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Process stability during fiber laser-arc hybrid welding of thick steel plates
Optics and Lasers in Engineering 102 (2018) 34–44
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Optics and Lasers in Engineering journal homepage: www.elsevier.com/locate/optlaseng
Process stability during fiber laser-arc hybrid welding of thick steel plates Ivan Bunaziv a,∗, Jan Frostevarg b, Odd M. Akselsen a,c, Alexander F.H. Kaplan b a
Norwegian University of Science and Technology, Department of Mechanical and Industrial Engineering, Richard Birkelands vei 2B, NO-7034 Trondheim, Norway Luleå University of Technology, Department of Engineering Sciences and Mathematics, SE-97187 Luleå, Sweden c SINTEF Materials and Chemistry, P.O. Box 4760 Sluppen, NO-7465 Trondheim, Norway b
a r t i c l e
i n f o
Keywords: Hybrid welding Process stability Fiber laser High speed imaging Thick section welding
a b s t r a c t Thick steel plates are frequently used in shipbuilding, pipelines and other related heavy industries, and are usually joined by arc welding. Deep penetration laser-arc hybrid welding could increase productivity but has not been thoroughly investigated, and is therefore usually limited to applications with medium thickness (5-15 mm) sections. A major concern is process stability, especially when using modern welding consumables such as metalcored wire and advanced welding equipment. High speed imaging allows direct observation of the process so that process behavior and phenomena can be studied. In this paper, 45 mm thick high strength steel was welded (butt joint double-sided) using the fiber laser-MAG hybrid process utilizing a metal-cored wire without pre-heating. Process stability was monitored under a wide range of welding parameters. It was found that the technique can be used successfully to weld thick sections with appropriate quality when the parameters are optimized. When comparing conventional pulsed and the more advanced cold metal transfer pulse (CMT+P) arc modes, it was found that both can provide high quality welds. CMT+P arc mode can provide more stable droplet transfer over a limited range of travel speeds. At higher travel speeds, an unstable metal transfer mechanism was observed. Comparing leading arc and trailing arc arrangements, the leading arc configuration can provide higher quality welds and more stable processing at longer inter-distances between the heat sources. © 2017 Elsevier Ltd. All rights reserved.
1. Introduction Laser-arc hybrid welding (LAHW) has a high potential in welding of thick steel plates for heavy industries. As explained by Ono et al. [1], this is due to its high penetration depths and production rates compared to arc welding. Moore et al. [2] investigated the possibilities of improving the mechanical properties of LAHW joints compared to autogenous laser beam welding (LBW), concentrating on the composition of the filler wire. Nowadays, modern flux- or metal-cored filler wires offer better weld metal toughness at low temperatures compared to solid wires according to Gook et al. [3]. However, the implementation of the process is still limited, partly due to the manifold parameters and complex laser-arc interactions. Steen [4] made the very first observations of LAHW with a camera, using a CO2 laser beam (10 μm wavelength) and a tungsten inert gas (TIG) arc. He identified an interaction between the laser and arc plasmas, enabling stabilization of the TIG process so that welding speeds could be increased. As described by Kristensen et al. [5], this augmentation gained a remarkable level of interest in the late 1990’s and was widely employed in the European shipbuilding heavy industry in the early 2000′s. Modern fiber lasers have a shorter wavelength (1 μm) and
∗
Corresponding author. E-mail address: [email protected] (I. Bunaziv).
https://doi.org/10.1016/j.optlaseng.2017.10.020 Received 4 July 2017; Received in revised form 13 October 2017; Accepted 26 October 2017 0143-8166/© 2017 Elsevier Ltd. All rights reserved.
can be guided by fiber optics. Another aspect of these lasers is that the amount of ionized plasma produced in the weld zone is almost negligible compared to CO2 lasers. During fiber laser-arc welding the arc plasma interacts with the vapor produced from the melt pool, which contains metal particles (especially near the keyhole) as investigated by Katayama et al. [6]. These particles have lower ionization temperatures than the weld shielding gas, as discussed by Hu et al. [28]. Cai et al. [25] and others have explained that this causes a distortion of the electric arc towards the keyhole, affecting drop detachment and droplet trajectory. High speed imaging (HSI), as explained by Eriksson et al. [7], is an excellent technique to achieve a better understanding of process behavior, offering direct observation of the process surface which can then be analyzed in detail, frame by frame. Frostevarg et al. [8] used HSI to compare the process differences between the conventional pulsed mode with the standard mode (continuous current), and the more recently developed cold metal transfer (CMT) arc modes for medium feed rates during LAHW. The droplet transfer for the standard mode changes from globular to spray, depending on settings, whilst the pulsed mode has globular transfer and CMT has short-arc with reversing wire motions to the melt pool. It was concluded that the CMT mode was the most stable and that any arc is stabilized by the presence of a joint gap, which also
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facilitates improved mixing throughout the depth of the weld, especially for the CMT arc mode, see Frostevarg et al. [9]. Productivity is affected by weld travel speed as well as the number of runs required. At sufficiently high travel speeds, both LBW [10] and especially MAG [11] processes became unstable, producing poor quality welds. An important setting in LAHW is the laser-arc inter-distance (DLA ), that for 10 μm (CO2 ) lasers was found by Abe et al. [12] to have a certain optimum value for stabilizing the electric arc. At higher DLA , Hayashi et al. [13] identified that the heat sources become more separated and therefore the process turns into a tandem process with little interaction between the sources, rather than hybrid with a shared melt pool. For 1 μm lasers, Fellman et al. [14] reported that the arc metal transfer mode and process stability is affected by the DLA and the torch arrangement (i.e. arc leading or trailing), where the DLA for optimum penetration was found to be further away from the keyhole for a leading torch. The reason is that at very short distances there is strong negative interaction due to impingement of molten droplets with the laser beam that partially reflects and reduces beam power into the keyhole opening. A fluctuating power delivery to the keyhole can cause porosity due to keyhole collapse. Porosity can also be generated by keyhole instabilities caused by overflow of melt, as when increasing arc power or filler wire rate, according to Fellman et al. [14]. On the other hand, in some cases porosity can be reduced by the same factors, as found by Katayama et al. [15]. Murakami et al. [16] used HSI to identify that a more favorable molten metal flow is a reason for the fact that a trailing arc arrangement provides higher quality welds in T-joints, giving less porosity and undercut. The melt flow was found to be affected by a strong pressure from the arc that pushes melt towards the front of the keyhole. Bunaziv et al. [17] have theoretically predicted that the flying droplets from the wire impinge on the keyhole opening more often when using a leading arc, also causing keyhole collapse. Frostevarg and Kaplan [18] used HSI to position the heat sources in order to tailor the melt pool flow in order to reduce undercutting, concluding that the optimum distance for producing a proper weld cap changes depending on arc size and power. The presence of an air gap has been found to have other benefits besides material mixing and arc stabilization. A gap also provides deeper penetration and stabilizes the keyhole, as explained by Piili et al. [19]. 1-4 mm air gaps were used by Hayashi et al. [13] when joining 22-25 mm thick plates using a 30 kW CO2 laser. They suggest that under the effect of the vaporization recoil pressure the keyhole forms immediately due to the presence of the molten metal in the groove when the arc is in a laser-leading configuration. Using this principle joints having air gaps exceeding the laser beam diameter can be welded successfully. The productivity potential of the LAHW process is especially high for thick section welding and has been the subject of a considerable amount of research over the last decade, e.g. for pipeline high strength low alloy (HSLA) steels using 1 μm lasers and multi-pass welding. Grünenwald et al. [20] joined 14 mm thick plates, achieving appropriate mechanical properties in joints created at 1-2 m/min travel speeds, which is substantially higher than those applicable in arc welding (0.5-0.6 m/min). In this case welding was performed by using LAHW in the first pass and MAG alone for the weld cap. Akselsen et al. [21] successfully joined 20 mm thick plates in two passes (double-sided), at 0.5-2 m/min travel speeds providing reasonable toughness (minimum 27 J at −40°C). More promising results were obtained by Pan et al. [22] where 12-24 mm thick plates were joined at 1-2 m/min travel speeds with good mechanical properties providing a minimum 100 J toughness at −40°C. The optimization of welding parameters in LAHW can be very challenging due to the manifold non-linearly related process parameters. Most studies of the LAHW process stability are limited to certain cases, with relatively narrow sets of weld parameters, or involve studying a particular phenomenon during processing of a specific metal, setup, laser optics and conditions. Therefore, there is no generally usable information on LAHW in thick carbon steel plates using 1 μm lasers and the CMT arc mode. This work aims to explain the process behavior for
Fig. 1. Experimental set-up: 1 – fiber laser equipment; 2 – MAG equipment; 3 – steel workpieces; 4 – heavy clamping system; 5 – illumination pulsed diode laser; and 6 – high speed imaging camera.
a wide range of conditions in order to explain generalizable process mechanisms so that findings may be used in future applications. The varied parameters discussed here include: DLA , weld travel speeds, process setup, joint groove preparations and gap widths. 2. Methodology 45 mm thick plates were joined with double-sided LAHW, using metal-cored wire, comparing the effects of CMT+P and pulsed arc modes in a variety of parameter settings in order to improve weld quality and productivity. The experimental methodology and analysis was as follows. 2.1. Equipment and materials The LAHW setup used for the experiments is shown in Fig. 1(a). During the experiments, a 15 kW (IPG Laser GmbH YLR-15000) ytterbium fiber laser was used, having: a fiber core diameter of 400 μm, beam parameter product of 10.3 mm•mrad, wavelength 1070 nm, continuous wave mode, 300 mm focal length optics, focus below the surface (−7 mm focal spot position giving a spot diameter of 800 μm and a Rayleigh length of ±4 mm). The laser was combined with electric arc welding equipment; TPS4000 VMT Remote from Fronius GmbH, with wire feeder unit VR7000 with a Robacta Drive unit (carrying out the back and forth motion of the wire tip, enabling the CMT process). To prevent high back reflections that can damage the laser source and optics, the laser was slightly tilted by 7°30ʹ. The arc torch had an angle of 60±2° with a stickout length of 15±1 mm. The shielding gas mixture of Argon with 18% CO2 flowed through the arc torch at a rate of 25 L/min. The welds were carried out using an articulated robot from Motoman. The arrangement of the equipment is shown in Fig. 1(b). The 45 mm thick HSLA steel plates (yield strength of 500 MPa) used in the experiments were plasma cut by 500 × 125 mm and after having their edges milled, sandblasted near the area of welding and heavily clamped prior to welding. A half-meter length of the plates was used to produce representative welds, so that the process stability and weld quality could be properly determined. The filler wire used was metalcored Kobelco Trustarc MX-A55T of 1.2 mm in diameter. For the HSI, the camera (Redlake) was used from the side, tilted by 55±5° from the horizontal. The region of interest had the resolution of 800 × 600 pixels with 10 bits pixel depth providing 3410 frames per second. The duration of filming was up to 1.2 s. In order to increase the quality of filming, the process light was filtered out by using a bandpass filter (808 nm) and the surface illuminated with a pulsed illumination laser (Cavitar), with 500 Watts peak power during a 3 μm shutter exposure time. 35
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Fig. 2. Design matrix of welding parameters.
current, voltage, and pulsing duration and frequency were set according to the synergetic lines of the arc weld machine. •
•
•
•
Fig. 3. Geometry of prepared grooves for welding: a) I-groove beveling with 5 mm face root width; b) Y-groove beveling with 5 mm face root width (not to scale). •
2.2. Design of variables and analysis
•
The welding parameters are shown within the design matrix in Fig. 2. In all the experiments, the laser output was set to 15 kW usually with a 3.5-4.5 mm DLA . Parameters that significantly affect process stability were changed between the experiments, including: travel speed vt (0.51.2 m/min), arc mode (CMT+P, pulsed), filler wire feed rate WFR (4-18 m/min), torch arrangement (leading or trailing arc), air gap between plates (0.2-1.0 mm), and joint edge preparation (I or Y groove). The beveled geometry of the joint edges is shown in Fig. 3. The WFR was chosen to properly fill the joint. The arc parameter characteristics for the
Set (i) experiments were treated as a reference for comparing differences between arc modes, joint edges and torch arrangement setup. They involve 0.4 mm air gap and 9.0 m/min WFR and vt of 0.8 m/min (this is high compared to pure MAG, where typically the travel speed is 0.5 m/min). Set (ii) was made with vt decreased to 0.6 m/min, air gap to 0.3 mm and WFR to 7 m/min. For the Y-groove the WFR was further increased by 1 m/min to adjust for the larger volume to be filled. Set (iii) investigated the effect of increasing air gaps to 0.6-0.8 mm, while having low and high levels of vt and WFR. Set (iv) expanded the parameter regime for the trailing arc configuration, to possibly improve mechanical properties due to increased filler wire mixing and supplying melt deeper into the plate by inward flow, as explained by Zhao et al. [23]. Variable parameters included: increased air gaps, vt and corresponding WFR. Set (v) included decreased vt to 0.5 m/min and WFR to 4 m/min, facilitating a pure CMT arc mode for 0.3 mm gap setup. Set (vi) included a few welds made with different DLA and slightly increased air gap.
For every combination of parameters in Fig. 2, one sample was made, but they are considered representative for production due to the length of the weld (500 mm in accordance with the ISO 15614-14:2013 standard). To evaluate the samples, the weld surfaces were visually inspected and cross sectional macrographs where made (polished and etched). HSI was applied to both sides during welding to verify identical behavior. In addition to other phenomena, droplet transfer stability, interactions between the arc and laser beam, keyhole stability and melt 36
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wire chopping and droplet explosion during forming and subsequent detachment, providing a more stable process. A possible reason for the suppression of wire chopping is that a deeper gouge is formed below the arc which is filled with a lower level of melt, resulting in a longer arc. Multiple-drop-per-pulse (MDPP, related to reduced pinch-effect) occurred more frequently in leading pulsed arc, sometimes resembling the spray transfer mode where detaching droplets form a molten metal channel, or stream, consisting of small elongated droplets, Fig. 6(c, d). MDPP in leading torch is usually formed regardless of travel speeds and WFR. Changes in drop detachment are related how the arc envelops the wire. One aspect of having a leading arc is that it pushes away the melt from the front of the gouge, exposing solid metal, Fig. 6(k). When using the CMT arc mode, the melt can flow back into the gouge during shortcircuiting, enabling surface tension drop transfer, see Fig. 6(l). The arc shape depends on weld conditions, which are the gouge shape, the melt pool shape and the physical and chemical state of the surrounding atmosphere. The arc shape in this case is distorted, decreasing the pinch-effect during arc pulses. MDPP also occurred in the more controlled CMT+P arc mode for both torch arrangements, Fig. 6(e, g, h). In the trailing torch setup, there was sometimes a delay in droplet detachment after an initial pulse when WFR was 7 m/min. After that a droplet micro-explosion might occur during pinch-effect in the subsequent pulse cycle. However, microexplosions during forming and detachment had no significant effect on overall process stability and surface quality. The leading torch arrangement had better process stability characterized by smoother and regular droplet detachment. Concerning droplet shape, it was observed that they were less spherical in the case of higher WFR due to the higher currents involved. An overall conclusion is that the stability of the metal transfer was improved when having a leading arc and lower vt , as regards instabilities such as wire chopping and droplet explosions during drop detachment. Sometimes during the short-circuiting phase, when using the CMT+P arc mode and increased vt , Fig. 7(g), the intended contact between the filler wire and the melt pool did not occur. This causes the melt at the wire tip to remain attached for several subsequent pulses, possibly causing droplet explosion (multiple small spatters) or wire chopping.
Fig. 4. HSI of laser-MAG hybrid welding for a) trailing torch arrangement and b) leading torch arrangement where: 1 – filler wire; 2 – arc plasma; 3 – molten droplet; 4 – liquid weld pool rim; 5 – laser keyhole; and 6 – air gap between plates.
flow in the process zone were all analyzed as they are critical for process stability. Example HSI frames are shown in Fig. 4(a, b) for both trailing and leading arcs, respectively. 3. Results and discussion 3.1. Weld results and process behavior of reference welds All reference experiments (set i), were performed at 0.8 m/min vt , with a 0.4 mm joint air gap and a WRF of 9 and 10 m/min for the I- and Y-groove gaps respectively (I-C1T, I-P1T, I-C1L, I-P1L, Y-C1T, Y-P1T, Y-C1L, and Y-P1L samples). A comparable set (ii) of experiments were conducted with reduced air gaps of 0.3 mm and 0.2 mm for I- and Ygroove respectively, vt was reduced (from 0.8 to 0.6 m/min) in order to maintain penetration depth, WFR (and corresponding arc settings) was adjusted to 7 m/min (I-C2T, I-P2T, I-C2L, I-P2L, Y-C2T, Y-P2T, Y-C2L, and Y-P2L samples). For all cases, the process behavior was identical for both sides of the weld. 3.1.1. Weld appearances Deposited weld appearances and macrosections are shown in Fig. 5. The appearance of the weld beads was very similar for both I- and Ygroove samples, due to almost identical parameters being used (only slightly more WFR was required to fill the gap in the Y-groove case). Therefore surface views are only shown for the I-grooves. Because of plate distortion during welding, the penetration depth can vary as slight gap variations occur during processing (in some cases also causing a little underfill). To prevent these issues from occurring, industrial implementation of thick sheet welding usually involves hydraulic clamping. The trailing torch provided slightly less spattering and higher quality weld seam appearance compared to the leading torch arrangement for both groove types. Spatter generation is probably linked to arc instabilities, causing changes in drop flight trajectories, possibly causing variations in weld seam appearance. An exception is the I-C1T sample (see Fig. 5(a)) that unintentionally had slightly shorter DLA (2.5-3 mm), resulting in occasional deposited droplet explosions and chopping of the wire instead of the arc forming at the wire tip. The pulsed arc mode provided better weld surface quality due to a lower incidence of arc and drop detachment instabilities such as wire chopping and droplet explosions.
3.1.3. Deposition trajectory of molten droplets The trajectory of the molten droplets can play an important role and should be considered when optimizing welding parameters. Material transfer from the wire frequently experiences deflection, or deviation, of droplet trajectory (see Fig. 6). For both trailing and leading arcs, the droplets tend to veer away from the gouge front due to changes in arc shape, but for different reasons. For the leading arc condition, the arc produces a gouge and pushes away the melt beneath so that the arc has a solid elongated surface from which to form, pushing the melt at the wire tip upwards so that the detached droplets fly for a longer distance (the pinch-effect is also reduced), Fig. 6(j). With a trailing arc, lower levels of ionization in the shielding gas (due to the presence of metal particles) shifts the arc central position closer to the keyhole. The sum of the arc force on the molten wire tip is then such as to deviate the droplet path from the weld front to the rear Fig. 6(i). The magnitude of drop trajectory deviation is strongly related to the welding parameters. The deviation is stronger when the arc power is higher and when the arc length is shorter (e.g. Fig. 6(e), I-C1T sample). Greater flight path deviation with higher WFR is also related to MDPP, where the droplets have smaller mass and more easily affected by the pressure from a stronger electron flow from the gouge front. Onedrop-per-pulse (ODPP) usually experiences less change in droplet path trajectory due to the larger mass of the droplets. Also, the flight path deviation during CMT+P was found to be reduced and the pinch-effect was less attenuated during arc pulsing, probably because of lower arc power levels. Notably, the deflection of droplet trajectory is not constant over time. For a leading torch the deflection of trajectory is visible for
3.1.2. Process behavior Fig. 6 contains illustrations of the process behavior recorded by HSI for all cases in set (i) and set (ii), showing arc sizes, melt pools, arc gouge shapes, melt flows, and droplet behavior. Droplet impingement is also observed and corresponding areas (as blue circle) are indicated. No significant differences were observed between the I- and Y-groove preparation samples. The pulsed arc provided larger arc sizes due to higher power levels, producing deeper gouges, especially in the leading torch arrangement. The leading torch arrangement slightly reduced 37
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Fig. 5. Macrosections and deposited weld seam appearances of I- and Y-groove samples.
more than half the flight time (up to 70%) compared to less than half time for the trailing torch (up to 40%). In the trailing torch arrangement, the increase of WFR which can disturb the keyhole dynamics due to excessive molten metal overflow is less of a problem, compared to the leading torch, due to a greater level of droplet trajectory deflection away from the keyhole. With a leading arc, the droplets were frequently landed close to the keyhole opening area, Fig. 6(c, d). The droplet impingement area in this case is close to the keyhole front and sometimes droplets strike the keyhole opening area. As a result, the keyhole is more prone to collapse and subsequent porosity may form. Note that although MDPP occurs more for a leading torch, the droplets are smaller than for ODPP. Therefore, each droplet has less effect on the keyhole collapse and porosity formation can be reduced
Fig. 5(e, f, m, n). Similar results related to porosity generation were obtained by Cao et al. [24]. Porosity was more severely pronounced in the Y-grooved specimens with both arc modes, see Fig. 5(h, d). This is probably due to droplets striking the keyhole opening and more vigorous internal melt flow. Y-groove specimens showed similar results to the Igroove samples as regards flight path deviations. However, it was noted that an increase of WFR by 1 m/min for the CMT+P arc (heat output increased by 27%) in Y-groove welding is characterized by double the amount of peak current pulses (from 6 to 12). 3.1.4. Wire chopping The wire chopping phenomenon observed during experiments is of particular interest because it has not been discussed or observed in pre38
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Fig. 6. a-h) LAHW process stability based on HSI for I-groove specimens where: 1 – arc plasma; 2 – droplet impingement area; 3 – molten droplet; 4 – melt flow; 5 – keyhole; 6 – air gap; 7 – variable arc length (where LA is arc length); and 8 – arc gouge (not covered by molten metal). i) droplet trajectory deviation for trailing and j) leading arc. k) melt is pushed from the gouge and l) during the short-circuiting phase the gouge is filled with melt.
vious scientific papers. However, evidence of the phenomenon can be found in earlier work done by Wahba et al. [26] who used solid wire. The detached wire can occur as spatter beside the weld seam, as shown in Fig. 5(aii ), or can possibly strike the keyhole opening, leading to a subsequent collapse. This could explain occasional macro porosities, see Fig. 5(ai ). The chopping mechanism could be related to metal-cored wire design, as solid wires are believed to be less prone to wire chopping. Wire chopping was observed to occur more frequently when using higher travel speeds, trailing torch arrangement, and CMT+P arc mode, indicated in Fig. 6. Figure 7 shows a sample with wire chopping occurring for CMT+P arc mode at 0.8 m/min vt . Streak imaging was used to analyze wire chopping phenomenon, see Fig. 7(d). Streak imaging uses a collection of stacked lines (one pixel wide) extracted from each frame of the video (vertical line in Fig. 7(c)). In the streak image, the perpendicular axis becomes time dependent, showing the time evolution of the chosen axis and allows for periodic behavior of the process to be analyzed and represented in a single picture. Frostevarg et al. [8] have applied this technique to examine causes of undercutting in different arc modes since it facilitates the analysis of complex phenomena. For wire chopping analysis, the streak line was selected at the wire at certain distance from the tip (along y axis direction) where wire chopping most frequently occurred. From this streak image, Fig. 7(d), it is possible to identify the relationship between wire chopping and a preceding short-circuiting phase that is of unusually short duration. The A-D positions in Fig. 7(d) cor-
responds to HSI frames in Fig. 7(e-h), showing the wire chopping event (A-B), and demonstrating that the arc then becomes larger until the process is stabilized again (C-D). In the CMT+P case presented in Fig. 7(d) the time required to re-stabilize the arc after the chopping event is about 27 ms, which correspond to the next short-circuiting phase (C). During re-stabilization, the arc length will be longer than normal (B), possibly causing disturbances on the weld pool and the resulting weld bead profile. Fig. 8 illustrates the wire chopping mechanism. Wire chopping occurs after an unusually short short-circuiting phase due to the wire not being sufficiently retracted to an appropriate distance from the molten pool. After such an event, the arc is very short. Before wire chopping, it seems that there is a buildup of heat in the wire with each pulse, see Fig. 8(a). The arc length gradually becomes longer with each pulse, but not reaching its normal length. Resistance of heated wire increases as its temperature approaches to the melting point. At some point after 3-5 pulses, the wire is prematurely melted up to a certain distance Fig. 8(b). At this distance, the wire is chopped off by Pinch-force during subsequent pulse (peak current) due to melting with increased resistance in this area, Fig. 8(c). Deposition trajectory of the chopped wire is unpredictable, depending on a combination of gravity and the sum of forces from the arc.
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Fig. 7. Deposited weld seam appearance (a, b) and HSI (55° inclined camera from surface) analysis of LAHW (Y-C1T sample) through streak image (d) technique of each selected line showing unstable processing (e) and stable (h) processing.
Fig. 8. Wire chopping mechanism.
chopping was observed when the arc length was very short. An increase of WFR by 1 m/min for the I-groove preparation slightly improved the stability of the process, compared to set (i) experiments. Spattering was significantly reduced due to a reduction in wire chopping.
3.2. Process behavior when changing process settings 3.2.1. Increase of joint air gap Increasing the air gap increases the requirement for filler material, as insufficient deposition rates can produce slight underfill, e.g. Fig. 9(a) (sample I-P3T). An increase of air gap from 0.4 to 0.6 mm for high (sample I-C5T; 9 m/min WFR and 0.8 m/min vt ) and low (sample I-C6T; 7 m/min WFR and 0.6 m/min vt ) level parameters for the trailing CMT+P arc mode, did not significantly improve the stability of the process. It is likely that the increase in gap width (0.2 mm) was too small to have any noticeable effect. Increasing the air gap to 0.8 mm (samples I-C3T, I-P3T, Y-C3T, and Y-P3T), slightly increased process stability, providing MDPP metal transfer for both trailing and leading arc modes (with 10 m/min WFR and 0.8 m/min vt ). For CMT+P in the I-groove case only, wire
3.2.2. Process stability after further increase in travel speed, wire feed rate and air gap Using CMT+P, I-groove joint preparation and increasing vt from 0.8 to 0.95 m/min (sample I-C7T), the process became very unstable with frequent wire chopping (of different lengths), a significant amount of spatter and some porosity. For both I- and Y-groove preparation, using pulsed arc modes and increasing the WFR to 13 m/min (samples I-P8T and Y-P4T), resulted in MDPP-to-spray droplet transfer and significantly less wire chopping with some underfill, as shown in Fig. 9(b) (sample 40
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free-flight mode, between short-circuits in CMT+P as shown in Fig. 11(ac) (sample I-C4T). Constant short-circuiting (direct contact with molten pool) is provided without chopping and the weaker arc pulses ensure calm and stable melt flow around the keyhole. Similar observations of LAHW with CMT were made by Frostevarg et al. [8]. As a result, no spatter was generated (see Fig. 11(d)). Streak imaging (Fig. 11(e)) reveals that duration of short-circuiting is stable, providing appropriate melting of the electrode wire (see Fig. 11(b, c)). This indicates that at lower travel speeds the rooting effect is well established. Despite the excellent process stability, the penetration depth was significantly reduced to 15-16 mm. From HSI it was identified that melt flow velocity were significantly slower compared to experiments with higher WFR and travel speed, indicating that the arc melt pool flow affects and reduces the penetration capability of the laser keyhole. 3.2.4. Stability effects of varying the laser-arc inter-distance In most welds, the DLA was 3.5-4.5 mm and no clear stability trends could be isolated or observed within this interval. To better observe the effect of different DLA upon process stability, the distance was increased to 8 mm for the leading arc and decreased to 2 mm for the arc trailing case. When using a leading pulsed arc and increasing the DLA from 4 mm to 8 mm (Y-P8L and Y-P9L samples), vt at 0.8 m/min and 0.6 mm air gap, the process became more stable for both 9 and 13 m/min WFR. There was a very low arc length variation, less spatter, and the droplets from the wire no longer impinged the keyhole opening area, see Fig. 12(a). In addition, no porosity formation was observed, (this porosity is probably linked to keyhole collapses, and none were observed in this case). Observations showed a low degree of arc and laser interaction and droplet transfer mode is still in MDPP-to-spray transfer mode, Fig. 12(ai, ii ), probably also related to increased WFR. As a result, in this configuration allows utilizing higher WFR there is a low probability of process instability. On rare occasions, due to arc pressures on the melt, the height of the melt around the keyhole was observed to increase (with waves in the melt pool), but this did not seem to disturb the process. When having a trailing pulsed arc and decreasing the DLA to 2 mm (samples Y-P6T and Y-P7T), with a 13 m/min WFR and using high vt at 0.8 m/min with a 0.6 mm air gap, the process was observed to be stable. Even though the arc was closer to the keyhole, droplet trajectory deviation away from the melt pool front was increased, so that droplets did not land in the vicinity of the keyhole Fig. 12(b). Also, the keyhole was not observed to be covered by melt. Therefore, porosity was not found in the samples. Spatter levels were low, Fig. 12(biv ), and wire chopping was not observed. At 9 m/min WFR, the droplet transfer mode alternated between globular with substantial droplet trajectory deviations as shown in Fig. 12(bi, ii ) (Y-P6T sample) and a MDPP resembling molten stream (see Fig. 12(biii )).
Fig. 9. The effect of WFR, travel speed and air gap on quality of welds in pulsed trailing arc.
I-P8T). Very low amounts of spatter were also produced (Fig. 9). Increasing the WFR to 17 m/min and vt to 1.0 m/min in a Y-groove setup (Y-P5T sample) eliminated wire chopping while also reducing porosity and spattering. Frostevarg et al. [8] has shown that at higher weld speeds, higher gap widths have an arc stabilizing effect. These observations verify that these trends are also valid for thick plate welding. Further increasing vt to 1.2 m/min and air gap to 1.0 mm (I-P9T sample), the WFR was also increased to 18 m/min to fill the gap properly. Due to increased vt at 1 mm gap, there is a significant underfill with large areas of premature solidification as shown in the Fig. 10(a-c) HSI sequences. The process was found to be relatively stable and the arc showed only very slight variations (Fig. 10(c)). However, due to the gap width, the laser keyhole opening is vertically lower than the top surface of the sheet Fig. 10(a). Therefore there is the melt pool could occasionally overflow into the keyhole (due to arc pressure), which could cause porosity. At higher vt , the reduction in wire chopping is related to increased WFR with MDPP improving the rooting effect. Process stabilization by wider air gaps was not being observed for these samples, even though such effects have previously been observed by Frostevarg et al. [8]. This is probably due to the increased amount of filler material which demands high arc powers. A large arc will produce a much wider gouge than the gap width, and will therefore not be constrained by it.
4. Generalized results From this analysis it can be concluded that high vt can result in unstable welding processing in terms of molten metal droplet transfer, especially for the CMT+P arc mode. Instabilities in filler wire deposition, especially wire chopping, during high vt are related to an inability to establish the so-called rooting effect [4] for the arc, and also to too short
3.2.3. Decreased travel speed and WFR Decreasing the WFR to 4 m/min and vt to 0.5 m/min in the presence of a narrow 0.3 mm gap (I-C4T and Y-C4T samples) resulted in an absence of intermittent pulses and drop detachment with subsequent
Fig. 10. a-c) HSI frames of high travel speed (1.2 m/min) for trailing pulsed arc mode (18 m/min WFR) and d) weld bead appearance. MF is melt flow and PS is premature solidification.
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Fig. 11. HSI at low travel speed (0.5 m/min) and trailing arc with 4 m/min WFR using the CMT mode.
Fig. 12. HSI of different process distances effect on process stability with different torch arrangements (0.8 m/min vt , 0.6 mm gap), where: A – MDPP transfer mode; B – keyhole boundaries; C – droplets impingement area; and D – flight path trajectory deviation.
5. Conclusions
an arc length. This makes the application of metal-cored wire more difficult compared to solid wires. However, instabilities (like wire chopping and problems with droplet detachment) can be reduced by increasing the WFR since stronger arcs have a better rooting effect. However, it should be kept in mind that stronger arcs have a more significant effect on droplet trajectories. In addition, an increase in WFR (in order to fill wider gaps) does not ensure more stable processing since a high WFR can disturb the keyhole, especially with a leading arc and shorter DLA (for both torch arrangements). This verifies the findings of Fellman et al. [14]. In addition, leading arc arrangement is less susceptible to wire chopping due to an elongation of the arc, which is related to a deeper arc gouge that is only partially filled with molten metal. Increasing WFR was also found to change droplet detachment towards MDPP due to the presence of higher temperatures. Fig. 13 is a matrix flow chart (MFC), first adopted by Karlsson et al. [27]. It shows generalized experimental trends. Note that identified trends only are shown for some variable values. Only a few experiments had a proper ODPP, since the arc gets disturbed by the gouge form and metal vapor from laser keyhole. Groove preparation type had only a slight effect on processing therefore it can be concluded that LAHW is fairly tolerant to gap variations and applications of different groove types.
LAHW of thick steel plates using metal-cored filler wire is a very efficient and promising fusion joining process, providing welds of acceptable quality within a certain optimized range of parameters. From the observed process behavior and resulting welds, the following conclusions can be drawn:
•
•
•
•
•
42
During LAHW, interactions between the heat sources and the melt flow affects the arc process, creating conditions for ODPP to MDPP or even the spray transfer mode, depending on process settings. For a leading arc, short laser-arc inter-distance is not recommended as there is an increased risk of keyhole collapse causing porosity, due to droplet impingement on the keyhole opening. Short inter-distances can be allowed for trailing arcs without causing keyhole collapse because the arc pressure on the melt causes droplets to fly away from the gouge front. The CMT+P arc mode is not recommended for travel speeds higher than 0.8 m/min due to the occurrence of wire chopping (of the metal-cored wire) which could be deleterious to weld quality. The type of groove preparation was observed to have little effect on process stability.
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Fig. 13. MFC of process stability based on performed experiments for different groove preparations and laser-arc arrangement setup. Notations of effects: 0 is a non-occurring event, (+) is an infrequent or weakly occurring event and (++) is a frequently or strongly occurring event.
•
An air gap between plates only has a significant influence on process stability when the gap is wider than 0.6 mm.
[10] Fabbro R. Melt pool and keyhole behavior analysis for deep penetration laser welding. J Phys D: Appl Phys 2010;43(44). [11] Sugino T, Tsukamoto S, Nakamura T, Arakane G. Fundamental study on welding phenomena in pulsed laser-GMA hybrid welding. In: In proc. of 24th international congress on applications of lasers and electro-optics (ICALEO), 31th October to 3rd November. Laser Institute of America; 2005. p. 108–16. [12] Abe N, Kunugita Y, Hayashi M, Tsuchitani Y. Dynamic observation of high speed laser-arc combination welding of thick steel plates. Trans JWRI 1997;26(2):7–11. [13] Hayashi T, Katayama S, Abe N, Omori A. High-power CO2 laser-MIG hybrid welding for increased gap tolerance. Hybrid weldability of thick steel plates with a square groove. Welding Int 2003;18(9):692–701. [14] Fellman A, Salminen A. Study of the phenomena of fiber laser-MAG hybrid welding. In: In proc. of 26th international congress on applications of lasers and electro-optics (ICALEO), October 29th to November 1st.. Laser Institute of America; 2007. p. 871–80. [15] Katayama S, Naito Y, Uchiumi S, Mizutani M. Physical phenomena and porosity prevention mechanism in laser-arc hybrid welding. Trans JWRI 2006;35(1):13–18. [16] Murakami T, Shin M-H, Nakata K. Effect of welding direction on weld bead formation in high power fiber laser and MAG arc hybrid welding. Trans JWRI 2010;39(2):175–7. [17] Bunaziv I, Akselsen OM, Ren X, Salminen A. Hybrid welding possibilities of thick sections for arctic applications. Phys Proc 2015;78:74–83. [18] Frostevarg J, Kaplan AFH. Undercut suppression in laser-arc hybrid welding by melt pool tailoring. J Laser Appl 2014;26:031501. doi:10.2351/1.4872062. [19] Piili H, Salminen A, Harkko P, Lehtinen J. Study of phenomenon of fibre-laser-Mig/Mag-hybrid-welding. In: In proc. of 27th international congress on applications of lasers and electro-optics (ICALEO), October 20-23rd. Laser Institute of America; 2008. p. 506–15. [20] Grünenwald S, Seefeld T, Vollertsen F, Kocak M. Solutions for joining pipe steels using laser-GMA-hybrid welding processes. Phys Proc 2010;5(Part B):77–87. [21] Akselsen OM, Wiklund G, Østby E, Sørgjerd A, Kaplan A. A first assessment of laser hybrid welding of 420 MPa steel for offshore structure applications. In: proceedings of the 14th nordic laser materials processing conference (NOLAMP), August 26-28th. Luleå University of Technology; 2013. p. 171–82. [22] Pan Q, Mizutani M, Kawahito Y, Katayama S. High power disk laser-metal active gas arc hybrid welding of thick high tensile strength steel plates. J Laser Appl 2016;28:012004. [23] Zhao L, Sugino T, Arakane G, Tsukamoto S. Influence of welding parameters on distribution of wire feeding elements in CO2 laser GMA hybrid welding. Sci Technol Weld Join 2009;14(5):457–67.
Acknowledgments The authors wish to thank the Research Council of Norway for funding through the Petromaks 2 Programme, Contract No. 228513/E30, as well as EU-RFCS project OptoSteel. The financial support from ENI, Statoil, Lundin, Total, JFE Steel Corporation, Posco, Kobe Steel, SSAB, Bredero Shaw, Borealis, Trelleborg, Nexans, Aker Solutions, FMC Kongsberg Subsea, Marine Aluminium, Hydro and Sapa are also acknowledged. References [1] Ono M, Shinbo Y, Yoshitake A, Ohmura M. Development of laser-arc hybrid welding. NKK Tech Rev 2002(No. 86):8–12. [2] Moore PL, Howse DS, Wallach ER. Microstructures and properties of laser/arc hybrid welds and autogenous laser welds in pipeline steels. Sci Technol Weld Join 2004;9(4):314–22. [3] Gook S, Gumenyuk A, Rethmeier M. Hybrid laser arc welding of X80 and X120 steel grade. Sci Technol Weld Join 2014;19(1):15–24. [4] Steen WM. Arc augmented laser processing of materials. J Appl Phys 1980;51(11):5636–41. [5] Kristensen JK, Webster S, Petring D. Hybrid laser welding of thick section steels ˗ The HYBLAS project. In: Proc. of 12th NOLAMP 2012 Conference, Copenhagen, Denmark; 2009 (August 2009). [6] Katayama S, Kawahito Y, Mizutani M. Elucidation of laser welding phenomena and factors affecting weld penetration and welding defects. Phys Proc 2010;5:9–17. [7] Eriksson I, Gren P, Kaplan AFH. New high-speed photography technique for observation of fluid flow in laser welding. Opt Eng 2010;49(10):100503. doi:10.1117/1.3502567. [8] Frostevarg J, Kaplan AFH, Lamas J. Comparison of CMT with other arc modes for laser-arc hybrid welding of 7 mm steel. Welding World 2014;58(5):649–60. [9] Frostevarg J. Comparison of three different arc modes for laser-arc hybrid welding steel. J Laser Appl 2016;28(2):15–35. 43
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[24] Cao X, Wanjara P, Huang J, Munro C, Nolting A. Hybrid fiber laser-arc welding of thick section high strength low alloy steel. Mater Des 2011;32(6):3399–413. [25] Cai C, Feng J, Li L, Chen Y. Influence of laser on the droplet behavior in short-circuiting, globular and spray modes of hybrid fiber laser-MIG welding. Opt Laser Technol 2016;83:108–18. [26] Wahba M, Mizutani M, Katayama S. Hybrid welding with fiber laser and CO2 gas shielded arc. J Mater Process Technol 2015;221:146–53.
[27] Karlsson J, Kaplan AFH. Analysis of a fibre laser welding case study, utilising a matrix flow chart. Appl Surface Sci 2011;257(9):4113–22. [28] Hu B, den Ouden G. Laser induced stabilisation of the welding arc. Sci Technol Weld Join 2005;10(1):76–81.
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Contents lists available at ScienceDirect
Optics and Lasers in Engineering journal homepage: www.elsevier.com/locate/optlaseng
Process stability during fiber laser-arc hybrid welding of thick steel plates Ivan Bunaziv a,∗, Jan Frostevarg b, Odd M. Akselsen a,c, Alexander F.H. Kaplan b a
Norwegian University of Science and Technology, Department of Mechanical and Industrial Engineering, Richard Birkelands vei 2B, NO-7034 Trondheim, Norway Luleå University of Technology, Department of Engineering Sciences and Mathematics, SE-97187 Luleå, Sweden c SINTEF Materials and Chemistry, P.O. Box 4760 Sluppen, NO-7465 Trondheim, Norway b
a r t i c l e
i n f o
Keywords: Hybrid welding Process stability Fiber laser High speed imaging Thick section welding
a b s t r a c t Thick steel plates are frequently used in shipbuilding, pipelines and other related heavy industries, and are usually joined by arc welding. Deep penetration laser-arc hybrid welding could increase productivity but has not been thoroughly investigated, and is therefore usually limited to applications with medium thickness (5-15 mm) sections. A major concern is process stability, especially when using modern welding consumables such as metalcored wire and advanced welding equipment. High speed imaging allows direct observation of the process so that process behavior and phenomena can be studied. In this paper, 45 mm thick high strength steel was welded (butt joint double-sided) using the fiber laser-MAG hybrid process utilizing a metal-cored wire without pre-heating. Process stability was monitored under a wide range of welding parameters. It was found that the technique can be used successfully to weld thick sections with appropriate quality when the parameters are optimized. When comparing conventional pulsed and the more advanced cold metal transfer pulse (CMT+P) arc modes, it was found that both can provide high quality welds. CMT+P arc mode can provide more stable droplet transfer over a limited range of travel speeds. At higher travel speeds, an unstable metal transfer mechanism was observed. Comparing leading arc and trailing arc arrangements, the leading arc configuration can provide higher quality welds and more stable processing at longer inter-distances between the heat sources. © 2017 Elsevier Ltd. All rights reserved.
1. Introduction Laser-arc hybrid welding (LAHW) has a high potential in welding of thick steel plates for heavy industries. As explained by Ono et al. [1], this is due to its high penetration depths and production rates compared to arc welding. Moore et al. [2] investigated the possibilities of improving the mechanical properties of LAHW joints compared to autogenous laser beam welding (LBW), concentrating on the composition of the filler wire. Nowadays, modern flux- or metal-cored filler wires offer better weld metal toughness at low temperatures compared to solid wires according to Gook et al. [3]. However, the implementation of the process is still limited, partly due to the manifold parameters and complex laser-arc interactions. Steen [4] made the very first observations of LAHW with a camera, using a CO2 laser beam (10 μm wavelength) and a tungsten inert gas (TIG) arc. He identified an interaction between the laser and arc plasmas, enabling stabilization of the TIG process so that welding speeds could be increased. As described by Kristensen et al. [5], this augmentation gained a remarkable level of interest in the late 1990’s and was widely employed in the European shipbuilding heavy industry in the early 2000′s. Modern fiber lasers have a shorter wavelength (1 μm) and
∗
Corresponding author. E-mail address: [email protected] (I. Bunaziv).
https://doi.org/10.1016/j.optlaseng.2017.10.020 Received 4 July 2017; Received in revised form 13 October 2017; Accepted 26 October 2017 0143-8166/© 2017 Elsevier Ltd. All rights reserved.
can be guided by fiber optics. Another aspect of these lasers is that the amount of ionized plasma produced in the weld zone is almost negligible compared to CO2 lasers. During fiber laser-arc welding the arc plasma interacts with the vapor produced from the melt pool, which contains metal particles (especially near the keyhole) as investigated by Katayama et al. [6]. These particles have lower ionization temperatures than the weld shielding gas, as discussed by Hu et al. [28]. Cai et al. [25] and others have explained that this causes a distortion of the electric arc towards the keyhole, affecting drop detachment and droplet trajectory. High speed imaging (HSI), as explained by Eriksson et al. [7], is an excellent technique to achieve a better understanding of process behavior, offering direct observation of the process surface which can then be analyzed in detail, frame by frame. Frostevarg et al. [8] used HSI to compare the process differences between the conventional pulsed mode with the standard mode (continuous current), and the more recently developed cold metal transfer (CMT) arc modes for medium feed rates during LAHW. The droplet transfer for the standard mode changes from globular to spray, depending on settings, whilst the pulsed mode has globular transfer and CMT has short-arc with reversing wire motions to the melt pool. It was concluded that the CMT mode was the most stable and that any arc is stabilized by the presence of a joint gap, which also
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facilitates improved mixing throughout the depth of the weld, especially for the CMT arc mode, see Frostevarg et al. [9]. Productivity is affected by weld travel speed as well as the number of runs required. At sufficiently high travel speeds, both LBW [10] and especially MAG [11] processes became unstable, producing poor quality welds. An important setting in LAHW is the laser-arc inter-distance (DLA ), that for 10 μm (CO2 ) lasers was found by Abe et al. [12] to have a certain optimum value for stabilizing the electric arc. At higher DLA , Hayashi et al. [13] identified that the heat sources become more separated and therefore the process turns into a tandem process with little interaction between the sources, rather than hybrid with a shared melt pool. For 1 μm lasers, Fellman et al. [14] reported that the arc metal transfer mode and process stability is affected by the DLA and the torch arrangement (i.e. arc leading or trailing), where the DLA for optimum penetration was found to be further away from the keyhole for a leading torch. The reason is that at very short distances there is strong negative interaction due to impingement of molten droplets with the laser beam that partially reflects and reduces beam power into the keyhole opening. A fluctuating power delivery to the keyhole can cause porosity due to keyhole collapse. Porosity can also be generated by keyhole instabilities caused by overflow of melt, as when increasing arc power or filler wire rate, according to Fellman et al. [14]. On the other hand, in some cases porosity can be reduced by the same factors, as found by Katayama et al. [15]. Murakami et al. [16] used HSI to identify that a more favorable molten metal flow is a reason for the fact that a trailing arc arrangement provides higher quality welds in T-joints, giving less porosity and undercut. The melt flow was found to be affected by a strong pressure from the arc that pushes melt towards the front of the keyhole. Bunaziv et al. [17] have theoretically predicted that the flying droplets from the wire impinge on the keyhole opening more often when using a leading arc, also causing keyhole collapse. Frostevarg and Kaplan [18] used HSI to position the heat sources in order to tailor the melt pool flow in order to reduce undercutting, concluding that the optimum distance for producing a proper weld cap changes depending on arc size and power. The presence of an air gap has been found to have other benefits besides material mixing and arc stabilization. A gap also provides deeper penetration and stabilizes the keyhole, as explained by Piili et al. [19]. 1-4 mm air gaps were used by Hayashi et al. [13] when joining 22-25 mm thick plates using a 30 kW CO2 laser. They suggest that under the effect of the vaporization recoil pressure the keyhole forms immediately due to the presence of the molten metal in the groove when the arc is in a laser-leading configuration. Using this principle joints having air gaps exceeding the laser beam diameter can be welded successfully. The productivity potential of the LAHW process is especially high for thick section welding and has been the subject of a considerable amount of research over the last decade, e.g. for pipeline high strength low alloy (HSLA) steels using 1 μm lasers and multi-pass welding. Grünenwald et al. [20] joined 14 mm thick plates, achieving appropriate mechanical properties in joints created at 1-2 m/min travel speeds, which is substantially higher than those applicable in arc welding (0.5-0.6 m/min). In this case welding was performed by using LAHW in the first pass and MAG alone for the weld cap. Akselsen et al. [21] successfully joined 20 mm thick plates in two passes (double-sided), at 0.5-2 m/min travel speeds providing reasonable toughness (minimum 27 J at −40°C). More promising results were obtained by Pan et al. [22] where 12-24 mm thick plates were joined at 1-2 m/min travel speeds with good mechanical properties providing a minimum 100 J toughness at −40°C. The optimization of welding parameters in LAHW can be very challenging due to the manifold non-linearly related process parameters. Most studies of the LAHW process stability are limited to certain cases, with relatively narrow sets of weld parameters, or involve studying a particular phenomenon during processing of a specific metal, setup, laser optics and conditions. Therefore, there is no generally usable information on LAHW in thick carbon steel plates using 1 μm lasers and the CMT arc mode. This work aims to explain the process behavior for
Fig. 1. Experimental set-up: 1 – fiber laser equipment; 2 – MAG equipment; 3 – steel workpieces; 4 – heavy clamping system; 5 – illumination pulsed diode laser; and 6 – high speed imaging camera.
a wide range of conditions in order to explain generalizable process mechanisms so that findings may be used in future applications. The varied parameters discussed here include: DLA , weld travel speeds, process setup, joint groove preparations and gap widths. 2. Methodology 45 mm thick plates were joined with double-sided LAHW, using metal-cored wire, comparing the effects of CMT+P and pulsed arc modes in a variety of parameter settings in order to improve weld quality and productivity. The experimental methodology and analysis was as follows. 2.1. Equipment and materials The LAHW setup used for the experiments is shown in Fig. 1(a). During the experiments, a 15 kW (IPG Laser GmbH YLR-15000) ytterbium fiber laser was used, having: a fiber core diameter of 400 μm, beam parameter product of 10.3 mm•mrad, wavelength 1070 nm, continuous wave mode, 300 mm focal length optics, focus below the surface (−7 mm focal spot position giving a spot diameter of 800 μm and a Rayleigh length of ±4 mm). The laser was combined with electric arc welding equipment; TPS4000 VMT Remote from Fronius GmbH, with wire feeder unit VR7000 with a Robacta Drive unit (carrying out the back and forth motion of the wire tip, enabling the CMT process). To prevent high back reflections that can damage the laser source and optics, the laser was slightly tilted by 7°30ʹ. The arc torch had an angle of 60±2° with a stickout length of 15±1 mm. The shielding gas mixture of Argon with 18% CO2 flowed through the arc torch at a rate of 25 L/min. The welds were carried out using an articulated robot from Motoman. The arrangement of the equipment is shown in Fig. 1(b). The 45 mm thick HSLA steel plates (yield strength of 500 MPa) used in the experiments were plasma cut by 500 × 125 mm and after having their edges milled, sandblasted near the area of welding and heavily clamped prior to welding. A half-meter length of the plates was used to produce representative welds, so that the process stability and weld quality could be properly determined. The filler wire used was metalcored Kobelco Trustarc MX-A55T of 1.2 mm in diameter. For the HSI, the camera (Redlake) was used from the side, tilted by 55±5° from the horizontal. The region of interest had the resolution of 800 × 600 pixels with 10 bits pixel depth providing 3410 frames per second. The duration of filming was up to 1.2 s. In order to increase the quality of filming, the process light was filtered out by using a bandpass filter (808 nm) and the surface illuminated with a pulsed illumination laser (Cavitar), with 500 Watts peak power during a 3 μm shutter exposure time. 35
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Fig. 2. Design matrix of welding parameters.
current, voltage, and pulsing duration and frequency were set according to the synergetic lines of the arc weld machine. •
•
•
•
Fig. 3. Geometry of prepared grooves for welding: a) I-groove beveling with 5 mm face root width; b) Y-groove beveling with 5 mm face root width (not to scale). •
2.2. Design of variables and analysis
•
The welding parameters are shown within the design matrix in Fig. 2. In all the experiments, the laser output was set to 15 kW usually with a 3.5-4.5 mm DLA . Parameters that significantly affect process stability were changed between the experiments, including: travel speed vt (0.51.2 m/min), arc mode (CMT+P, pulsed), filler wire feed rate WFR (4-18 m/min), torch arrangement (leading or trailing arc), air gap between plates (0.2-1.0 mm), and joint edge preparation (I or Y groove). The beveled geometry of the joint edges is shown in Fig. 3. The WFR was chosen to properly fill the joint. The arc parameter characteristics for the
Set (i) experiments were treated as a reference for comparing differences between arc modes, joint edges and torch arrangement setup. They involve 0.4 mm air gap and 9.0 m/min WFR and vt of 0.8 m/min (this is high compared to pure MAG, where typically the travel speed is 0.5 m/min). Set (ii) was made with vt decreased to 0.6 m/min, air gap to 0.3 mm and WFR to 7 m/min. For the Y-groove the WFR was further increased by 1 m/min to adjust for the larger volume to be filled. Set (iii) investigated the effect of increasing air gaps to 0.6-0.8 mm, while having low and high levels of vt and WFR. Set (iv) expanded the parameter regime for the trailing arc configuration, to possibly improve mechanical properties due to increased filler wire mixing and supplying melt deeper into the plate by inward flow, as explained by Zhao et al. [23]. Variable parameters included: increased air gaps, vt and corresponding WFR. Set (v) included decreased vt to 0.5 m/min and WFR to 4 m/min, facilitating a pure CMT arc mode for 0.3 mm gap setup. Set (vi) included a few welds made with different DLA and slightly increased air gap.
For every combination of parameters in Fig. 2, one sample was made, but they are considered representative for production due to the length of the weld (500 mm in accordance with the ISO 15614-14:2013 standard). To evaluate the samples, the weld surfaces were visually inspected and cross sectional macrographs where made (polished and etched). HSI was applied to both sides during welding to verify identical behavior. In addition to other phenomena, droplet transfer stability, interactions between the arc and laser beam, keyhole stability and melt 36
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wire chopping and droplet explosion during forming and subsequent detachment, providing a more stable process. A possible reason for the suppression of wire chopping is that a deeper gouge is formed below the arc which is filled with a lower level of melt, resulting in a longer arc. Multiple-drop-per-pulse (MDPP, related to reduced pinch-effect) occurred more frequently in leading pulsed arc, sometimes resembling the spray transfer mode where detaching droplets form a molten metal channel, or stream, consisting of small elongated droplets, Fig. 6(c, d). MDPP in leading torch is usually formed regardless of travel speeds and WFR. Changes in drop detachment are related how the arc envelops the wire. One aspect of having a leading arc is that it pushes away the melt from the front of the gouge, exposing solid metal, Fig. 6(k). When using the CMT arc mode, the melt can flow back into the gouge during shortcircuiting, enabling surface tension drop transfer, see Fig. 6(l). The arc shape depends on weld conditions, which are the gouge shape, the melt pool shape and the physical and chemical state of the surrounding atmosphere. The arc shape in this case is distorted, decreasing the pinch-effect during arc pulses. MDPP also occurred in the more controlled CMT+P arc mode for both torch arrangements, Fig. 6(e, g, h). In the trailing torch setup, there was sometimes a delay in droplet detachment after an initial pulse when WFR was 7 m/min. After that a droplet micro-explosion might occur during pinch-effect in the subsequent pulse cycle. However, microexplosions during forming and detachment had no significant effect on overall process stability and surface quality. The leading torch arrangement had better process stability characterized by smoother and regular droplet detachment. Concerning droplet shape, it was observed that they were less spherical in the case of higher WFR due to the higher currents involved. An overall conclusion is that the stability of the metal transfer was improved when having a leading arc and lower vt , as regards instabilities such as wire chopping and droplet explosions during drop detachment. Sometimes during the short-circuiting phase, when using the CMT+P arc mode and increased vt , Fig. 7(g), the intended contact between the filler wire and the melt pool did not occur. This causes the melt at the wire tip to remain attached for several subsequent pulses, possibly causing droplet explosion (multiple small spatters) or wire chopping.
Fig. 4. HSI of laser-MAG hybrid welding for a) trailing torch arrangement and b) leading torch arrangement where: 1 – filler wire; 2 – arc plasma; 3 – molten droplet; 4 – liquid weld pool rim; 5 – laser keyhole; and 6 – air gap between plates.
flow in the process zone were all analyzed as they are critical for process stability. Example HSI frames are shown in Fig. 4(a, b) for both trailing and leading arcs, respectively. 3. Results and discussion 3.1. Weld results and process behavior of reference welds All reference experiments (set i), were performed at 0.8 m/min vt , with a 0.4 mm joint air gap and a WRF of 9 and 10 m/min for the I- and Y-groove gaps respectively (I-C1T, I-P1T, I-C1L, I-P1L, Y-C1T, Y-P1T, Y-C1L, and Y-P1L samples). A comparable set (ii) of experiments were conducted with reduced air gaps of 0.3 mm and 0.2 mm for I- and Ygroove respectively, vt was reduced (from 0.8 to 0.6 m/min) in order to maintain penetration depth, WFR (and corresponding arc settings) was adjusted to 7 m/min (I-C2T, I-P2T, I-C2L, I-P2L, Y-C2T, Y-P2T, Y-C2L, and Y-P2L samples). For all cases, the process behavior was identical for both sides of the weld. 3.1.1. Weld appearances Deposited weld appearances and macrosections are shown in Fig. 5. The appearance of the weld beads was very similar for both I- and Ygroove samples, due to almost identical parameters being used (only slightly more WFR was required to fill the gap in the Y-groove case). Therefore surface views are only shown for the I-grooves. Because of plate distortion during welding, the penetration depth can vary as slight gap variations occur during processing (in some cases also causing a little underfill). To prevent these issues from occurring, industrial implementation of thick sheet welding usually involves hydraulic clamping. The trailing torch provided slightly less spattering and higher quality weld seam appearance compared to the leading torch arrangement for both groove types. Spatter generation is probably linked to arc instabilities, causing changes in drop flight trajectories, possibly causing variations in weld seam appearance. An exception is the I-C1T sample (see Fig. 5(a)) that unintentionally had slightly shorter DLA (2.5-3 mm), resulting in occasional deposited droplet explosions and chopping of the wire instead of the arc forming at the wire tip. The pulsed arc mode provided better weld surface quality due to a lower incidence of arc and drop detachment instabilities such as wire chopping and droplet explosions.
3.1.3. Deposition trajectory of molten droplets The trajectory of the molten droplets can play an important role and should be considered when optimizing welding parameters. Material transfer from the wire frequently experiences deflection, or deviation, of droplet trajectory (see Fig. 6). For both trailing and leading arcs, the droplets tend to veer away from the gouge front due to changes in arc shape, but for different reasons. For the leading arc condition, the arc produces a gouge and pushes away the melt beneath so that the arc has a solid elongated surface from which to form, pushing the melt at the wire tip upwards so that the detached droplets fly for a longer distance (the pinch-effect is also reduced), Fig. 6(j). With a trailing arc, lower levels of ionization in the shielding gas (due to the presence of metal particles) shifts the arc central position closer to the keyhole. The sum of the arc force on the molten wire tip is then such as to deviate the droplet path from the weld front to the rear Fig. 6(i). The magnitude of drop trajectory deviation is strongly related to the welding parameters. The deviation is stronger when the arc power is higher and when the arc length is shorter (e.g. Fig. 6(e), I-C1T sample). Greater flight path deviation with higher WFR is also related to MDPP, where the droplets have smaller mass and more easily affected by the pressure from a stronger electron flow from the gouge front. Onedrop-per-pulse (ODPP) usually experiences less change in droplet path trajectory due to the larger mass of the droplets. Also, the flight path deviation during CMT+P was found to be reduced and the pinch-effect was less attenuated during arc pulsing, probably because of lower arc power levels. Notably, the deflection of droplet trajectory is not constant over time. For a leading torch the deflection of trajectory is visible for
3.1.2. Process behavior Fig. 6 contains illustrations of the process behavior recorded by HSI for all cases in set (i) and set (ii), showing arc sizes, melt pools, arc gouge shapes, melt flows, and droplet behavior. Droplet impingement is also observed and corresponding areas (as blue circle) are indicated. No significant differences were observed between the I- and Y-groove preparation samples. The pulsed arc provided larger arc sizes due to higher power levels, producing deeper gouges, especially in the leading torch arrangement. The leading torch arrangement slightly reduced 37
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Fig. 5. Macrosections and deposited weld seam appearances of I- and Y-groove samples.
more than half the flight time (up to 70%) compared to less than half time for the trailing torch (up to 40%). In the trailing torch arrangement, the increase of WFR which can disturb the keyhole dynamics due to excessive molten metal overflow is less of a problem, compared to the leading torch, due to a greater level of droplet trajectory deflection away from the keyhole. With a leading arc, the droplets were frequently landed close to the keyhole opening area, Fig. 6(c, d). The droplet impingement area in this case is close to the keyhole front and sometimes droplets strike the keyhole opening area. As a result, the keyhole is more prone to collapse and subsequent porosity may form. Note that although MDPP occurs more for a leading torch, the droplets are smaller than for ODPP. Therefore, each droplet has less effect on the keyhole collapse and porosity formation can be reduced
Fig. 5(e, f, m, n). Similar results related to porosity generation were obtained by Cao et al. [24]. Porosity was more severely pronounced in the Y-grooved specimens with both arc modes, see Fig. 5(h, d). This is probably due to droplets striking the keyhole opening and more vigorous internal melt flow. Y-groove specimens showed similar results to the Igroove samples as regards flight path deviations. However, it was noted that an increase of WFR by 1 m/min for the CMT+P arc (heat output increased by 27%) in Y-groove welding is characterized by double the amount of peak current pulses (from 6 to 12). 3.1.4. Wire chopping The wire chopping phenomenon observed during experiments is of particular interest because it has not been discussed or observed in pre38
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Fig. 6. a-h) LAHW process stability based on HSI for I-groove specimens where: 1 – arc plasma; 2 – droplet impingement area; 3 – molten droplet; 4 – melt flow; 5 – keyhole; 6 – air gap; 7 – variable arc length (where LA is arc length); and 8 – arc gouge (not covered by molten metal). i) droplet trajectory deviation for trailing and j) leading arc. k) melt is pushed from the gouge and l) during the short-circuiting phase the gouge is filled with melt.
vious scientific papers. However, evidence of the phenomenon can be found in earlier work done by Wahba et al. [26] who used solid wire. The detached wire can occur as spatter beside the weld seam, as shown in Fig. 5(aii ), or can possibly strike the keyhole opening, leading to a subsequent collapse. This could explain occasional macro porosities, see Fig. 5(ai ). The chopping mechanism could be related to metal-cored wire design, as solid wires are believed to be less prone to wire chopping. Wire chopping was observed to occur more frequently when using higher travel speeds, trailing torch arrangement, and CMT+P arc mode, indicated in Fig. 6. Figure 7 shows a sample with wire chopping occurring for CMT+P arc mode at 0.8 m/min vt . Streak imaging was used to analyze wire chopping phenomenon, see Fig. 7(d). Streak imaging uses a collection of stacked lines (one pixel wide) extracted from each frame of the video (vertical line in Fig. 7(c)). In the streak image, the perpendicular axis becomes time dependent, showing the time evolution of the chosen axis and allows for periodic behavior of the process to be analyzed and represented in a single picture. Frostevarg et al. [8] have applied this technique to examine causes of undercutting in different arc modes since it facilitates the analysis of complex phenomena. For wire chopping analysis, the streak line was selected at the wire at certain distance from the tip (along y axis direction) where wire chopping most frequently occurred. From this streak image, Fig. 7(d), it is possible to identify the relationship between wire chopping and a preceding short-circuiting phase that is of unusually short duration. The A-D positions in Fig. 7(d) cor-
responds to HSI frames in Fig. 7(e-h), showing the wire chopping event (A-B), and demonstrating that the arc then becomes larger until the process is stabilized again (C-D). In the CMT+P case presented in Fig. 7(d) the time required to re-stabilize the arc after the chopping event is about 27 ms, which correspond to the next short-circuiting phase (C). During re-stabilization, the arc length will be longer than normal (B), possibly causing disturbances on the weld pool and the resulting weld bead profile. Fig. 8 illustrates the wire chopping mechanism. Wire chopping occurs after an unusually short short-circuiting phase due to the wire not being sufficiently retracted to an appropriate distance from the molten pool. After such an event, the arc is very short. Before wire chopping, it seems that there is a buildup of heat in the wire with each pulse, see Fig. 8(a). The arc length gradually becomes longer with each pulse, but not reaching its normal length. Resistance of heated wire increases as its temperature approaches to the melting point. At some point after 3-5 pulses, the wire is prematurely melted up to a certain distance Fig. 8(b). At this distance, the wire is chopped off by Pinch-force during subsequent pulse (peak current) due to melting with increased resistance in this area, Fig. 8(c). Deposition trajectory of the chopped wire is unpredictable, depending on a combination of gravity and the sum of forces from the arc.
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Fig. 7. Deposited weld seam appearance (a, b) and HSI (55° inclined camera from surface) analysis of LAHW (Y-C1T sample) through streak image (d) technique of each selected line showing unstable processing (e) and stable (h) processing.
Fig. 8. Wire chopping mechanism.
chopping was observed when the arc length was very short. An increase of WFR by 1 m/min for the I-groove preparation slightly improved the stability of the process, compared to set (i) experiments. Spattering was significantly reduced due to a reduction in wire chopping.
3.2. Process behavior when changing process settings 3.2.1. Increase of joint air gap Increasing the air gap increases the requirement for filler material, as insufficient deposition rates can produce slight underfill, e.g. Fig. 9(a) (sample I-P3T). An increase of air gap from 0.4 to 0.6 mm for high (sample I-C5T; 9 m/min WFR and 0.8 m/min vt ) and low (sample I-C6T; 7 m/min WFR and 0.6 m/min vt ) level parameters for the trailing CMT+P arc mode, did not significantly improve the stability of the process. It is likely that the increase in gap width (0.2 mm) was too small to have any noticeable effect. Increasing the air gap to 0.8 mm (samples I-C3T, I-P3T, Y-C3T, and Y-P3T), slightly increased process stability, providing MDPP metal transfer for both trailing and leading arc modes (with 10 m/min WFR and 0.8 m/min vt ). For CMT+P in the I-groove case only, wire
3.2.2. Process stability after further increase in travel speed, wire feed rate and air gap Using CMT+P, I-groove joint preparation and increasing vt from 0.8 to 0.95 m/min (sample I-C7T), the process became very unstable with frequent wire chopping (of different lengths), a significant amount of spatter and some porosity. For both I- and Y-groove preparation, using pulsed arc modes and increasing the WFR to 13 m/min (samples I-P8T and Y-P4T), resulted in MDPP-to-spray droplet transfer and significantly less wire chopping with some underfill, as shown in Fig. 9(b) (sample 40
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free-flight mode, between short-circuits in CMT+P as shown in Fig. 11(ac) (sample I-C4T). Constant short-circuiting (direct contact with molten pool) is provided without chopping and the weaker arc pulses ensure calm and stable melt flow around the keyhole. Similar observations of LAHW with CMT were made by Frostevarg et al. [8]. As a result, no spatter was generated (see Fig. 11(d)). Streak imaging (Fig. 11(e)) reveals that duration of short-circuiting is stable, providing appropriate melting of the electrode wire (see Fig. 11(b, c)). This indicates that at lower travel speeds the rooting effect is well established. Despite the excellent process stability, the penetration depth was significantly reduced to 15-16 mm. From HSI it was identified that melt flow velocity were significantly slower compared to experiments with higher WFR and travel speed, indicating that the arc melt pool flow affects and reduces the penetration capability of the laser keyhole. 3.2.4. Stability effects of varying the laser-arc inter-distance In most welds, the DLA was 3.5-4.5 mm and no clear stability trends could be isolated or observed within this interval. To better observe the effect of different DLA upon process stability, the distance was increased to 8 mm for the leading arc and decreased to 2 mm for the arc trailing case. When using a leading pulsed arc and increasing the DLA from 4 mm to 8 mm (Y-P8L and Y-P9L samples), vt at 0.8 m/min and 0.6 mm air gap, the process became more stable for both 9 and 13 m/min WFR. There was a very low arc length variation, less spatter, and the droplets from the wire no longer impinged the keyhole opening area, see Fig. 12(a). In addition, no porosity formation was observed, (this porosity is probably linked to keyhole collapses, and none were observed in this case). Observations showed a low degree of arc and laser interaction and droplet transfer mode is still in MDPP-to-spray transfer mode, Fig. 12(ai, ii ), probably also related to increased WFR. As a result, in this configuration allows utilizing higher WFR there is a low probability of process instability. On rare occasions, due to arc pressures on the melt, the height of the melt around the keyhole was observed to increase (with waves in the melt pool), but this did not seem to disturb the process. When having a trailing pulsed arc and decreasing the DLA to 2 mm (samples Y-P6T and Y-P7T), with a 13 m/min WFR and using high vt at 0.8 m/min with a 0.6 mm air gap, the process was observed to be stable. Even though the arc was closer to the keyhole, droplet trajectory deviation away from the melt pool front was increased, so that droplets did not land in the vicinity of the keyhole Fig. 12(b). Also, the keyhole was not observed to be covered by melt. Therefore, porosity was not found in the samples. Spatter levels were low, Fig. 12(biv ), and wire chopping was not observed. At 9 m/min WFR, the droplet transfer mode alternated between globular with substantial droplet trajectory deviations as shown in Fig. 12(bi, ii ) (Y-P6T sample) and a MDPP resembling molten stream (see Fig. 12(biii )).
Fig. 9. The effect of WFR, travel speed and air gap on quality of welds in pulsed trailing arc.
I-P8T). Very low amounts of spatter were also produced (Fig. 9). Increasing the WFR to 17 m/min and vt to 1.0 m/min in a Y-groove setup (Y-P5T sample) eliminated wire chopping while also reducing porosity and spattering. Frostevarg et al. [8] has shown that at higher weld speeds, higher gap widths have an arc stabilizing effect. These observations verify that these trends are also valid for thick plate welding. Further increasing vt to 1.2 m/min and air gap to 1.0 mm (I-P9T sample), the WFR was also increased to 18 m/min to fill the gap properly. Due to increased vt at 1 mm gap, there is a significant underfill with large areas of premature solidification as shown in the Fig. 10(a-c) HSI sequences. The process was found to be relatively stable and the arc showed only very slight variations (Fig. 10(c)). However, due to the gap width, the laser keyhole opening is vertically lower than the top surface of the sheet Fig. 10(a). Therefore there is the melt pool could occasionally overflow into the keyhole (due to arc pressure), which could cause porosity. At higher vt , the reduction in wire chopping is related to increased WFR with MDPP improving the rooting effect. Process stabilization by wider air gaps was not being observed for these samples, even though such effects have previously been observed by Frostevarg et al. [8]. This is probably due to the increased amount of filler material which demands high arc powers. A large arc will produce a much wider gouge than the gap width, and will therefore not be constrained by it.
4. Generalized results From this analysis it can be concluded that high vt can result in unstable welding processing in terms of molten metal droplet transfer, especially for the CMT+P arc mode. Instabilities in filler wire deposition, especially wire chopping, during high vt are related to an inability to establish the so-called rooting effect [4] for the arc, and also to too short
3.2.3. Decreased travel speed and WFR Decreasing the WFR to 4 m/min and vt to 0.5 m/min in the presence of a narrow 0.3 mm gap (I-C4T and Y-C4T samples) resulted in an absence of intermittent pulses and drop detachment with subsequent
Fig. 10. a-c) HSI frames of high travel speed (1.2 m/min) for trailing pulsed arc mode (18 m/min WFR) and d) weld bead appearance. MF is melt flow and PS is premature solidification.
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Fig. 11. HSI at low travel speed (0.5 m/min) and trailing arc with 4 m/min WFR using the CMT mode.
Fig. 12. HSI of different process distances effect on process stability with different torch arrangements (0.8 m/min vt , 0.6 mm gap), where: A – MDPP transfer mode; B – keyhole boundaries; C – droplets impingement area; and D – flight path trajectory deviation.
5. Conclusions
an arc length. This makes the application of metal-cored wire more difficult compared to solid wires. However, instabilities (like wire chopping and problems with droplet detachment) can be reduced by increasing the WFR since stronger arcs have a better rooting effect. However, it should be kept in mind that stronger arcs have a more significant effect on droplet trajectories. In addition, an increase in WFR (in order to fill wider gaps) does not ensure more stable processing since a high WFR can disturb the keyhole, especially with a leading arc and shorter DLA (for both torch arrangements). This verifies the findings of Fellman et al. [14]. In addition, leading arc arrangement is less susceptible to wire chopping due to an elongation of the arc, which is related to a deeper arc gouge that is only partially filled with molten metal. Increasing WFR was also found to change droplet detachment towards MDPP due to the presence of higher temperatures. Fig. 13 is a matrix flow chart (MFC), first adopted by Karlsson et al. [27]. It shows generalized experimental trends. Note that identified trends only are shown for some variable values. Only a few experiments had a proper ODPP, since the arc gets disturbed by the gouge form and metal vapor from laser keyhole. Groove preparation type had only a slight effect on processing therefore it can be concluded that LAHW is fairly tolerant to gap variations and applications of different groove types.
LAHW of thick steel plates using metal-cored filler wire is a very efficient and promising fusion joining process, providing welds of acceptable quality within a certain optimized range of parameters. From the observed process behavior and resulting welds, the following conclusions can be drawn:
•
•
•
•
•
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During LAHW, interactions between the heat sources and the melt flow affects the arc process, creating conditions for ODPP to MDPP or even the spray transfer mode, depending on process settings. For a leading arc, short laser-arc inter-distance is not recommended as there is an increased risk of keyhole collapse causing porosity, due to droplet impingement on the keyhole opening. Short inter-distances can be allowed for trailing arcs without causing keyhole collapse because the arc pressure on the melt causes droplets to fly away from the gouge front. The CMT+P arc mode is not recommended for travel speeds higher than 0.8 m/min due to the occurrence of wire chopping (of the metal-cored wire) which could be deleterious to weld quality. The type of groove preparation was observed to have little effect on process stability.
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Fig. 13. MFC of process stability based on performed experiments for different groove preparations and laser-arc arrangement setup. Notations of effects: 0 is a non-occurring event, (+) is an infrequent or weakly occurring event and (++) is a frequently or strongly occurring event.
•
An air gap between plates only has a significant influence on process stability when the gap is wider than 0.6 mm.
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Acknowledgments The authors wish to thank the Research Council of Norway for funding through the Petromaks 2 Programme, Contract No. 228513/E30, as well as EU-RFCS project OptoSteel. The financial support from ENI, Statoil, Lundin, Total, JFE Steel Corporation, Posco, Kobe Steel, SSAB, Bredero Shaw, Borealis, Trelleborg, Nexans, Aker Solutions, FMC Kongsberg Subsea, Marine Aluminium, Hydro and Sapa are also acknowledged. References [1] Ono M, Shinbo Y, Yoshitake A, Ohmura M. Development of laser-arc hybrid welding. NKK Tech Rev 2002(No. 86):8–12. [2] Moore PL, Howse DS, Wallach ER. Microstructures and properties of laser/arc hybrid welds and autogenous laser welds in pipeline steels. Sci Technol Weld Join 2004;9(4):314–22. [3] Gook S, Gumenyuk A, Rethmeier M. Hybrid laser arc welding of X80 and X120 steel grade. Sci Technol Weld Join 2014;19(1):15–24. [4] Steen WM. Arc augmented laser processing of materials. J Appl Phys 1980;51(11):5636–41. [5] Kristensen JK, Webster S, Petring D. Hybrid laser welding of thick section steels ˗ The HYBLAS project. In: Proc. of 12th NOLAMP 2012 Conference, Copenhagen, Denmark; 2009 (August 2009). [6] Katayama S, Kawahito Y, Mizutani M. Elucidation of laser welding phenomena and factors affecting weld penetration and welding defects. Phys Proc 2010;5:9–17. [7] Eriksson I, Gren P, Kaplan AFH. New high-speed photography technique for observation of fluid flow in laser welding. Opt Eng 2010;49(10):100503. doi:10.1117/1.3502567. [8] Frostevarg J, Kaplan AFH, Lamas J. Comparison of CMT with other arc modes for laser-arc hybrid welding of 7 mm steel. Welding World 2014;58(5):649–60. [9] Frostevarg J. Comparison of three different arc modes for laser-arc hybrid welding steel. J Laser Appl 2016;28(2):15–35. 43
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[27] Karlsson J, Kaplan AFH. Analysis of a fibre laser welding case study, utilising a matrix flow chart. Appl Surface Sci 2011;257(9):4113–22. [28] Hu B, den Ouden G. Laser induced stabilisation of the welding arc. Sci Technol Weld Join 2005;10(1):76–81.
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