Analysis of additional electromagnetic force for mitigating the humping bead in high-speed gas metal arc welding

Accepted Manuscript Title: Analysis of additional electromagnetic force for mitigating the humping bead in high-speed gas metal arc welding Author: Y. Li C.S. Wu L. Wang J.Q. Gao PII: DOI: Reference:

S0924-0136(15)30121-7 http://dx.doi.org/doi:10.1016/j.jmatprotec.2015.09.014 PROTEC 14549

To appear in:

Journal of Materials Processing Technology

Received date: Revised date: Accepted date:

23-7-2015 4-9-2015 5-9-2015

Please cite this article as: Li, Y., Wu, C.S., Wang, L., Gao, J.Q., Analysis of additional electromagnetic force for mitigating the humping bead in highspeed gas metal arc welding.Journal of Materials Processing Technology http://dx.doi.org/10.1016/j.jmatprotec.2015.09.014 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

Analysis of additional electromagnetic force for mitigating the humping bead in high-speed gas metal arc welding Y. Li, C. S. Wu* [email protected], L. Wang, J. Q. Gao MOE Key Lab for Liquid-Solid Structure Evolution and Materials Processing, and Institute of Materials Joining, Shandong University, Jinan, China *Corresponding author.

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Abstract To mitigate the occurrence of humping bead in high-speed gas metal arc welding, an external magnetic field is exerted into the weld pool to produce the forward electromagnetic force and to brake the backward flow molten jet. A thermal-magnetic coupling model is developed to analyze the distribution of the additional electromagnetic force in weld pool. The interaction of the external magnetic field in the arc region with the arc/liquid metal stream at the wire tip is taken into consideration, and the excitation current and the wire-magnet distance are optimized. The welding experiments on mild steel plates (Q235B) demonstrate that with help of the forward additional electromagnetic force, good weld bead quality without humping is obtained in high-speed gas metal arc welding. Keywords: External magnetic field; high-speed GMAW; numerical analysis; additional electromagnetic force; humping bead

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1. Introduction Gas metal arc welding (GMAW) is the most widely used materials joining process in manufacturing industry due to its advantages of low cost, easy operation and good adaptability. When it is employed to weld the structures of thin metal plates, higher welding speed is preferred to achieve higher productivity, as pointed by Nguyen et al. (2006). However, when the welding speed exceeds a certain value, humping bead is formed and deterioration of the weld bead quality will occur. Nguyen et al. (2005; 2007) found that larger levels of both arc force and droplet impingement associated with higher welding current in high-speed GMAW results in a gouged region at the front of the weld pool, which makes the molten metal flows backward with very high velocity. Wu et al. (2009a; 2009b) conducted experimental observations of the weld pool behaviors in high-speed GMAW, and found that the backward flow molten jet with high momentum in weld pool is responsible for the formation of humping bead. Chen and Wu (2010) also developed a mathematical model to analyze the effect of the flow velocity of molten jet inside the weld pool on the formation of humping bead. For the sake of suppressing humping bead in high-speed GMAW, it is critical to reduce or mitigate the momentum of the backward flow molten jet in weld pool. Recently, some investigators have developed modified processes and equipment to decrease the momentum of the backward flow molten metal in weld pool. Kiran et al. (2011) used two-wire tandem submerged arc welding for probing the influence of welding current on weld quality. Qin et al. (2015) developed high speed tandem gas tungsten arc welding process for welding thin stainless steel plates. Michie et al. (1999) investigated the process characteristics and applications of twin-wire GMAW. Meng et al. (2014) combined a tungsten inert gas arc with a metal inert gas arc to increase the welding speed. Li et al. (2007) developed a double-electrode GMAW process. Wu et al. (2012) used a double-electrode GMAW system to improve the weld bead quality. All these process modifications aim to partition the total welding current among the multiple wires/arcs. In this way the force from each arc is lowered and the distribution region of arc pressure is widened, which results in a weaker surface depression of weld pool and a lower momentum of the backward flow molten metal in weld pool. Choi et al. (2006) employed laser-GMAW hybrid welding to improve the fluid flow and heat transfer status in weld pool so that the momentum of the 3

backward flow molten jet is decreased. However, more electrodes or heat sources could cause the process complexity, operation difficulty and larger investment of equipment. Thus, it is of great significance to take low-cost and flexible technical measures to manipulate the backward flow molten jet in weld pool. Because magnetic control unit is cost effective and easy to implement, as reviewed by Li et al. (2015), it is an effective means to control the arc/weld pool behaviors and the weld quality by applying external magnetic field in various welding processes. Kou and Le (1985) used low frequency transverse magnetic field to make arc oscillation along welding direction, and found it can improve the quality of joints. Nomura et al. (2012) used a cusp type magnetic field produced by four magnetic poles to change the cross section of arc plasma from circular to elliptical shape, and good bead appearance was obtained in high speed welding. Shoichi et al. (2013) used a magnet to produce a upward electromagnetic force in hot wire tungsten inert gas welding so that the molten metal is lifted up and excessive sag of underside weld is prevented. In laser beam welding, Avilov et al. (2012) applied a noncontact inductive electromagnetic weld pool support system to suppress gravity dropout of the melt and eliminate sagging of the weld pool root side surface when AlMg3 plates of up to 30 mm thickness is welded. All these works either changed the arc-pool surface contact area or produced supportive force to weld pool to overcome the gravity effect, but did not concern with appropriate adjustment of fluid flow in weld pool. Kern et al. (2000) firstly applied transverse magnetic field to suppress humping bead in CO2 laser beam welding, and found that the Hartmann effect and thermoelectric voltage induced by temperature gap between weld bead and weld pool are beneficial to suppress humping bead. Zhou and Tsai (2007) used an external electromagnetic force to increase the back filling speed of the liquid metal during the keyhole collapse process for preventing the development of porosity in laser welding. Bachmann et al. (2012) used a steady external magnetic field to control fluid flow dynamics in laser beam weld pool, and the wineglass shape of the weld bead on the transverse cross-section could almost completely be eliminated. Bachmann et al. (2013) developed a numerical model to investigate the influence of a steady magnetic field on the flow pattern in the melt, and a braking of the flow velocities in the weld pool due to the Hartmann effect was observed in laser beam welding. However, the aforementioned cases are all related to laser beam welding where there is no 4

welding current flowing in weld pool, and there is no strong interaction between the welding current with the external magnetic field. In order to slow down the backward flow molten jet in weld pool during high-speed GMAW, Yang et al. (2014) developed an experimental system of high-speed GMAW with an external magnetic field device to mitigate the momentum of the backward flow molten jet inside the weld pool. Although it could operate with a welding speed up to 2 m min-1, there were severe spatters during the welding process. Wang et al (2015) found that the spatter is resulted from the unfavorable interaction of the external electromagnetic field with the arc column and the liquid metal stream at wire tip. It was experimentally verified that a good weld bead with neither humping nor spatter can be obtained if the external electromagnetic field is appropriately combined with the backward inclination angle of the welding torch. To this end, the distribution of the exerted additional electromagnetic force must be selected and controlled in a suitable way. To effectively employ the external electromagnetic device and optimize its parameters, it is essential to conduct numerical analysis of both the external magnetic field and the additional electromagnetic force in molten pool. In this study, a thermal-magnetic coupling model is developed to analyze the distribution of the exerted magnetic field and electromagnetic force in weld pool. The excitation current level and the relative position of the external magnetic device with respect to the torch are optimized. High-speed GMAW tests are performed to validate the model and the process effectiveness.

2. Numerical simulation of external magnetic field Fig. 1 schematically shows the experimental set up of high-speed GMAW with the external magnetic field apparatus. The external magnetic field is produced by the two coils wounded around each end of the magnetic core. The shape and size of the magnet are described in Fig. 2a, and the turn number of each coil is 320. The main function of the external magnet is to produce a transverse magnetic field in weld pool, which interacts with the welding current in weld pool so that a forward electromagnetic force is produced inside the weld pool, and the backward flow molten jet is slowed down. To minimize the unwanted effect of the external magnetic field on the welding arc, the magnet apparatus is installed beneath the workpiece. For generating additional electromagnetic force towards the welding direction in weld pool, two coils straddle the welding line (seeing Fig. 1), and the flowing 5

direction of the excitation current in the coils must ensure that the direction of the external magnetic field is parallel to the workpiece surface and perpendicular to the welding line (along y-direction). During the welding process, the workpiece is driven to move along the welding direction (positive x-direction) while both the torch and the magnet device are stationary.

2.1 Model for external magnetic field Based on the arrangement of external magnetic field apparatus, it needs to establish a three dimensional finite element model to calculate the distribution of the magnetic field strength in the whole domain including the excitation coils, magnet core and workpiece. In addition, the influence of external magnetic field on the arc during welding should be considered. Thus, a cylinder subdomain with dimension of 20 mm in diameter and 20 mm in height is used to calculate the distribution of external magnetic field in the arc area, as shown in Fig. 2b. The size of workpieces is 180 mm in length, 60 mm in width, and 3 mm in thickness. Because the excitation power produces a steady magnetic field, the three dimensional static magnetic scalar method of ANSYS electromagnetic module is used to calculate the external magnetic field. SOURCE36 loop current element is adopted to simulate the magnet coils, which is a specific element that does not need to generate finite element mesh for the magnet coils. The thermal-magnetic coupling element of SOLID96 is used for the other parts of magnetic domain. Finer mesh is generated around the weld pool and the arc zone, while coarse mesh is used for the rest regions to enhance the calculation efficiency. In global Cartesian coordinate system of the calculation domain, the Maxwell's equations are written as follows: B = ∇× A

(1)

∇Ax =

∂ 2 Ax ∂ 2 Ax ∂ 2 Ax + 2 + 2 = − µ Jx ∂x 2 ∂y ∂z

(2)

∇Ay =

∂ 2 Ay ∂ 2 Ay ∂ 2 Ay + 2 + 2 = −µ Jy ∂x 2 ∂y ∂z

(3)

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∇Az =

∂ 2 Az ∂ 2 Az ∂ 2 Az + 2 + 2 = − µ Jz ∂x 2 ∂y ∂z

(4)

where B is the magnetic flux density, A is the magnetic potential vector, Ax, Ay and Az are the three components of the magnetic potential vector A, µ is the magnetic permeability of medium, and Jx, Jy and Jz are the three components of excitation current density. The magnetic flux density B can be obtained by Eq. (5) Bx =

∂Az ∂Ay − ∂y ∂z

By =

∂Ax ∂Az − ∂z ∂x

Bz =

∂Ay ∂Ax − ∂z ∂x

(5)

The boundary conditions of external magnetic field calculation are written as follows: Hn =

Ht =

1 ∂Az =0 µ ∂t 1 ∂Az

µ ∂n

(6)

=0

(7)

where H is the magnetic field strength, and Hn and Ht are the normal and tangential components of the magnetic field strength. Eq. (6) means that there is no normal component of magnetic field strength along the tangential direction of magnet core surface, and this boundary condition is applied on the outer surface of all magnetic medium to reduce divergence of magnetic field as soon as possible. Eq. (7) indicates that the magnetic lines of flux pass the surface of magnetic medium in perpendicular direction, and this boundary condition is imposed on the end surface of magnet core and the lateral surface in longitudinal direction of workpiece.

2.2 External magnetic field before welding First, the distribution of external magnetic field on the workpiece is calculated before welding. The turn number of each coil is 320, the excitation current is 5 A, and the air gap between the top surface of the magnet and the bottom surface of the workpiere is 5 mm. The values of magnetic permeability for air and mild steel (Q235B) workpiece are 1 and 200, respectively. Fig. 3 is the calculated magnetic flux density vector B on the workpiece surface before welding. It can be seen that the magnetic flux density vector in the workpiece is distributed basically along y direction (perpendicular to the welding direction), while its components in x direction (welding direction) and z direction (workpiece thickness direction) are quite low. 7

This is indeed the transverse magnetic field, which is exactly what is needed. Thereby, only the transverse magnetic field strength By is used hereinafter. Since the variation of By in z direction is only about 0.1 mT when the plate thickness is 3 mm, only the distribution of By along x and y direction are considered. Fig. 4 shows the predicted magnetic flux density By on the workpiece surface in x and y directions. The distribution of By is symmetric with respect to the action point of magnet (x=0), and at this location there exists the maximum value of By in x direction (66.4 mT). Away from this point, the magnetic field intensity decreases, but its dropping extent is small (about 10 mT). This indicates that a relatively uniform distribution of transverse magnetic field exists in the zone around the action point of magnet (x=0) before welding. For the distribution of By in y direction, two peak values locate at the workpiece edges, just as demonstrated in Fig. 3. To validate the magnetic field model, the transverse magnetic flux density in the workpiece is measured by a Tesla-Meter. Ten points near the center of magnetic pole is taken to measure the transverse magnetic flux density, and the distribution of By at these detection points is measured with the excitation current of 5 A. Fig. 5 shows that the calculated and measured data of transverse magnetic flux density in the workpiece under the same excitation situation before welding are in good agreement.

2.3 External magnetic field in weld pool The mild steel workpiece (Q235B) is a ferromagnetic material, and its magnetic permeablity changes with temperature variation. Therefore, the measured data of magentic flux intensity before welding process cannot represent the actual distribution of electromagnetic field in weld pool during welding. To calculate the distribution of electromagnetic field in weld pool, thermal-magnetic coupled analysis method in software ANSYS is used to solve the three-dimensional heat conduction and Maxwell's differential equations iteratively. The influence of high temperature in welding process on the permeability of ferromagnetic workpiece is taken into account. When the temperature is above the Curie temperature (Curie point, 770 oC for mild steel), the permeability of ferromagnetic workpiece decreases and approaches to the value in vacuum. Table 1 lists the welding conditions used in this study. First，Case 1 is used to examine 8

the variation of magnetic field intensity when the temperature field is considered. Fig. 6a demonstrates the distribution of the temperature in quasi-steady state on the top surface of workpiece in welding. The zone with red color in Fig. 6a is the weld pool. Fig. 6b compares the calculated magnetic field intensity along the weld line (y=0) under the conditions with and without coupling the thermal field. The magnetic flux density is very low in the workpiece region heated by the welding arc directly (the weld pool and the solidified weld bead). In the workpiece region ahead of the weld pool, the magnetic flux density is much higher because the temperature there is relatively much lower. Therefore, the magnet can be arranged behind the torch axis, and the wire-magnet distance (Dwm) can be selected as a critical parameter to adjust the distribution of external magnetic field in weld pool. In this way, the value of magnetic flux density in weld pool is ensured to satisfy with the requirement of adjusting backward flow molten jet. Fig. 7 shows the distribution of transverse magnetic flux density in weld pool under different wire-magnet distances, and the welding direction goes in positive x-direction. As the wire-magnet distance (Dwm) increases, the value of By in the front of weld pool increases. It is clear that the magnetic field strength in weld pool is completely changed by the wire-magnet distance, and the value of transverse magentic flux density is maximal if the wire-magnet distance is 15 mm.

3. Deflection of the arc/liquid metal stream In high-speed GMAW, higher level of welding current is usually employed to ensure the heat input. With Ar-rich shielding gas, a liquid metal stream (column of liquid metal) is formed at the wire tip, and its end disperses into droplets which are transferred into the weld pool. When a transverse magnetic field is applied to high speed GMAW, both the arc and the liquid metal stream may be affected, even this effect is minimized by assembling the magnet at underside. Therefore, it is necessary to investigate this effect.

During high-speed GMAW, the images of the arc and liquid metal stream are captured by a camera with a filter. The welding conditions are Case 3 in Table 1, with the flow rate 20 L/min of shielding gas (Ar+8% CO2 ). For the used experimental system shown in Fig.1, the applied magnetic field is along the y-direction to produce a forward electromagnetic force. As shown in Fig. 8, both the arc and the liquid metal stream are pushed forward by the 9

external magnetic field. With increasing of excitation current (associated with higher external magnetic field), more sever deflection of the arc and the liquid metal stream is formed. Fig. 9 illustrates schematically the forward deflection of the arc and the liquid metal stream. Under the action of external magnetic field, the interaction of the current in the arc column/the liquid metal stream with the external magnetic field results in a forward electromagnetic force Fx, which pushes the arc and the liquid metal stream forward in welding direction. For simplification, the current density on any cross-section of the liquid metal stream is written as: j = R =

I

π R2

（8）

R bH − ( R b − R a ) z H

（9）

where j is the welding current density, I is the welding current, R is the liquid stream radius at any cross-section, Ra is the wire radius, Rb is the droplet radius, H is the vertical length of liquid metal stream, and z is the longitudinal coordinate. Therefore, with applied external transverse magnetic field, the additional electromagnetic force acting on the liquid metal stream along x direction can be obtained: f ( z ) = ( j × B ) x = By( z ) ⋅

I

H2

π [ R b H − ( R b − R a ) z ]2

（10）

where By(z) is the transverse magnetic flux density in the arc region. Correspondingly, the electromagnetic force F acting on a droplet at the end of liquid metal stream is as follows: F = f (0) ⋅ Vd

（11）

where f (0) is the additional electromagnetic force at the end of liquid metal stream, and Vd is the droplet volume. Then, the force F causes the x-direction acceleration of the droplet, i.e.,

ax =

F f (0) = m ρ

（12）

where m is the droplet mass，and ρ is the droplet density. Therefore, the applied transverse external magnetic field makes the droplet move a distance (s) forward, which may be written as follows:

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1 1 h 2 s = ax ( ∆t ) = ax   2 2 v

2

2

1 f (0)  h  1 I h s= B y (0) ⋅   = 2 ρ v 2ρ π Rb2  v 

（13） 2

（14）

where h is the distance between the end of liquid metal stream and the workpiece surface， and v is the downward velocity of droplet. Based on the movement of droplet, the offset distance between the liquid metal stream axis and the wire axis can be obtained: 2

s( z) =

θ (z) =

1 f (z)  h  1 I H2 h By( z )   = 2  2 ρ v 2ρ π [ R b H − ( R b − R a ) z ]  v 

180

π

 s(z)  arctan    H 

2

（15）

（16）

where θ is the deflection angle between the liquid stream centerline and the wire centerline (Fig. 9). Through combining the distribution of external magnetic field in the arc region, the offset distances between the liquid stream centerline and wire centerline with different excitation conditions are calculated. The measured data are also obtained by processing the captured images of liquid metal stream. As shown in Fig.10, the calculated results are in good agreement with the measured data, especially near the end of liquid metal stream (z=0.0-0.5 mm). However, there is a relatively big difference between the measured and predicted results at the upper part of liquid metal stream (from z=1.0 mm to z=2.5 mm). The reason is that Eqs. (15)-(16) are based on a few simplified assumptions. Two of them are as follows: (1) Eq.(15) implies that the offset distance s(z) is determined according to the forward movement of droplet at the end of liquid metal stream, and (2) Eqs. (8)-(9) are based on the assumption that the liquid metal stream is a truncated cone and the current density on any cross-section of the liquid metal stream is uniform. Nonetheless, the calculated offset distance s(z) near the end of liquid metal stream (z=0.0-0.5 mm) matches well with the measurement, and will be used to calculate the offset distance on the workpiece surface through the extrapolated centerline of the liquid stream, seeing Fig. 9 and Eq.(17). Fig. 10 shows that the end of liquid metal stream is pushed forward by a distance of around 2.4 mm. This deflection has two consequences. First, the forward deflection of the arc column changes the distribution range of the current density, the heat flux and the arc pressure on the weld pool surface. Second, the deflection of the liquid metal stream changes 11

the impingement region of droplets on the weld pool. The distribution model of the conduct current density is assumed unchanged after the arc deflection, but its distribution region is varied. The arc offset distance Sarc is considered. As shown in Fig. 9, in the Cartesian coordinate system (the coordinate origin is located at the arc center on the workpiece surface), the distribution function of the conduct current density along z direction is expressed as: j

cz

=

 ( x − s arc ) 2 + y 2   z  exp −3  1 −  2 2 π rH rH L   3I

（17）

where rH is the radius of the arc column on the workpiece surface，Sarc is the arc offset distance, and L is the thickness of workpiece. Then，the component of current density along x direction can be obtained, and the distribution model is treated as uniformly distributed along the workpiece thickness direction. jcx =

 1 − e x p 2π L   I

 ( x − s a rc ) 2 + y 2   | x − s a rc |  − 3   2 2 2 rH    ( x − s a rc ) + y

S a rc = L a rc • t a n θ

（18）

（19）

where θ is the angle between the wire centerline and the arc centerline after deflection, and Larc is the distance from the wire tip to the workpiece, as shown in Fig. 9.

4 . Additional electromagnetic force in weld pool The unidirectional conduct current I (welding current) which flows through the wire into the weld pool is the main source of current density existed in weld pool to induce electromagnetic force. The interaction between the strong unidirectional conduct current density jc and the external transverse magnetic field By produces an additional electromagnetic force Fc in the weld pool,

Fc = jc × B y

（20）

As shown in Fig.11, the components of welding current density in z and x directions are jcz and jcx, which is defined by Eqs. (17) and (18), respectively. At the front of weld pool, jcx interacts with the external magnetic field to produce a upward force Fcz. At the rear part of 12

weld pool, jcx interacts with the external magnetic field to produce a downward force Fcz. Especially, the current density along z-direction jcz interacts with the external magnetic field to produce a forward force Fcx. Based on Eqs. (17) and (18), jcx is much lower than jcz. Thus, the force Fcz is also much lower than the forward force Fcx. And the forward force Fcx is what is needed to slow down the backward flow molten jet in weld pool. In addition, there is Hartman effect in weld pool. The backward flow molten jet cuts the magnetic lines of force and produces the inductive current density. When the later interacts with the external magnetic field, an electromagnetic force FH is produced. According to Li’s study (2015), the ratio of the forward force Fcx to the Hartman force FH is about 300. Thus, only the forward force Fcx is considered in this study. With considering both the thermal-magnetic coupling and the arc/liquid metal stream deflection, the additional electromagnetic force due to the external magnetic field is numerically calculated. The welding condition is Case 2 in Table. 1. Fig. 12 shows the calculated additional electromagnetic forces Fcx and Fcz in the weld pool under different wire-magnet distances. The longitudinal component Fcx of the additional electromagnetic force are mainly distributed in the weld pool within a region from x=-5 mm to x= 7.5 mm beneath the arc, which is due to the centralized distribution of the conduct current density in this region. With increasing of the wire-magnet distance Dwm, the magnitude of Fcx becomes larger, and its distributed region expands a little. Since the arc center on the weld pool surface is deviated to a point ahead of the wire axis by the external magnetic field, the peak value of Fcx is emerged at the front part of weld pool, and the distribution of the additional electromagnetic force Fcx is symmetric with respect to the arc center ahead of the wire axis. The peak value of Fcx should shift closer to the wire center (x=0) as Dwm gets larger, because a lager Dwm is corresponding to a less extent of forward arc inclination (and closer distance between the arc center and the wire axis). However, the peak value of Fcx shifts a little bit to positive x-values between Dwm=0 and 5 mm. This may be resulted from the calculation error from discretization of the calculation domain on the workpiece. The vertical component Fcz of the additional electromagnetic force, induced by the transverse magnetic field and x-direction current density, is mainly distributed in the weld pool within a range from x=-15 mm to x= 7.5 mm. The peak value of the vertical component Fcz is much less than that of the 13

longitudinal component Fcx, and is only about 20% of the later. Ahead of the arc center, Fcz is upward (positive), while Fcz is downward (negative) behind the arc center, which is exactly the case illustrated schematically in Fig. 11. It is clear that the value of Fcx is much greater than Fcz in weld pool, while the distribution range of Fcz is much broader. Fig. 13 is the pseudo-color distribution of the longitudinal component Fcx of the additional electromagnetic force in weld pool. Fcx is mainly distributed in the front part of weld pool beneath the welding arc, and is very low in the rear region of weld pool. Under such a case, near the arc center in the front of weld pool (about 2.3 mm ahead of the wire), the maximum value of Fcx is around 100 kN/m3 when the wire-magnet distance is 15 mm (corresponding to a 2.3 mm arc deflection in this case). Fig. 14 shows the calculated values of additional electromagnetic force in weld pool under different levels of excitation current. The value of additional electromagnetic force at longitudinal direction goes up remarkably with an increase of excitation current. At the same time, the deflection of the arc and liquid metal stream is also affected by the excitation current level. In fact, the additional electromagnetic force Fcx is the main force to slow down the backward fluid flow in weld pool during high-speed GMAW. If the excitation current increases from 6 A to 12 A, the maximum value of the additional electromagnetic force Fcx will rise from 80 kN/m3 to 160 kN/m3. Therefore, it is feasible to mitigate the momentum of backward flow molten jet in weld pool by changing the level of excitation current.

5. Experimental validation To examine the process feasibility of the developed system, i.e., the external magnetic field is employed to slow down the backward flow molten jet and suppress humping bead in high-speed GMAW, welding experiments without and with the external magnetic field are conducted. The experimental conditions are as follows: welding current 310 A, arc voltage 34 V, welding speed 2.01-2.37 m/min, flow rate of shielding gas (Ar+8% CO2) 20 L/min, mild steel (Q235B) workpieces of thickness 3 mm, and wire ER70S-6 of diameter 1.2 mm. Fig. 15 presents the weld bead under different levels of welding speed without external magnetic field. When the welding speed is over 2.01 m/min, humping bead and undercuts occur (Fig. 15a). With further increasing of the welding speed, the distance between adjacent 14

humps is decreased, i.e., it is easier to form humping bead (Fig. 15 b, c, d). Fig. 16 shows the weld bead formation with applied external magnetic field. Comparing the weld beads without/with the applied external magnetic field under constant welding speed in Figs. 15 and 16, it can be seen that the periodic humping beads formed with zero excitation current in Fig. 15 are disappeared under the action of external magnetic field. When the welding speed is 2.01 m/min, the minimum excitation current to suppress humping bead is 4.12 A. With increasing of welding speed, the minimum excitation current used to suppress humping bead is rising (Fig.16b-d). However, for a specific welding speed, there is an appropriate level of excitation current, which cannot be varied in a large range. The experimental results indicate that it is feasible to suppress humping bead and improve weld bead quality by using external magnetic field. It is noteworthy that the optimal excitation current (Fig. 16) becomes twice as high (from 4.12 A to 8.75 A) for an increase of the welding speed of only 20% (from 2.01 m/min to 2.37 m/min). The reason is that a little bit increase of welding speed results in larger variations of the heat input, the temperature profile on the workpiece, the weld pool shape and size, and the location of droplet impingement, etc. Because the correlation between the welding speed and the thermophysical conditions during the welding process is quite nonlinear and very complicated, further deep investigation needs to be conducted to give a reasonable explanation. In addition, some of the weld beads in Fig. 16 are without humping but with undercutting, even though the extent is not so severe. Next step, some technical means will be taken to eliminate such undercutting.

6. Conclusions (1) With considering the effect of the arc /liquid metal stream deflection, a thermal-magnetic coupling model is developed to analyze the transverse magnetic field and the additional electromagnetic force in weld pool. (2) The influences of excitation current level and the wire-magnet distance on the additional electromagnetic force in weld pool are quantitatively analyzed. For the study cases, the additional electromagnetic force achieves its peak value at the front of weld pool when the wire-magnet distance is 15 mm. (3) The additional electromagnetic force can effectively mitigate the humping bead in 15

high-speed GMAW. With the excitation current of 4.12 A, GMAW can be carried out at the welding speed of 2.01 m/min without humping bead. (4) As the welding speed increases, the minimum excitation current required to suppress humping bead is rising, and the correlation between them is obtained.

Acknowledgements The authors are grateful to the financial support for this research from the National Natural Science Foundation of China (Grant No. 51275276) and the Research Fund for the Doctoral Program of Higher Education of China (Grant No. 20120131130009).

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References Avilov V., Gumenyuk A, Lammers M, et al, 2012. PA position full penetration high power laser beam welding of up to 30 mm thick AlMg3 plates using electromagnetic weld pool support. Sci. Technol. Weld. Join.,17(2): 128-133. Bachmann M., Avilov V., Gumenyuk A., 2012. Numerical simulation of full-penetration laser beam welding of thick aluminium plates with inductive support. J. Phys. D: Appl. Phys., 45(3), 035201. Bachmann M., Avilov V., Gumenyuk A., Rethmeier M., 2013. About the influence of a steady magnetic field on weld pool dynamics in partial penetration high power laser beam welding of thick aluminium parts. Int. J. Heat Mass Trans, 60, 309–321. Chen J., C. S. Wu C.S., 2010. Numerical analysis of forming mechanism of hump bead in high speed GMAW. Weld. World, 54, R286-R291. Choi H. W., Farson D. F., Cho M. H., 2006. Using a hybrid laser plus GMAW process for controlling the bead humping defect. Weld. J., 85(8), 174s-179s. Kern M., Berger P., Hugel H., 2000. Magneto-fluid dynamic control of seam quality in CO2 laser beam welding. Weld. J., 79(3), 72s-78s. Kiran D. V., Basu B., Shah A. K., Mishra S., A. De A., 2011. Probing influence of welding current on weld quality in two wire tandem submerged arc welding of HSLA steel. Sci. Technol. Weld. Join., 15, 111-116. Kou S., Le Y., 1985. Improving weld quality by low frequency arc oscillation. Metall. Trans. A, 16(10),1887–1896. Li K. H., Chen J. S., Zhang Y. M., 2007. Double-electrode GMAW process and control. Weld. J., 86, 231s-237s. Li Y.B., Li D.L., Lin Z.Q., David A., Feng Z., Tang W., 2015. Review: magnetically assisted resistance spot welding. Sci. Technol. Weld. Join., http://dx.doi.org/10.1179/1362171815Y.0000000059. Li Y., 2015. Numerical simulation of external magnetic field for adjusting the backward flow jet in weld pool during high-speed GMAW. M.S. Thesis, Shandong Univetsity, China. Meng X.M., Qin G.L., Zhang Y.Q., Zou Z.D., 2014. High speed TIG-MIG hybrid arc welding of mild steel plate. J. Mater. Process. Technol., 214(11), 2417-2424. Michie K., Blackman S., Ogunbiyi T. E. B., 1999. Twin-wire GMAW: process characteristics and applications. Weld. J., 78 (5), 31-34. Nguyen T. C., Weckman D. C., Johnson D. A. et al. 2005. The humping phenomenon during high speed gas metal arc welding. Sci. Technol. Weld. Join., 10(4): 447-459. Nguyen T. C., Weckman D. C., Johnson D. A., et al, 2006. High speed fusion weld bead defects. Sci. Technol. Weld. Join., 11(6): 618-633. 17

Nguyen T .C., Weckman D. C., Johnson D. A., 2007. Predicting onset of high speed gas metal arc weld bead defects using dimensional analysis techniques’, Sci. Technol. Weld. Join., 12, 634-648. Nomura K., Ogino Y., Hirata Y., 2012. Shape control of TIG arc plasma by cusp-type magnetic field with permanent magnet. Weld. Int., 26, 759-764. Qin G.L., Meng X.M., Fu B.L., 2015. High speed tandem gas tungsten arc welding process of thin stainless steel plate. J. Mater. Process. Technol., 220, 58-64. Shoichi M., Yukio M., Koki T., Yasushi T., Yukinori M., Yusuke M., 2013. Study on the application for electromagnetic controlled molten pool welding process in overhead and flat position welding. Sci. Technol. Weld. Join., 18, 38-44. Wang L., Wu C.S., Gao J.Q., 2015. Suppression of humping bead in high speed GMAW with external

magnetic

field.

Sci.

Technol.

Weld.

Join.,

DOI

10.1179/1362171815Y.0000000074. Wu C. S., Hu Z. K., Zhang Y. M., 2009a. Suppression of weld-bead defects and increase in the critical welding speed during high-speed arc welding. P. I. Mech. Eng. B-J Eng., 223,751-757. Wu C. S., Zhong L. M., Gao J. Q., 2009b. Visualization of hump formation in high speed gas metal arc welding. Meas. Sci. Technol., 20 (11), 115702. Wu C. S., Hu Z. H., Zhong L. M., 2012. Prevention of humping bead associated with high welding speed by double-electrode gas metal arc welding. Int. J. Adv. Manuf. Tech., 63, 573-581. Zhou, J., Tsai, H.L., 2007. Effects of electromagnetic force on melt flow and porosity prevention in pulsed laser keyhole welding. Int. J. Heat Mass Transfer, 50, 2217–2235. Yang F. Z., Wu C. S., Gao J. Q., 2014. Suppressing of humping bead using an external magnetic field in high speed gas metal arc welding. China Welding, 23(2): 22-28.

18

Figure Captions Power source

+ -

GMAW torch

z y x

Magnetic field direction

Molten pool

Workpiece

Coil

Magnet core

Fig. 1. Schematic of high-speed GMAW with external magnetic apparatus.

19

10

40

120

φ 32 φ 20

15

Coil

Magnet core

20

30

40

(a) (b) Fig. 2. Geometric model for the external magnetic field (11-workpiece; 12-coil 1; 13-coil 2; 16-arc region): (a) dimension of the magnet, (b) calculation domain.

20

Fig. 3. Distribution of magnetic flux density vector B in workpiece (unit: Tesla).

21

150

x- direction y- direction

By (mT)

100 50 0 -50 -100 -30

-20

-10

0

10

20

30

x, y (mm)

Fig. 4. Transverse magnetic flux density (By) on the workpiece surface before welding.

22

Measured Predicted

70 60

By (mT)

50 40 30 20 10 -60

-40

-20

0

20

40

60

x (mm)

Fig. 5. Distribution of transverse magnetic flux density (By) on the workpiece surface before welding.

23

100

Uncoupled Coupled

By (mT)

80 60 40 20 0 -60

-40

-20

0

20

40

60

x (mm)

(a) (b) Fig. 6 The calculated temperature profile (a) and the external magnetic field along weld line on workpiece surface during welding (Case 1).

24

Fig. 7. Effect of the wire-magnet distance on the transverse magnetic field on the workpiece surface (Case 2).

25

(a)

(b)

(c)

(d)

Fig. 8. Captured images of liquid metal stream at wire tip (Case 3): (a) Ie=2 A, (b) Ie=4 A, (c) Ie=6 A , (d) Ie=8 A.

26

Fig. 9. Schematic of the arc and liquid metal stream deflection.

27

3.0 2.5

Measured Calculated

z (mm)

2.0 1.5 1.0 0.5

Wire center

0.0 0.0

0.5

1.0

1.5

2.0

2.5

Offset distance s (mm)

Fig. 10. Deflection distance of liquid metal stream with external magnetic field (Case 4).

28

+

z

Welding direction

x

Arc area Molten pool

jcx

By

Fcz

×

Fcz

Fcx jcz

jcx

-

Fig. 11. Schematic of additional external electromagnetic forces in weld pool.

29

(a)

(b)

Fig. 12. Calculated results of additional electromagnetic force in weld pool (Case 2): (a) Fcx , (b) Fcz .

30

(a)

(b) Fig. 13. The predicted longitudinal component Fcx of additional electromagnetic force in weld pool (Case 5): (a) top view, (b) side view.

31

200 Wire center

160

Fcx (kN m-3)

120 80

Ie= 6 A Ie= 8 A Ie= 10 A Ie= 12 A

40 0 -40 -40

-30

-20

-10

0

10

x (mm) Fig. 14. Effect of excitation current on the additional electromagnetic force in weld pool (Case 5).

32

(a)

(b)

(c)

(d) Fig. 15. Weld bead morphology under different levels of welding speed: (a) v=2.01 m/min, (b) v=2.13 m/min, (c) v=2.25 m/min, (d) v=2.37 m/min.

33

(a)

(b)

(c)

(d) Fig. 16. Effect of external magnetic field on weld bead morphology: (a) Ie=4.12 A (v=2.01 m/min), (b) Ie=5.58 A (v=2.13 m/min), (c) Ie=6.7 A (v=2.25 /min), (d) Ie=8.75 A (v=2.37 m/min).

34

Tables Table1. The welding conditions in different study cases Study Case number

Welding current (A)

Arc voltage (V)

Welding speed (m/min)

Excitation current (A)

1 2 3

300 300 298

30 30 32

1.8 1.8 2.01

5 8 2-8

4

298

32

2.01

5

300

30

1.8

35

Air gap (mm)

Wire-magnet distance (mm)

8

5 3 3 3

0 0-15 5 5

8

3

15