Proceedings of the 7th World Congress on Mechanical, Chemical, and Material Engineering (MCM'21)
Prague, Czech Republic Virtual Conference – August 2021
Paper No. HTFF 130
DOI: 10.11159/htff21.130

Effect of Substrate Orientation on Melt Pool during Multi-Layer
Deposition in V-Groove with Gas Metal Arc
Pradip Aryal1, Kjell Hurtig1, Fredrik Sikström1, Håkan Nilsson2, Isabelle Choquet1
University West
Department of Engineering Science, 461 86 Trollhättan, Sweden
pradip.aryal@hv.se; kjell.hurtig@hv.se; fredrik.sikstrom@hv.se; isabelle.choquet@hv.se
2
Chalmers University of Technology
Mechanics and Maritime Sciences, Fluid Dynamics, SE-412 96 Gothenburg, Sweden
hakan.nilsson@chalmers.se
1

Abstract – Thermo-fluid dynamic and experimental approaches are used to investigate the influence of 20° uphill, downhill and sideway
substrate orientation during metal deposition over a previously deposited bead in a V-groove. The computational fluid dynamic modelwith free surface deformation and metal transfer gives insight into the melt pool flow and causes of defect formation observed on thesolidified beads. The experimental metallographs, high-speed images and computational results show good agreement. It is found thatthe deposition of a second layer on a smooth first layer cooled down to room temperature leads to large changes in melt pool flow patternat 20° substrate inclination compared to flat condition. It results in undercut and humps with the uphill orientation and undercut with theside inclination. Therefore, lower angle range is necessary for multilayer gas metal arc deposition for these two last configurations.
Keywords: metal deposition, gas metal arc welding, V-groove, substrate orientation, melt flow, reinforced bead, hump,

OpenFOAM

1. Introduction

Metal deposition for multi-pass joining and additive manufacturing with gas metal arc (GMA) processing using nonhorizontal substrate orientation would avoid workpiece repositioning necessitating manipulation by large and accurate robots
when manufacturing large components. However, the freedom for tilting the arc axis to maintain it normal to a non-horizontal
substrate can be limited by defect formation such as overflow, undercut and hump. The effect of the welding position on themolten pool behaviour and resulting bead geometry was studied both experimentally (e.g., Park et al. [1]) and numerically(e.g., Hu et al. [2], Cho et al. [3]) with GMA processing, single bead deposit, and flat, overhead, and vertical welding position.
These authors observed that while the flat position generated continuous bead without noticeable defects, the overhead andvertical up positions were prompt to hump formation. Vertical down position increased bead width, while it reduced heightand penetration depth. Besides, it is known that for multi-layer deposition the melt pool can have remarkably different
behaviour in the root pass and in the subsequent layers for which it becomes more sensitive to the process parameters.
However, the effect of non-flat welding position in GMA multi-layer deposition is to our knowledge poorly documented.
This study thus aims at understanding this effect on melt pool thermal flow, free surface deformation, bead geometry, possible
defect formation when depositing above the root pass. It is known from earlier studies that overhead and vertical position
cannot be an option with GMA heat source. Therefore, in this study the welding position is varied by only 20° uphill,
downhill, and sideway to investigate, using both experimental and computational fluid dynamics approaches, whether these
conditions could provide an acceptable process window.

2. Experimental setup

Figure 1(left) shows an overview of the experimental setup. A precise movement ABB robot operated a gas metal arc
torch with TPS-4000 Fronius power supply. It was used to deposit Invar 36 at the travel speed of 7.5 mm/s from a 1.2 mm
diameter electrode wire on Invar 36 plates with same nominal composition. The deposition was shielded by a mixture of
98% Ar and 2% CO2 flowing at the rate of 15 l/min. The plates (200 mm × 143.5 mm × 10.6 mm) were prepared with a Vgroove at 60° with no root gap and mounted on the worktable. This table that can be rotated about its central horizontal axis
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was inclined to position the substrate as illustrated in Fig. 1(right): flat (reference angular position at 0°), downhill (at -20°),
uphill (at 20°), and side inclined (at 20°). A data acquisition system was integrated with this setup to measure the arc current
and voltage at a sampling frequency of 4 kHz. A vision camera was mounted on the robotic arm and focused behind thenozzle region to capture melt pool images at a frequency of 100 Hz. For each substrate positioning two layers of metal were
deposited using the GMA welding system. The torch axis was in each case perpendicular to the substrate, the arc length wasapproximately 5 mm, and uniform contact-tip-to-work distance was maintained. The first layer deposit (root pass) generated
continuous and uniform bead without noticeable defects. It was let to cool down to room temperature and the reinforced bead
was 3D scanned before the second layer was deposited. A fifth-degree polynomial that curve fit the profile of the root passbead reinforcement along the transverse direction was generated.

Fig. 1 : Left - Experimental setup; right - schematic of the substrate orientation: a) flat, b) downhill, c) uphill, d) side inclined.

3. Mathematical model

A transient 3-dimensional melt pool model with metal alloy melting, re-solidification, metal transfer, and free surface
deformation was implemented in the open-source software OpenFOAM®. It assumes incompressible and Newtonian fluids(atmosphere and liquid alloy with the Boussinesq approximation) and laminar flow. The alloy-gas interface is tracked witha volume of fluid method and the solid-liquid alloy transition region with a mushy zone approach. The model consists in thefollowing system of equations governing mass, momentum, thermal energy, and metal volume fraction, respectively:𝜕𝜕𝑡𝑡 𝜌𝜌 + ∇ ∙ (𝜌𝜌u�⃗) = 𝜌𝜌̇ drop(1)

𝜕𝜕𝑡𝑡 �𝜌𝜌𝑐𝑐𝑝𝑝 𝑇𝑇� + ∇ ∙ �𝜌𝜌𝑐𝑐𝑝𝑝 𝑇𝑇𝑢𝑢�⃗� = ∇ ∙ (𝑘𝑘𝑘𝑘𝑘𝑘) + ℎ𝑠𝑠𝑠𝑠 [𝜕𝜕𝑡𝑡 (𝜌𝜌𝑓𝑓𝑙𝑙 ) + 𝛻𝛻 ∙ (𝜌𝜌𝑢𝑢�⃗𝑓𝑓𝑙𝑙 )] + Q̇ arc(3)

2

𝜕𝜕𝑡𝑡 (𝜌𝜌u�⃗) + ∇ ∙ (𝜌𝜌u�⃗⨂u�⃗) = −∇𝑝𝑝 + ∇ ∙ �μ(∇u�⃗ + (𝛻𝛻𝑢𝑢�⃗)𝑇𝑇 ) − 3 𝜇𝜇(𝛻𝛻 ∙ 𝑢𝑢�⃗)𝑰𝑰� + 𝜌𝜌𝑚𝑚 𝑔𝑔⃗ [1 − 𝛽𝛽(𝑇𝑇 − 𝑇𝑇𝑚𝑚 )](1 − 𝑓𝑓𝑙𝑙 )2
𝑑𝑑σ
2𝜌𝜌
𝑢𝑢
�⃗ + �σ 𝜅𝜅𝑛𝑛�⃗ +�𝛻𝛻𝛻𝛻 − 𝑛𝑛�⃗(𝑛𝑛�⃗ ∙ 𝛻𝛻𝛻𝛻)� + 𝑝𝑝arc � |𝛻𝛻𝛻𝛻|+ 𝐹𝐹⃗arc−𝐶𝐶𝐷𝐷 3𝑑𝑑𝑑𝑑
𝜌𝜌𝑙𝑙 + 𝜌𝜌𝑔𝑔
𝑓𝑓𝑙𝑙 + 𝜀𝜀𝐷𝐷(2)𝜕𝜕𝑡𝑡 α + ∇ ∙ (αu�⃗) + 𝐶𝐶α ∇ ∙ [α(1 − α)𝑢𝑢�⃗𝑐𝑐 ] = 𝛼𝛼̇ drop

(4)

where 𝑡𝑡 is the time. The variable 𝑢𝑢�⃗ is the velocity vector, 𝑝𝑝 the pressure, 𝑇𝑇 the temperature and ∝ is the volume fraction of
metal. The one-fluid density is the volume-weighted average 𝜌𝜌 = 𝛼𝛼𝜌𝜌𝑚𝑚 + (1 − 𝛼𝛼)𝜌𝜌𝑔𝑔 , where 𝜌𝜌𝑚𝑚 and 𝜌𝜌𝑔𝑔 are the metal and the
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atmospheric gas density, respectively. The source terms of metal mass, 𝜌𝜌̇ drop , and volume fraction, 𝛼𝛼̇ drop , related to metaltransfer in the form of molten droplets are specified hereafter. 𝑰𝑰 denotes the identity matrix. The one-fluid viscosity μ, specificheat capacity 𝑐𝑐𝑝𝑝 , and thermal conductivity 𝑘𝑘 are volume-weight averaged using the temperature dependent properties ofInvar36 alloy [4]. In the buoyancy force 𝑔𝑔⃗ is the gravitational acceleration, 𝛽𝛽 the coefficient of volume expansion for theliquid alloy [4], and 𝜌𝜌𝑚𝑚 the alloy density at the melting temperature 𝑇𝑇𝑚𝑚 . In the second line of Eq. (2) the first term is theDarcy damping. It is active in the mushy zone where the fraction of liquid alloy, 𝑓𝑓𝑙𝑙 , is strictly bounded between zero andone. This fraction is a function of temperature, as defined in [5]. The Darcy constants are set to 𝐶𝐶𝐷𝐷 = 107 and 𝜀𝜀𝐷𝐷 = 10−3.
The surface forces are estimated using the sharp surface force model introduced by Shams et al. [6]. As the densities of thetwo fluid phases differ by several orders of magnitude, the multiplier 2𝜌𝜌/(𝜌𝜌𝑚𝑚 + 𝜌𝜌𝑔𝑔 ) introduced by Brackbill et al. is applied[7]. The capillary and thermocapillary forces depend on the surface tension, σ, the unit vector normal to the free surface, 𝑛𝑛�⃗,and the surface curvature, 𝜅𝜅. The latent heat of fusion is denoted ℎ𝑠𝑠𝑠𝑠 . In equation (4), the third term is the free-surface
sharpening term, which is a function of the numerical compression velocity, �⃗𝑢𝑢𝑐𝑐 , [8]. The numerical compression factor is set
to 𝐶𝐶α = 1 to satisfy conservation. The effect of the arc is modeled through the arc pressure, 𝑝𝑝arc , the electromagnetic force𝐹𝐹⃗arc , and the heat source, 𝑄𝑄̇arc . These three source terms are respectively estimated using the expression proposed by Tsaiand Eagar [9], by Kou and Sun [10], and𝑄𝑄̇arc =𝜂𝜂arc 𝑉𝑉𝑉𝑉
𝑟𝑟 2
exp
�−
�
2𝜋𝜋𝜎𝜎𝑞𝑞 2
2𝜎𝜎𝑞𝑞 2(5)where 𝑟𝑟 is the radial distance from the center of the arc, 𝑉𝑉 the arc voltage and 𝐼𝐼 the arc current. The arc efficiency, 𝜂𝜂arc , isestimated from the total GMA process efficiency 𝜂𝜂 = 0.8 [11], and the droplet efficiency according to [12]. The same valueof the distribution parameter 𝜎𝜎𝑞𝑞 is assumed for the heat flux, the arc pressure, and the current density according to therelationship proposed in [13]. Based on the experimental measurements, V=25.2 V, I=270 A and 𝜎𝜎𝑞𝑞 =1.4 mm.
Figure 2 shows the computational domain of dimensions 60 mm × 20 mm × 13 mm, which is shorter and narrower thanthe experimental substrate. For the uphill configuration, a thicker atmosphere layer of up to z = 23 mm was used to capture
the alloy deposition (see section 4). The results presented below were computed with a mesh size of 0.2 mm uniform in x, z,
and in -6 ≤ y ≤ 6 mm. A cell-to-cell expansion ratio of 1.2 was applied outside this region. The root pass bead profile scanned
during the experiment served to initialise the first layer in the V-groove, as shown in Fig. 2. For the configurations symmetric
about y=0, only one half of the domain was simulated; this concerns the flat, uphill, and downhill positioning.

Fig. 2: Computational domain after initializing the root pass in the V-groove.

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The test cases were computed with an arc torch initially in x = 10 mm, so that the alloy remained solid on the domainboundaries. The boundary condition for the gas phase was zero velocity gradient if the flow was outward. Otherwise, thecomputed pressure was used to determine the magnitude of the normal velocity at the boundary face. The metal alloy had
zero-velocity and continuous temperature gradient conditions at the boundaries, making the workpiece semi-infinite for heat
conduction. Based on experimental measurements the source terms 𝜌𝜌̇ drop and 𝛼𝛼̇ drop reproduced spherical droplets of radius
0.54 mm transferred from the electrode wire at the periodic frequency of 207 s-1. The location of the arc and droplet injection
was moved at the welding travel speed for a period of 6 s. An adjustable time stepping with maximum CFL number of 0.1,
and maximum time step of 10-5 s was applied. The pressure-velocity coupling was computed with a PISO algorithm. The
convergence criteria imposed on the (final) residuals at each physical time step when solving for 𝛼𝛼, 𝑝𝑝, 𝑢𝑢�⃗ components and𝑇𝑇 was 10-12, 10-10, 10-8, and 10-8, respectively.

4. Results and discussion

Figure 3 presents computed results showing the melt pool followed by the resolidified bead and experimental postprocess bead for the 20° uphill case (different scales). A non-continuous bead with intermittent humps can be seen in theright. The computational results also show an unstable melt pool with humping bead geometry. As the computational domain
was shorter than the experimental substrate, the simulation results did only predict the formation of the first hump. With theother configurations, experimental and computed process and bead surface were all smooth.

Fig. 3: Bead morphology for the uphill deposit condition. Left: simulation. Right: experiment (different scale).

Figure 4 compares, for each configuration, computational results with experimental macrographs in transverse crosssections. The computed interface between fusion zone (orange/red in Fig. 4) and heat affected zone (yellow in Fig. 4) for the
2nd deposit is plotted at the arc axis location (e.g., section S1 in Fig. 3) when the melt pool has reached quasi-steadydimensions. Superposed on it is the computed bead reinforcement (blue grey in Fig. 4) extracted from the resolidified region(e.g., section S2 in Fig. 3 left). With the flat configuration a) the computed and experimental reinforced bead heights are 0.7
and 0.5 mm above the original workpiece height, respectively. A small depression that can lead to an undercut defect isobserved in the metallograph a). To predict it a finer grid size would be required since its length is much lower than 1 mm.
With the downhill configuration b) the height at the centre of the reinforced bead reaches the level of the original workpiecesurface in the experimental metallographs while it is approximately 0.06 mm lower in the simulations. With the uphill
configuration c) and contrary to the other cases, the bead surface is not regular and presents humps as can be seen in Fig. 3.
The macrograph image shown in Fig. 4 c) was obtained from a cross section located at 21 mm from the arc start for bothexperiment and simulation. There, the computed and experimental reinforced bead heights are 0.8 and 0.6 mm above theoriginal workpiece height, respectively. With the side inclined configuration d) the height of the reinforced bead at the centreof the V-groove is at the level of the original workpiece surface in the experimental metallographs. For the numericalsimulation it is 0.55 mm above. Furthermore, the computed bead elevation is slightly lower uphill the groove centre andhigher downhill. The undercut defect where the deposited metal fails to fill the groove can be seen on the bead edge in theexperimental and computational images. Good qualitative agreement is observed in all the computed conditions. The small

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differences compared to the experimental bead elevation might be due to the simplified arc physics, uncertainty in arc
efficiency, or/and to small differences (≈ 2.5%) in average value of current and voltage between the cases that were notconsidered in the simulations.

Fig. 4: Computational (left) and experimental (right) bead cross-sections for a) flat, b) uphill, c) downhill, and d) side inclined position.

Figure 5 visualises for each case at time t = 5.0 s the top view of the fully developed melt pool, its free surface geometry,
the temperature field, and the velocity vectors on the free surface. At the rim of the pool crater (corresponding to the redtemperature area) the velocity vectors are as expected distributed radially outward due to the thermocapillary force. For theflat deposit a) the maximum velocity outside the pool crater, which also results from the thermocapillary flow, is located in
the middle and rear part of the melt where the pool width constricts. A vortex can be seen in the tail end of the melt pool dueto the counterflow returning after striking the rear resolidified front. Similar observations can be made for the side inclined
position d), but the melt flow is then asymmetric, towards the direction of substrate inclination due to the gravity. The meltpool length is approximately the same in a) and d). Figure 5 shows also that the melt pool is the shortest for the downhill
position b) and the longest for the uphill position c). This is due to the gravity force that in the former case supplements thethermocapillary force ahead of the arc to accelerate the flow towards the front part of the melt pool (and decelerate the rearflow). In the latter case it acts on opposite direction, thus accelerates the rear flow (and decelerates the pool front flow) while
surface tension is not sufficient to counteract the resultant increased pressure. This can also be seen with the velocity vectors,e.g., the melt flow shows larger velocity vectors towards the forefront in downhill position b) and towards the pool rear in
uphill position c). Also, the maximum velocity is the lowest in b) and the largest in c). Due to the weaker flow for the
downhill position b) compared to the other cases, no vortex is observed at the rear part of the melt pool in b). Besides, thesechanges in flow velocity compared to the flat case a) induce pressure changes on the free surface that the surface tension
does not sustain. Thus, the pool deformation leading to wider pool front in downhill position b), longer pool rear in uphillposition c), and asymmetric pool edge in side inclined position d) compared to the flat case a). The temperature distributions
within the melt pools are in close ranges, with a maximum in the vicinity of 2920 K for all cases except the downhill position
that is 200 K colder. The melt pool is also wider for the downhill position b) than for the other cases, providing higher degreeof cooling rate. The computed melt pool geometries of Fig. 5 are consistent with the top view images of the pool free surface
behind the arc acquired during the experiments and shown in Fig 6. These images indeed show the shortest melt pool length
for the downhill position b) and an asymmetric pool with more melt towards the downhill direction (indicated by the whitearrow in Fig. 6 d)) of the substrate for the side inclined position. For the uphill position c), the melt pool width is significantly
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narrower, and gaps indicated with arrows can be seen between the melt pool and the groove wall. They contribute to theundercut defects seen in Fig 4.

Fig. 5: Temperature and velocity vectors on the computed melt pool free surface; a) flat, b) downhill, c) uphill, d) inclined case.

Fig. 6: Melt pool top view captured with a vision camera; a) flat, b) downhill, c) uphill, and d) side
inclined (the white arrow points downhill). The red arrows indicate the arc centre travel direction.

Figure 7 shows for each case (at same scale) a longitudinal cross-section passing through the arc axis of the computed
melt pool (grey in Fig. 7) and the velocity vectors at time t = 5.0 s where pool dimensions are quasi-steady. The horizontal
red line indicates the upper surface of the first layer deposit. It should be noticed that the velocity vector plotted at the arc
axis location orients differently depending on the dynamic formation/restoration of the pool crater with drop impact. It cannot
be compared between the instantaneous images a) to d) since although synchronised with the drop detachment frequency the
impact is not exactly at the same time as surface oscillations can differ with the configuration. Two regions recirculating in
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opposite directions can be seen in the melt pool in the flat position a). The recirculation flow that is closest to the arc is
clockwise, therefore it enhances the heat transfer under the arc and the penetration depth. For the other positions, theclockwise recirculation is absent. For the downhill position b) a significantly shallower melt pool can be observed. The
counterclockwise recirculation is then much more extended than in the flat position a) and it flows at a velocity of ≈ 0.3 m/s,
which is about 0.1 m/s larger than in the counterclockwise recirculation of a), indicating a stronger returning flow in theshallower melt pool. For the uphill position c) a significant free surface depression can be seen underneath the arc column
since a large volume of melted alloy flows downward, towards the pool rear under the action of the thermocapillary force
enhanced by the gravitational force. This fluid accumulates in the trailing region, thus the high alloy thickness visible above
the red line indicator in Fig. 7 c), resulting in a hump upon solidification. For the side inclined deposit position d) the meltpool shape at the centerplane is relatively shallow compared to the flat deposit position since some of the liquid metal thenflows towards the downhill edge (see Fig. 5d).

Fig. 7: Melt pool and velocity vectors in longitudinal section through the arc axis for a) flat b) downhill c) uphill d) side position.

4. Conclusion

A thermo-fluid model for metal fusion with GMA heat source with tracking of free surface deformation and metaltransfer developed in OpenFOAM was applied to study the effect of different substrate orientations on metal layer deposition
over a pre-deposited layer in a V-groove. The numerical study was supplemented with experimental observation and the
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results were in good qualitative agreement. However, the undercut defects observed in the experimental metallographs were
only partially captured by the numerical simulations. A much finer grid size would be needed for a precise prediction ofthese small undercut defects. Nevertheless, it was found that deposition of a second layer on a smooth first layer that had
been cooled down to room temperature leads,- for flat welding position, to a liquid alloy recirculation pattern that enhances the penetration depth under the arc, andto smooth and continuous reinforced bead,- for the 20° downhill position, to a large and strong returning flow behind the arc axis towards the melt front, which
increases both the melt pool width and its cooling, while it decreases its length and the bead height.
- for the 20° uphill position, to flow acceleration by gravity towards the pool rear, resulting in a narrowing of the pool
width, increased pool length, and liquid alloy accumulation at the pool rear resulting in undercuts and bead humps
upon solidification.
- for the 20° side inclined position, the trailing part of the melt pool tends to flow downhill resulting in incomplete
fusion on the opposite side of the V groove and undercut defects.
Therefore, uphill and side inclined multilayer GMA deposition are not recommended with these process conditions. Further
investigations are needed to establish the range of inclination angles where acceptable deposition results are maintained.

Acknowledgements

This research work is supported by grants from the EU project - Horizon 2020: INTEGRADDE, which is gratefully
acknowledged. The computations were performed on resources provided by the Swedish National Infrastructure for
Computing (SNIC) at NSC which is gratefully acknowledged.

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