maximum and mean temperatures are reached. Table 3 lists the average temperatures of the melt pool volume and surface for both profiles. It shows that although the average temperature of the melt pool volume are comparable, the averaged surface temperature of LINE is higher, resulting in a higher averaged evaporation pressure of the surface for LINE and therefore a more pronounced overall evaporation for LINE. 3.3. Hot crack tendency When working with aluminium copper alloys special attention has to be paid to hot cracking. For determining the existence of hot cracks, 184
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Fig. 10. No visible hot cracks in HCMW; a) and c) crosssection and polished top view of RING with Pw = 3.0 kW and v = 25 mm/s; b) and d) cross-section and top view of Points with Pw = 0.7 kW and v = 25 mm/s.
Fig. 11. Weld area and depth at a fixed v of a) and b) 50 mm/s, and c) and d) 800 mm/s.
At higher traverse speed the melt pool cross-section areas differ depending on the used BP (see Fig. 11c and d). The high-speed videos provide an explanation for this behaviour: The specific point energy (SPE ∝ Iint ∙w2 ∙/v) [40], which corresponds to the accumulated local energy from the laterally shaped intensity, is too low to melt the oxide layer. The layer remains partly (see Fig. 13b – d) or completely unmolten (see Fig. 13a) on the hidden melt pool (dashed red lines), causing a lower absorption [41] of laser irradiation compared to the liquid aluminium alloy [42]. At POINTS and DEF only the centers of the spots have a high enough SPE to melt the oxide layer. For LINE the SPE is too low to melt the layer but still a huge bow wave most likely increasing the absorption by changing the angle of incidence and enlarging the area in the laser beam irradiated zone can be observed. This might be the reason for LINE having the highest weld area compared to the other intensity profiles with comparable Pw . This behaviour is also visible when comparing the welt pool length (see Fig. 12). LINE has a longer weld pool compared to POINTS for all investigated traverse speeds. This is also seen in simulations, where the temperature drops slower with increasing distance to the interaction zone (see Fig. 9). The decrease of the length for POINTS at high speeds, which is not found for LINE, further indicates that less energy is absorbed when the oxide layer is not completely molten. Additionally the overall melt pool size and geometry, especially the shape of the re-solidifying front, varies significantly with the BP at constant process parameters. While the BP translates directly to the shape of the front of the melt pool, the rear end cannot be directly related to the BP due to heat conduction.
Fig. 12. Comparison of the weld pool length lW of LINE and POINTS at constant Pw = 1.3 kW.
Inspecting the cross-sections of the welds created by the different intensity profiles, a change in the form is visible as well. At larger melt pool sizes (see Fig. 14), the welds look similar for the shaped profiles as the initial intensity profile integrated in traverse direction and therefore the temperature field becomes less distinct due to heat conduction inside the melt pool. In contrast, with decreasing melt pool size (see Fig. 15) the original intensity profile becomes more visible. DEF creates a half circle like weld seam. The three other profiles produce rather flat sections. The weld cross-section of POINTS has a rectangular shape with quite sharp edges and the individual spots shining through at smallest melt pools (see Fig. 15b red arrows). The cross-section of RING comes with slightly 185
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Fig. 13. Fully and partly hidden melts pools below oxide layer at high traverse speeds (800 mm/s); a) RING Pw = 1.35 kW [vis5]; b) POINTS Pw = 1.31 kW [vis6]; c) DEF Pw = 1.71 kW [vis7]; d) LINE Pw = 1.35 kW [vis8].
Fig. 14. Etched cross-section at v = 50 mm/s a) RING Pw = 0.67 kW b) POINTS Pw = 0.66 kW c) DEF Pw = 0.86 kW d) LINE Pw = 0.68 kW.
Fig. 15. Etched cross-section at v = 800 mm/s a) RING Pw = 1.35 kW b) POINTS Pw = 1.31 kW c) DEF Pw = 1.71 kW d) LINE Pw = 1.35 kW.
Fig. 16. Unpolished bright-field top-view images of a) RING with Pw = 1.35 kW, v = 25 mm/s and a clearly visible grain structure b) POINTS with Pw = 1.31 kW, v = 50 mm/s and oxide layer leftovers c) LINE with Pw = 1.35 kW, v = 50 mm/s d) DEF Pw = 1.71 kW, v = 100 mm/s.
rounder edges and is slightly deeper at both sides (Fig. 15a). LINE has flat tapered flanks induced by the experimental BP. 3.5. Surface roughness The lowest SA was obtained with RING. The low melt pool dynamics and therefore, a high melt pool stability enables a surface finish where even grains become visible (see Fig. 16 a). POINTS suffers from its passage channels. The oxide layer in HCMW remains on top of the melt pool and reduces the surface quality (see Fig. 16b). LINE and DEF need higher traverse speeds to stay in HCMW. Consequently, the dynamics inside the melt pool are considerably higher and the surface of the weld is permeated with solidified waves (see Fig. 16c and d). The effect of the traverse speed and laser power on the resulting surface roughness is subsequently exemplarily discussed on the RING profile because of its large process window for HCMW (see Fig. 17). First
Fig. 17. Achieved surface roughness with RING depending on Pw and v.
of all, a high enough SPE is needed to break the oxide layer which is not the case for Pw = 0.67 kW at all traverse speeds and Pw = 1.35 kW and 186
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traverse speed higher than 200 mm/s. At the lowest applied laser power Pw = 0.67 kW an interesting phenomenon can be observed. Starting at 25 mm/s a melt pool in size equal to RING is created leading to a good surface quality most likely by reformation of the sand blasted surface by surface tension of the melt pool below the oxide layer. At 50 mm/s vibrations going throw the melt pool can be observed causing a local SA peek. Presumably the liquid is partly solidified in the central area of RING and remolten at the second interaction with the laser beam. That sequence of melting, solidification and remelting might be the reason for the observed behaviour. At higher speeds the centre for the RING stays solid and the surface roughness is getting better again. This explanation should be valid for the peak at Pw = 1.35 kW and v = 400 mm/s as well. As soon as the oxide layer is broken the surface roughness is positively affected. SA with the curvature of the weld seam having the highest impact down to 0.2 μm at v = 25 mm/s, 50 mm/s and Pw = 2.02 kW can be obtained. Those values are normally only achieved by laser polishing [43]. With increasing traverse speed the surface quality is decreasing again due to the increasing melt pool dynamics.
and length, can be influenced, where the front of the melt pool mimics the BP directly, while the re-solidification front cannot be directly deduced. Numerical simulations predict that the coupling efficiency of the laser power can be enhanced by beam shaping which is in agreement with experimental results showing different cross-sectional areas of the weld seam for different profiles at the same processing parameters. Simulations also show that although the mean temperature of the melt pool volume is comparable, the mean surface temperature and resulting evaporation pressure can differ between the BPs. Therefore, this work shows a huge potential of beam shaping for optimising laser welding processes. Beam shaping promises to improve this process due to larger attainable scanning speeds with large crosssection areas and melt pools with low process dynamics. To achieve the goal of beam profiles tailored to the individual requirements of each welding process further studies investigating more elaborate beam profiles and their influence on process dynamics and the quality of the weld are needed.
4. Conclusion
Acknowledgements
Beam shaping by diffractive optical elements enables the possibility to influence melt pool size, shape, and dynamics. Therefore, the process window for HCMW can be significantly increased by using shaped intensity profiles and surface qualities comparable to laser polishing can be obtained. In the presented study, a ring-profile had the largest process window and yielded the best surface qualities. In contrast to welding with a Gaussian spot, for beam shaping a high peak intensity does not necessarily lead to an instable melt pool, but the distribution itself plays the major role as could be shown by the different process windows for LINE and POINTS. The size of the melt pool, i. e. the depth
The authors want to thank the German Research Foundation (DFG) for funding the Collaborative Research Center 814 (CRC 814) – Additive Manufacturing, sub-project A5. The authors gratefully acknowledge the support provided by the Erlangen Graduate School in Advanced Optical Technologies. The work was performed with the support of the Ministry of Education and Science of the Russian Federation, decree N220, state Contract No. 14.Z50.31.0023 Appendix
Fig. 18. Weld dimensions of RING. Area AW and depth DW versus v for Pw .
Fig. 19. Weld dimensions of POINTS. AW and DW versus v for different Pw .
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Fig. 20. Weld dimensions of LINE. AW and DW versus v for different Pw .
Fig. 21. Weld dimensions of DEF. AW and DW versus v for different Pw .
Supplementary materials
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