Understanding the thermal process during laser assisted ultra-high frequency induction deposition with wire feeding

International Journal of Heat and Mass Transfer 153 (2020) 119536

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Understanding the thermal process during laser assisted ultra-high frequency induction deposition with wire feeding Rui Sun, Yongjun Shi∗, Xiaogang Wang, Yankuo Guo, Xiaoyu Zhou College of Mechanical & Electronic Engineering, China university of Petroleum, Qingdao, Shangdong 266580, China

a r t i c l e

i n f o

Article history: Received 29 October 2019 Revised 9 January 2020 Accepted 16 February 2020

Keywords: Ultra-high frequency (UHF) induction heat Thermal process Numerical simulation Deposition

a b s t r a c t This study proposes a novel metal deposition method referred to as laser assisted ultra-high frequency induction (UHF) deposition. In this method, the UHF induction heat is used as the main heat source to melt the deposited metal, and the laser heat acts as an auxiliary heat source that provides a high-temperature substrate surface for efficient fusion between the deposited metal and substrate. A numerical model coupled with electromagnetic and temperature fields is developed to understand the thermal process of laser assisted UHF induction deposition. The thermal process with different combination states of the two heat sources is numerically investigated to reveal the influence mechanism of the two heat sources on the penetration depth of the deposited layer. Results show that the UHF induction heat increases the penetration depth of the deposited layer by raising the temperature of the deposited metal, and laser heat leads to an increment in penetration depth by providing a high substrate surface temperature. Decreasing the distance between the laser beam and metal wire also increases the substrate surface temperature, thereby increasing the penetration depth. Criteria for characteristic temperatures Tpeak1 , Tpeak2 , and Tinterval are proposed based on thermal process analysis to control thermal process and prevent the deposited layer from having a large penetration depth. Deposition experiments reveal the process feasibility of the proposed deposition method and validate the numerical model. The performance evaluation of the deposited layers proves that controlling the thermal process is the key for ensuring the performance of the deposited layer. The numerical model and criteria for characteristic temperatures provide an efficient way for controlling the thermal process during deposition; thus, reasonable performance of the deposited layer can be ensured. © 2020 Elsevier Ltd. All rights reserved.

1. Introduction Metal deposition, a promising technique in surface cladding, direct 3D metal part forming, and other additive manufacturing applications, has become a new optional method for the equipment manufacturing industry [1-3]. During deposition, metallic wires or powders are melted by high-energy heat sources and metallurgically bonded to a substrate. Electric arc [4,5] and laser beam [6, 7] are two commonly used high-energy heat sources for metal deposition. The two heat sources are characterized by a high heat input. Although the high heat input improves processing efficiency, it also results in high residual stress and high dilution rate of the deposited layers, which directly lead to poor performance of these layers [8-10]. Extensive studies have been performed on process optimization for metal deposition methods to enhance the performance of deposited metal layers. Researchers have begun to pursue perfor-

∗

Corresponding author.

https://doi.org/10.1016/j.ijheatmasstransfer.2020.119536 0017-9310/© 2020 Elsevier Ltd. All rights reserved.

mance improvement of the deposited metal layer by changing the high-energy heat source type. Heating metal wires and substrates separately with two types of heat sources is an efficient means to achieve a low heat input for substrates. Wen et al. [11] used resistance heat to increase the temperature of metal wires close to the melting point during laser welding; this method decreases the need for high laser power because no extra laser heat is required to melt the wires. Therefore, a low heat input can be achieved. Barroi et al. [12] combined laser heat and non-transferred arc in surface cladding. During cladding, the non-transferred arc is mainly used for melting the metal wire, which significantly reduces the excessive heat input for the substrate. The laser heat is utilized to produce a suitable substrate surface temperature for a sound bond between the coating and substrate. Practically, heat sources that can melt metal materials within seconds should have the potential for metal deposition application. The induction heating method converts electromagnetic energy into heat energy and has been widely used in thermal processing [13,14]. When the frequency of induction power is increased to hundreds of kilohertz, the skin

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depth during induction heating becomes less than 1 mm, and the heating rate of induction power is greatly improved. The increase in power frequency allows the induction heat to realize the rapid melting of metal materials with a small dimension, such as metal wires. Given the precise heating and good controllability of the induction heating method, Vega et al. [15] introduced induction heat into metal deposition. Metal wires with a low melting point are melted by induction heat and dropped from a nozzle. Hascoet et al. [16] applied a high-frequency induction heating method to the deposition of 316 L stainless steel, which has a high melting point. In this approach, the metal wire continuously passes through the induction coil during deposition. Due to the ‘ring effect’ of the induction heating method, the eddy current is mainly concentrated in the wire tip instead of substrate, which thereby decreased the heat input for substrate. With joule heat produced by eddy current, the tip of metal wire is molten and deposited onto the substrate. However, the concentration of eddy current in the metal wire implies the presence of insufficient heat for melting the surface layer of the substrate. With a relatively low temperature of the substrate surface, metallurgical bonding between the deposited layer and the substrate can hardly be obtained because the heat transferred by the molten metal cannot ensure proper melting of the substrate surface. A novel deposition method that combines ultra-high frequency (UHF) induction heat and auxiliary laser heat is proposed in this study to achieve a low heat input for the substrate and metallurgical bonding between the deposited layer and the substrate. This method follows the principle of reducing the heat input by using two types of heat sources to heat the metal wire and substrate separately. In this method, the metal wire is melted and deposited by UHF induction heat, and a low-power laser beam simultaneously irradiates the substrate surface zone located in front of the deposited molten metal. The surface layer of the substrate smoothly fuses with the deposited molten metal due to the laser preheating effect and the thermal conduction of the deposited molten metal. Thus, effective metallurgical bonding between the deposited layer and substrate is obtained. This method can be defined as a compound machining method because of two independent heat sources in the process. Unlike in the laser-induction hybrid deposition method proposed by Zhou et al. [17], the induction heat in the laser assisted UHF induction deposition method serves as a main heat source that melts the depositing material instead of an auxiliary heat source that preheats the substrate. Aside from being a heat source, UHF induction heating exerts another dual effect on the deposition process, namely, the electromagnetic force induced by the alternative magnetic field during UHF induction heating has a stirring effect on the molten metal, which has been proven to be beneficial for grain refinement [18-20]. The induction heating process in the proposed method performs electromagnetic stirring and material melting, a feature that reduces device complexity in comparison with the previous metal deposition method assisted by electromagnetic stirring. This study investigates the proposed laser assisted UHF induction deposition method in terms of process feasibility and thermal process to present a basic understanding of this method. Compared with other metal deposition methods using a single heat source, the laser assisted UHF induction deposition adopts a more complex thermal process for the combination of UHF induction heat and auxiliary laser heat. The thermal process plays a key role in determining the performance of the deposited layer mainly because it controls the penetration depth of the deposited layer and influences the layer’s dilution rate. Therefore, a multi-physics coupling numerical model that involves electromagnetic and temperature fields is established to investigate the thermal process during deposition. The moving mesh approach is used to simulate the relative motion between the metal wire and substrate. The effect of

the combination state of the two heat sources on the thermal process is numerically investigated to reveal the influence mechanism of the two heat sources on the penetration depth of the deposited layer and present criteria for controlling the thermal process. By using a self-designed deposition platform, experiments with different processing parameters are conducted to validate the process feasibility of the proposed deposition method and the established numerical model. The performance of the deposited layers with different processing parameters is also investigated, and the variation in performance is discussed in view of the control of the thermal process. 2. Description of the experimental approach The principle and experimental setup of the laser assisted UHF induction deposition method are illustrated in Fig. 1. During deposition, a metal wire with a diameter of 2 mm is transferred to the induction coil by the wire feeder. With the eddy current generated by the induction coil, the metal wire is rapidly melted and deposited onto the substrate, which is placed on a moving platform. Pure argon as shielding gas is used to protect metal materials from oxidation. Meanwhile, an auxiliary laser beam irradiates the surface area in front of the deposited molten metal. The current that flows through the induction coil is supplied by UHF induction power. According to induction heating theory, the induced eddy current that heats the metal material is mainly concentrated within the skin depth. The skin depth should be smaller than the radius of the metal wire to utilize the eddy current heat. On the basis of the relationship between current frequency and skin depth, the frequency of the UHF induction power is set to 700–1150 kHz. Substrate motion is performed by the moving platform, and a vertical sliding device is used to adjust the distance between the induction coil and substrate. The induction coil is of three turns. The substrate and metal wire applied in the experiment are 316 L stainless steel and Inconel 625 alloy, respectively. The chemical compositions of the substrate and metal wire are listed in Table 1. An infrared thermal imaging camera, Flir A615, is used to record the temperature distribution during deposition, and the measured results are compared with the numerical simulation ones to verify the established numerical model. The measuring range of the infrared thermal imaging camera is 30 0–20 0 0 °C, with a measuring error within ±2%. In the record of the high-temperature distribution, the emissivity of metal materials significantly affects the measurement accuracy. Therefore, for emissivity calibration, the measurements obtained by the infrared thermal imaging camera are compared with the measurements detected by thermocouples. The emissivity is set to 0.6 based on the comparisons and empirical value provided in Flir user’s guide. 3. Numerical model 3.1. Physical model A numerical model coupled with electromagnetic and temperature fields is developed with COMSOL Multiphysics. Fig. 2 shows the geometry and calculation domains of the numerical model. The numerical model is composed of five calculating domains, namely, air, induction coil, metal wire, deposited layer, and substrate. The air domain is further divided into five subdomains (i.e., Air1 –Air5 ) and is supposed to be an auxiliary approach to realize relative motion between the substrate and metal wire. As shown in Fig. 2, the metal wire is included in the numerical model, unlike in previous numerical studies on laser and arc deposition with wires. Nie et al. [21] investigated the temperature variation during laser wire additive manufacturing, and the metal wire

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Fig. 1. Deposition principle and experimental setup of the laser assisted UHF induction deposition method: (a) deposition principle and (b) experimental setup .

Table 1 Chemical composition of 316 L and Inconel 625 alloy (wt%). Materials

Ni

Cr

Si

Mo

Mn

Nb

Co

C

S

P

Fe

316L Inconel 625

10.0–14.0 ≥58

16.0–18.0 20.0–23.0

≤1.0 ≤0.5

2.0–3.0 8.0–10.0

≤2.0 ≤0.5

– 3.15–4.15

– ≤1.0

≤0.03 ≤0.015

≤0.03 ≤0.008

≤0.045 ≤0.024

Bal. Bal.

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Fig. 2. Geometry and domain composition of the numerical model. Table 2 Dimensions of the numerical model. Item

Definition of the item

Value (mm)

l w h h1 h2 hw hcs hd wd dw l1 l2 l3 l4 l5 dco dc

Overall length of the model Overall width of the model Overall height of the model Height of domain Air1 -Air3 Height of domain Substrate Height of domain Metal wire Distance between induction coil and substrate Height of domain Deposited layer Width of domain Deposited layer Diameter of the Metal wire Length of domain Air1 Length of domain Air2 Length of domain Air3 Length of domain Air4 , Air5 Length of domain Substrate and Deposited layer Outer diameter of the induction coil Diameter of the induction coil

65 30 20 16 4 16 5 3 (According to the experimental samples) 4.2 (According to the experimental samples) 2 5 10 50 8 57 1.5 5

was excluded in the numerical model. Bai et al. [22] simulated the temperature distribution during wire arc additive manufacturing on the basis of a numerical model and also ignored the metal wire. An approach that ignores the metal wire in the numerical model is suitable for laser and arc deposition methods. It is mainly because that the laser and electric arc heat source act in forms of heat flux density during numerical calculation, and disregarding the metal wire do not cause significant changes in heat flux density of the two heat sources. However, for an induction heating method, the temperature increment originates from the eddy current excited by the magnetic field. The characteristics of the magnetic field depend on the magnetic properties of the calculation domain; hence, disregarding the metal wire would cause discrepancy in the magnetic field between the numerical simulation and actual experiment. This condition would further cause the distribution of eddy current (i.e., the distribution of the heat source) to be inconsistent with that in actual experimental conditions. Given the consistency between the numerical model and experimental condition, the metal wire is included in the present numerical model. Hascoet et al. [16] simulated the heat transfer in the deposition process with high frequency induction heat, and the metal transfer behavior was simplified as the temperature movement from the metal wire tip to the deposited layer. For calculation simplification, the metal transfer behavior in our numerical simulation is also simplified by the temperature movement. The metal wire in this study is in contact with the deposited layer, as shown in Fig. 2. The total size of the numerical model is 65 × mm 30 × mm 20 × mm, and the specific dimensions of the domains in this

model are listed in Table 2. The cross-section dimensions of the deposited layer are set according to the experimental samples. A triangle mesh is utilized to discretize the numerical model, and the total number of mesh elements is 1.5 × 105 . In consideration of calculation efficiency and accuracy, the meshing size varies in different domains. As shown in Fig. 3, domain Air3 -Air5 are coarsely meshed, while domain Air1 is meshed with a small element size of 1 mm. For the domain Air2, the region that adjacent to the metal wire and Air1 is fine meshed and the rest is coarsely meshed. Domains that undergoes drastic changes in material properties and temperature, including domain Metal wire, Deposited layers and the center part of the domain Substrate are fine meshed with a maximum element size of 0.6 mm. The deposition process involves a significant temperature increment, and the temperature dependency of material properties is considered. The variations in material properties with temperature are illustrated in Fig. 4 [23-25]. 3.2. Implementation of relative motion and real-time addition of deposited material During deposition, a relative motion occurs between the metal wire and substrate; thus, the relative position of the two domains should be updated in real time. Previous studies usually dealt with relative motion between calculation domains by rebuilding the model and changing the material properties in different positions before the next time step [26,27]. Such a method is applicable to numerical models with low computational complexity. For numerical models with high complexity and nonlinearity, such a method leads to huge consumption of computer resources and low calcu-

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Fig. 3. Finite element mesh of the numerical model.

Fig. 4. Material properties of (a) Inconel 625 and (b) 316 L.

lation efficiency for rebuilding the model geometry in every time step. The moving mesh method, which deals with relative motion by deforming the mesh elements instead of rebuilding the model repeatedly, can be utilized to improve the calculation efficiency of numerical models that contain relative motion and high nonlinearity. Artinov et al. [28] used the moving mesh method to address the translational motion of the heat source in laser full-penetration welding, which simplifies the complexity of calculation and can obtain the thermal history of the entire heating process with low time consumption. In this study, the moving mesh method is introduced into the proposed numerical model. The implementation of the moving mesh method and the mesh elements at different times can be seen in Fig. 5. Only half of the established model is shown in Fig. 5 due to symmetry. In Fig. 5a, Air1 –Air3 , metal wire, and induction coil domains are regarded as the moving mesh part in which the mesh elements can move and deform freely, namely, the coordinate of mesh elements in these domains changes with time. During simulation, the united part which is composed of Air2 , metal wire, and induction coil domains moves at the moving velocity of vs , the black arrow denotes the moving direction of

the united part. Due to the motion of the united part, the mesh elements in Air1 and Air3 domains, as shown in Fig. 5b and 5c, are stretched and compressed. The Air4 –Air5 , substrate, and deposited layer domains are taken as the fixed mesh part, in which the coordinate of mesh elements keep unchanged. Continuity condition for electromagnetic and temperature fields exists at the contact interface between the moving and fixed mesh parts. In the actual experiment, the moving parts are the substrate; in the numerical model, the metal wire and induction coil are the moving parts in consideration of modeling simplicity. We maintain the same relative motion velocity between the metal wire and the substrate as that in the actual experiment. The laser heat that acts on the substrate also moves in the same direction and velocity. Therefore, the consistency between the numerical simulation and experiment is not affected. With the relative motion between the metal wire and substrate, the metal wire is melted and deposited onto the substrate. Hence, in the deposited layer domain, the deposited metal wire should be added in real time according to the instantaneous position of the metal wire. At the beginning of the calculation, elements in the

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Fig. 5. Implementation of the moving mesh method: (a) moving and fixed mesh parts of the model, (b) mesh elements at time t = 0 s, and (c) mesh elements at time t = 3 s.

deposited layer are characterized with air properties. As the calculation proceeds, the mesh elements in the deposited layer domain are divided into two types, namely, activated and inactivated, as shown in Fig. 6. Elements in the deposited layer underneath the metal wire are activated along with the relative motion. The electromagnetic and thermal properties of the metal wire are given to these activated elements. The front edge geometry of the activated elements is fitted by a 1/4 ellipsoidal surface based on the actual depositing condition. The expression of the ellipsoidal surface is illustrated in Fig. 6b, where a equals 3 mm, b equals the width of deposited layer wd , and c equals the height of deposited layer hd . The center point P of the ellipsoidal surface is located on the center of the projection of the metal wire on the substrate. The coordinate of point P, marked as (x_focus, y_focus, z_focus), is changed in real time along with the position of the metal wire. 3.3. Governing equations 3.3.1. Electromagnetic field Based on electromagnetic theory, the electromagnetic field during induction heating can be described by Maxwell equations,

which are expressed as follows:

 = J+ ∇ ×H

∂ B ∂t  = ρe ∇ ·D ∇ · B = 0 ∇ × E =

  − → − → ∂D ∂D = Je + Js + ∂t ∂t

(1) (2) (3) (4)

 is the magnetic field intensity, A/m; E is the electric field where H  is electric displacement, C/m2 ; B  is magnetic flux intensity, V/m; D 2  density, T; J is current density, A/m ; ρ e is electric density, C/m3 ; − → − → and Je (A/m2 ) and Js (A/m2 ) are eddy and source current densities, respectively. → , H , D  , E , and − Field variables B Je obey the following equations.

 = μ0 μr H  B   D = ε0 εr E − → Je = σ E

(5) (6) (7)

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Fig. 6. Real-time addition of the deposited layer (a) activated elements and inactivated elements (b) front edge geometry of activated elements in deposited layer.

where μ0 = 4π × 107 H/m, denotes absolute magnetic permeability; μr denotes relative magnetic permeability; ɛ0 is vacuum permittivity which is equal to 8.85 × 10−12 F/m; ɛr is relative permittivity; and σ denotes electrical conductivity, S/m.  and According to Helmholtz theorem, magnetic flux density B electric field intensity E can be further expressed as follows:

  =∇ ×A B ∂ A E = − − ∇ϕ ∂t

(8) (9)

 is the magnetic vector potential in A/m2 and ϕ is the where A scalar potential in V. On the basis of Eqs. (7) and (9), current density J can be described as follows:

− → − → ∂ A J = σ E + Js = −σ − σ ∇ϕ + Js ∂t

(10)

During induction heating, electric displacement density ∂∂Dt is negligible compared with the conduction current. Therefore, the total current law, Eq. (1), can be derived as follows with the as

sistance of Eqs. (5), (8), and (10)

∇×

1

μ0 μr

∇ × A = −σ

− → ∂ A − σ ∇ϕ + Js ∂t

(11)

 and scalar potential ϕ , For the confirmation of vector potential A Eq. (11) is rewritten in a divergence form as follow:



   − → ∂ A  ∇ · −σ + ∇ (−σ ∇ϕ ) + ∇ · Js = ∇ · ∇ × H ∂t

(12)

− → Source current density Js is constant during heating; thus, Eq. (12) can be further described as follows:

∇ 2ϕ +

∂   ∇ ·A =0 ∂t

(13)

In accordance with the uniqueness theorem of vector poten tial, for certain boundary conditions, magnetic vector potential A  are determined is unique only when the divergence and curl of A  = 0 and penalty function uniquely. Hence, the Coulomb gauge ∇ · A ∇ μ1 (∇ · A ) are introduced into Eq. (11), and the following equation is derived.

∇×

1

μ

∇ × A − ∇

 − → ∂ A + σ ∇ϕ = Js ∇ · A + σ ∂t

1

μ

(14)

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The magnetic insulation condition is specified on the boundary of the model, which is depicted as follow:

=0 n×A

(15)

where n denotes the unit normal of the boundary. The continuity condition is applied to the contact interface (Fig. 5a) between moving and fixed mesh domains. Thus, the following equation should be satisfied at the interface.

 dst = A  src A

(16)

where the subscript dst denotes the destination domains (fixed mesh domains) and the subscript src denotes the source domains (moving mesh domains). 3.3.2. Temperature field In the deposition process, the variation process of temperature field is governed by Fourier thermal equation:

ρC p

∂T − ∇ · ( k∇ T ) = Qe ∂t

(17)

where ρ is the material density, Cp is the specific heat, J/(kg•K); k is thermal conductivity, W/(m•K); and Qe is eddy current heat intensity, W/m3 . On the basis of Joule law, Qe is further expressed as: kg/m3 ;

Qe = σ |E |

2

(18)

During deposition, the metal wire continuously passes through the induction coil. Changes in the temperature field resulting from metal wire feeding should thus be considered. The transient thermal process in metal wire and deposited layer domains just underneath the metal wire is as follows:

ρC p

∂T  · ∇ T − ∇ · ( k∇ T ) = Qe + ρC p u ∂t

(19)

 is the metal wire feeding speed in m/s. where u The latent heat of phase change should be considered in the calculation because of solid–liquid phase change during deposition. The latent heat of the phase change is incorporated into the calculation process by the apparent heat capacity method, which can be expressed as follows:

  T −Tm 

Cp

app

exp − = Cp + √

δT

πδ T

Lm

(20)

The apparent heat capacity method considers latent heat by increasing the specific heat capacity in the phase change stage [28,29]. δ T is the half width of the phase change temperature range, and the δ T values of the metal wire and substrate are 30 and 25 K, respectively. Lm is the latent heat of the material during phase change, which is 261 and 280 kJ/kg for the metal wire and substrate, respectively. Tm represents the melting temperature, which is 1593 and 1658 K for the metal wire and substrate, respectively [25]. Cp is the temperature-dependent specific heat capacity of the materials. The specific heat capacity shown in Fig. 4 is the apparent heat capacity of the materials. In the laser-irradiated area, an inward heat flux occurs, namely,

 · k∇ T = Qlaser n

(21)

where Qlaser is the laser heat flux density in W/m2 . The laser heat is considered a Gaussian distribution heat source, that is,

  α Pω −ωr 2 Qlaser = exp π dl2 dl2

(22)

where P is the laser power, W; dl represents the radius of the laser beam, mm; α is the laser absorption coefficient, which equals 0.45; and ω is the energy distribution factor, which equals 2.

In consideration of radiation loss and thermal convection, the boundary condition at the interface between the heated area and air is described as follows:



 · k∇ T = −h(T − Ta ) − εemi σb T 4 − Ta4 n



(23) W/(m2 K);

where h denotes the heat transfer coefficient, Ta is the ambient temperature, K; ɛemi is the radiation coefficient; and σ b represents the Stefan–Boltzmann constant, which equals 5.67 × 10−8W/(m2 K4 ). The continuity condition in the temperature field should also be applied to the contact interface (Fig. 5a) between moving and fixed mesh domains. The governing equations are as follows:

−ndst · qdst = nsrc · qsrc

(24)

Tdst = Tsrc

(25)

At the outermost boundary of the model, the boundary condition is specified as a constant temperature, which is expressed as follows:

T = Tb

(26)

Tb equals 293.15 K. The initial temperature of the model is set to 293.15 K. 3.4. Solution procedures In the process of laser assisted UHF induction deposition, mutual coupling exists between electromagnetic and temperature fields; the eddy current heat induced by the electromagnetic field gives rise to a temperature increment, and the temperature increment leads to changes in electromagnetic and thermal properties, which further affect the distribution characteristics of the electromagnetic and temperature fields at the next moment. The relative motion and real-time addition of deposited materials also make the numerical simulation highly complex. The sequential coupled method is adopted due to the mutual coupling between the multiphysics field and the time variation of the numerical model. The flowchart of the multi-physics coupling calculation process is depicted in Fig. 7. The total deposition time is 14 s (i.e., tend = 14 s). The processing parameters and boundary conditions are specified at the beginning of the calculation. The electromagnetic field calculation is conducted based on these conditions, and the eddy current heat is obtained correspondingly. The eddy current heat and laser heat are then used in the calculation of the temperature field. Afterward, whether current time t equals end time tend is determined. If they are not equal, the position of the metal wire in the next time step is calculated, and the mesh movement is executed. Considering the real-time addition of the deposited material, the mesh elements in the deposited layer domain are activated, and the material properties of the activated elements are changed according to the position of the metal wire. Subsequently, the temperature-dependent material properties, including electrical conductivity, thermal conductivity, and specific heat, are updated based on the temperature field results. Then, the electromagnetic field calculation is restarted, and the process described above is repeated until the end time of the deposition. During calculation, the direct solver PARDISO is used to solve the electromagnetic field results, temperature field results, and the moving mesh method due to its robustness. The time step is controlled by an implicit time-stepping algorithm based on backward differentiation formulas. 3.5. Processing parameters Laser assisted UHF induction deposition uses two independent heat sources. As a result, more processing parameters need to be controlled compared with deposition methods using a single heat

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Fig. 7. Flowchart of the multi-physics coupling calculation process.

Table 3 Processing parameters used in the experiment and numerical simulation.

Fig. 8. Illustration of processing parameters during deposition.

source. The key processing parameters are illustrated in Fig. 8. The specific value of these parameters applied during the experiment and numerical simulation are listed in Table 3. 4. Results and discussion 4.1. Thermal process during deposition The thermal process during deposition involves UHF induction and laser heating processes. For the UHF induction heating process,

Parameters

Illustration

Value

I f P dl dlw vf vs

Current intensity Current frequency Laser power Laser diameter Distance between laser beam and metal wire Feeding velocity of metal wire Moving velocity

30 A 800 kHz 800 W 6 mm 3 mm 410 mm/min 2.4 mm/s

the eddy current mirrors the heat intensity and heat distribution characteristics. Fig. 9 displays the density and distribution of the generated eddy current. Eddy current also exists in the deposited layer adjacent to the coil. Therefore, both the metal wire that travels though the coil center and the deposited metal that is neighboring to the coil can be heated by UHF induction heat. Comparison of the eddy current condition in the metal wire and deposited layer indicates that the eddy current is mainly concentrated in the metal wire tip. The maximum eddy current densities in the metal wire and deposited layer are 2.14 × 108 and 1.07 × 108 A/m2 , respectively. The time-average heat intensity resulting from eddy current is calculated to quantitatively estimate the heating effect of the UHF induction heat on the metal wire and deposited layer. The calculation results show that the time-average heat intensity in the metal wire and deposited layer are 168.4 and 43.3 W, respectively. Hence, the UHF induction heat mainly acts in melting the metal wire tip.

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Fig. 9. Distribution and density of eddy current during deposition (t = 12 s).

Fig. 10 presents the temperature distribution at t = 4, 8, 12 s. The high-temperature zone in the temperature field consists of two parts. The formation of the fore-part high temperature zone, which appears in front of the deposited molten metal, is mainly due to the laser irradiation. The rear-part high temperature zone is formed by the deposited molten metal. As shown in Fig. 10, the average maximum temperature of the metal wire tip during deposition is approximately 2041.1 K. For the fore-part high temperature zone, the average maximum temperature is approximately 2010.9 K. The white dashed line in Fig. 10 marks the scope of the molten metal at different times, the boundary of the molten metal in fore-part and rear-part high temperature zone are determined according to the melting temperature of 316L(substrate) and Inconel 625(metal wire), respectively. On the basis of the geometry of the molten metal during deposition, it can be conclude that under such a processing parameter condition, laser irradiation actually forms a small molten pool in the fore-part high temperature zone. Behind the small molten pool, the metal wire molten by UHF induction heat is deposited onto the substrate. Fig. 11 presents the evolution of the temperature distribution on the substrate surface. The high-temperature area on the substrate surface appears as a “∞” shape. For the high-temperature area on the substrate surface, the front part is the fore-part high temperature zone indicated in Fig. 10, and the rear part is the substrate– deposit interface. The maximum temperature of the two parts is illustrated in Fig. 11. The average maximum temperature in the front and back parts is 2010.9 and 1995.8 K, respectively. During deposition, laser irradiation produces a small molten pool on the substrate surface, which leaves a high substrate surface temperature for the subsequent deposited molten metal. Then, the metal wire that melts through UHF induction heating is deposited onto the substrate. A relatively high substrate surface temperature is fa-

vorable for the metallurgical bonding between the deposited layer and substrate. The thermal cycles of characteristic points P1–P3 are depicted in Fig. 12 to understand the thermal condition evolution at the interface between the deposited layer and substrate. The three characteristic points are located on the center line of the substrate surface, as shown in Fig. 12. All of the thermal cycles contain two temperature peaks. This thermal phenomenon is mainly caused by the sequential heat input of laser heat and UHF induction heat. The thermal cycle of point P2 is used to explain this phenomenon. In the deposition process, point P2 initially undergoes laser irradiation, reaching its first peak temperature (Tpeak1 ) of 2020.8 K at t = 7.4 s. The molten metal has not reached the location of point P2 at this time; consequently, the temperature begins to drop as the laser beam moves forward. Within about 1 s, the temperature decreases to 1643.9 K, which is slightly lower than the solid temperature, Tsolid . Here, the minimum temperature between temperatures Tpeak1 and Tpeak2 is defined as Tinterval , which represents the substrate surface temperature before molten metal is deposited. At this moment, the deposited molten metal reaches point P2 and transfers the UHF induction heat to the substrate. Thus, a temperature increment occurs again. Fig. 12 illustrates that the temperature of point P2 is increased to the second peak temperature (Tpeak2 ) of 1999.5 K at t = 9.4, indicating that the material at P2 turns to molten state once again. The temperature distribution in the cross-section y = 0.015 m is described in Fig. 13, where the boundary of molten pool is marked with a white solid line. Given that we focus on the temperature in the deposited layer and substrate, the temperature of the metal wire is not shown in the figure. As claimed in the former analysis, laser irradiation forms a small molten pool in the fore-part high temperature zone. The molten pool depth in fore-part and rear part high temperature region are obtained by calculating the

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Fig. 10. Temperature distribution at different time. (a) t = 4 s (b) t = 8 s (c) t = 12 s.

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Fig. 11. Evolution of the temperature distribution on the substrate surface: (a) t = 4 s, (b) t = 8 s, and (c) t = 12 s.

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deposition layer, including microhardness and corrosion resistance. For the metal deposition method that adopts electric arc or laser beam as the main heat source, due to the high heat input, excessive melting of the substrate occurs, and the dilution rate of the deposited layer is generally higher than 30% [30,32,33]. To ensure the performance of the deposited layer, dilution rates within 15% are acceptable [34,35]. Therefore, careful control of the penetration depth of the deposited layer is required. In fact, the penetration depth is sensitive to the thermal process during deposition. Consequently, precise control of the thermal process in deposition is essential for avoiding large penetration depth of the deposited layer. The analysis of the thermal process during laser assisted UHF induction deposition implies that the thermal process is determined by laser heat and UHF induction heat. Insights into the thermal process with different combination states of UHF induction heat and laser heat should be acquired to realize precise control of the thermal process in deposition. Fig. 12. Thermal cycles of the characteristic points.

4.2. Thermal process with different combination state of two heat sources

Fig. 13. Temperature distribution and molten pool profile in cross-section y = 0.015 m (t = 8 s).

maximum vertical distance from the bottom of molten pool to the substrate surface. Fig. 13 depicts that the depth of the laser-caused molten pool is 0.43 mm. While in the rear-part high temperature zone, the depth of the molten pool here is 0.56 mm. The molten pool depth in the rear-part high temperature zone is in fact the penetration depth of the deposited layer. The penetration depth indicates the volume of the substrate melted during deposition. In metal deposition technology, the ratio of penetration depth to the sum of the penetration depth and height of the deposited layer is defined as dilution rate D [30,31], that is,

D=

hp × 100% h p + hd

(27)

where hp and hd refer to the penetration depth and height of the deposited layer, respectively, as denoted in Fig. 13. The height of the deposited layer obtained with the processing parameters in Table 3 is 3 mm. Therefore, with the penetration depth of 0.56 mm, the dilution rate of the deposited layer is 15.7%. The increasing dilution rate, namely, the increasing penetration depth, means that a substantial part of the substrate material is melted and mixed with the deposited molten metal. For the metal deposition method used for surface cladding applications, a high dilution rate should be avoided. Excessive mixing of the deposited molten metal and substrate will affect the element composition of the deposited layer, which will further degrade the performance of the

In this section, the effect of the combination state of UHF induction heat and laser heat on thermal process is evaluated from the aspects of molten pool profile and thermal condition at the substrate–deposit interface during deposition. Here, the thermal condition at the substrate–deposit interface is reflected by the thermal cycle of characteristic point P2. In the analysis of the molten pool profile, the variation in penetration depth is predicted according to the molten pool depth in the rear-part high temperature zone. The influencing mechanism of the two heat sources on penetration depth is further revealed based on the thermal condition at the substrate–deposit interface. In accordance with the thermal cycle analysis and influencing mechanism of the two heat sources on penetration depth, the criteria for temperatures Tpeak1 , Tpeak2 , and Tinterval are proposed to prevent large penetration depth and high dilution rate. Among the processing parameters during laser assisted UHF induction deposition, current intensity I, laser power P, and distance between the laser beam and wire dlw are regarded as the most critical parameters that affect the combination state of the two heat sources. Current intensity I and laser power P influence the heat intensity of the two heat sources. An increase in current density indicates an increment in the proportion of UHF induction heat for the combined heat source. A high level of laser power P consequently refers to an increasing laser heat proportion in the combined heat source. The distance between the laser beam and wire, dlw , determines the time intervals between the actions of the two heat sources, which also changes the combination state of the two heat sources. Therefore, a change in the combination state of the two heat sources is implemented by varying these critical parameters. Other processing parameters, such as feeding velocity of the metal wire vf and moving velocity vs , are consistent with the values in Table 3. Parameters vf and vs are unchanged; hence, the variations in the geometry size of deposited layers hd and wd are ignored. 4.2.1. Thermal process with varying UHF induction heat Fig. 14 depicts the variations in the thermal cycle with different current intensities. The laser power and distance between the laser beam and metal wire are set as 800 W and 3 mm, respectively. The current intensity mainly affects temperature Tpeak2 . The appearance of temperature Tpeak2 is caused by the heat transferred by the deposited molten metal; therefore, temperature Tpeak2 reflects the temperature of the deposited molten metal. In the case with a current intensity of 31 A, temperature Tpeak2 is up to 2037.7 K.

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As shown in Fig. 15, in the condition of current intensity of 28 A, the penetration depth is 0.26 mm. The dilution rate in this condition is approximately 7.9%. With a current intensity of 31 A, the penetration depth increases to 0.64 mm, and the dilution rate is correspondingly increased to 17.6%. The thermal cycle in Fig. 14 indicates that the increasing UHF induction heat raises temperature Tpeak2. An increase in temperature Tpeak2 means an increase in the heat transferred from the deposited molten metal to the substrate. Therefore, a substantial part of the substrate material is melted. It can be concluded that the UHF induction heat determines the penetration depth by improving the temperature of the deposited molten metal. As shown in Fig. 14, temperature Tpeak2 at the current intensity of 30 A is 1999.5 K. The penetration depth and dilution rate are 0.56 mm and 15.7%, respectively, according to Fig. 15. Temperature Tpeak2 should not exceed 20 0 0 K to avoid excessive melting of the substrate and limit the dilution rate of the deposited layer within ~15%.

Fig. 14. Thermal cycles with varying current intensity I.

When the current intensity is set to 28 A, which refers to a combination state with low UHF induction heat, temperature Tpeak2 is decreased to 1799.6 K. With an increase in UHF induction heat, the metal wire is melted and deposited at a high temperature. Temperature Tpeak2 is consequently improved. Temperature Tpeak1 shows no changes due to the constant laser power. Temperature Tinterval is also slightly changed by current intensity, which is improved from 1581.6 K to 1668.0 K when the current intensity increases from 28 A to 31 A.

4.2.2. Thermal process with varying laser heat Thermal cycles with different laser power are shown in Fig. 16. The current intensity and distance between the laser and metal wire are set as 30 A and 3 mm, respectively. A variation in laser heat mainly affects temperature Tpeak1 . As indicated in Fig. 16, temperature Tpeak1 at laser power of 700 W is 1831.6 K. For the heat source combination state with high-intensity laser heat, namely, a laser power of 900 W, temperature Tpeak1 is as high as 2201.9 K. An increment in temperature Tpeak1 results in increased temperature Tinterval . When the combined heat source contains high-intensity laser heat, the substrate surface is in a high-temperature condition when the molten metal is deposited. Although temperature Tinterval is high under high-intensity laser

Fig. 15. Molten pool profile and penetration depth with varying current intensity I: (a) 28 A, (b) 29 A, (c) 30 A, and (d) 31 A.

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Fig. 16. Thermal cycle with varying laser power P.

heat, temperature Tpeak2 does not consequently increase because the UHF induction heat is constant. The increasing laser heat raises temperature Tpeak1 . Increasing temperature Tpeak1 in turn leads to an increment in temperature Tinterval , namely, it provides a high substrate surface temperature. Thus, the surface layer of the substrate can be melted with relatively less heat. Therefore the penetration depth is consequently increased. The temperature Tinterval with laser power of 900 W is as high as 1728.9 K, it can be inferred that the substrate surface material is in molten state when the molten metal wire is deposited. As a result, the penetration depth in this condition, as shown in Fig. 17, is up to 0.62 mm, leading to a high dilution rate of 17.1%. Therefore, to prevent the large penetration depth, there should be a limitation for the temperature Tpeak1 . The penetration depth and dilution rate under the laser power of 800 W are decreased to 0.56 mm and 15.7%, respectively, and Tpeak1 in this condition is 2020.8 K. Hence, Tpeak1 should be lower than 20 0 0 K to achieve a dilution rate within the acceptable level of ~15%. 4.2.3. Thermal process with varying distance between laser beam and metal wire Fig. 18 shows the thermal cycle with varying distance between the laser beam and wire dlw . The laser power and current density are set to 800 W and 30 A, respectively. Changing dlw mainly causes a variation in temperature Tinterval . Temperature Tinterval in

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Fig. 18. Thermal cycle with varying distance between laser beam and wire dlw .

the case of parameter dlw of 2 mm is 1872.8 K. With increasing parameter dlw , the distance between the UHF induction heat and laser heat is increased, indicating a long time interval between the actions of the two heat sources for a certain area. The temperature of this area keeps decreasing during the period. As parameter dlw increases to 5 mm, temperature Tinterval drops to 1286.9 K. Temperature Tpeak1 decreases by approximately 60 K as dlw increases from 2 mm to 5 mm, although the laser power keeps constant. In the case of dlw of 2 mm, the distance between the laser beam and metal wire is relatively small. When laser beam irradiates the area of point P2, not only the laser heat but also the thermal conduction of the rear deposited molten metal contribute to the temperature increment of P2. As the distance between the laser beam and metal wire increases, the contribution of thermal conduction of the rear deposited molten metal is weakened, and temperature Tpeak1 decreases slightly. Temperature Tpeak2 here does not vary with the parameter dlw . Although the UHF induction heat and laser heat are constant, the penetration depth can differ with varying distance between the laser beam and metal wire, as shown in Fig. 19. The penetration depth shows an increasing tendency as dlw decreases from 5 mm to 2 mm. The aforementioned analysis clarifies that the decreasing dlw increases temperature Tinterval , which provides a high substrate surface temperature. This phenomenon is similar to the condition

Fig. 17. Molten pool profile and penetration depth with varying laser power P: (a) 700 W, (b) 800 W, and (c) 900 W.

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Fig. 19. Penetration depth and molten pool profile with varying distance between laser beam and metal wire dlw : (a) 2 mm, (b) 3 mm, (c) 4 mm, and (d) 5 mm. Table 4 Critical parameters used in the experiments. Cases

Current intensity I (A)

Laser power P (W)

Distance between laser beam and metal wire dlw (mm)

1 2 3 4 5

30 28 30 30 30

800 800 800 900 800

3 3 4 3 2

when applying high laser heat, which means that the melting of the substrate during deposition requires relatively less heat. Consequently, dlw mainly increases the penetration depth by increasing Tinterval . Tinterval under a dlw of 2 mm is as high as 1872.8 K; therefore, the penetration depth is increased to 0.61 mm, leading to a dilution rate of 16.9%. In the condition of dlw of 3 mm, which results in a penetration depth of 0.56 mm and a dilution rate of 15.7%, Tinterval is 1643.9 K. Given the control of dilution rate, Tinterval must be restricted within 1600 K to prevent a large penetration depth. 4.3. Experiment validation and performance investigation Deposition experiments with varying critical parameters are conducted based on the established experimental platform. The critical process parameters are shown in Table 4. Other parameters are consistent with those in Table 3. The experimental samples are transversely cross-sectioned by a wire-cutting electrical discharge machine. The cross section of the experimental samples is sanded with sandpaper and then finally polished. Afterwards, the cross section of the experimental samples is etched with aqua re-

gia for 90 s. The cross section of the experimental samples is then observed with a scanning electron microscope (SEM), and the element distribution is detected with an energy-dispersive spectroscope (EDS). Fig. 20a–e presents the cross section of the experimental samples. It can be seen that there is an obvious boundary between deposited layer and substrate. This boundary is in fact a very thin transition region between deposited layer and substrate, as denoted in Fig. 20c. No cracks origin from the transition region in all the five experimental samples. Thus, the deposited layers have efficiently bonded with the substrate, indicating the feasibility of the laser assisted UHF induction deposition method. Fig. 21a shows the microstructure of the transition region in case 3. The element distribution along lines A–A’ is detected and shown in Fig. 21b. It can be seen that element diffusion occurs in the transition region, meaning that the deposited layer and substrate are also metallurgically bonded. The penetration depth shows significant changes with the variation in parameters. The penetration depth of the experimental samples is measured by SEM and compared with the numerical simulations, as described in Fig. 22. Former analysis of the thermal process have shown that there are improper thermal process un-

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Fig. 20. Cross-section of the experimental samples: (a) case 1, (b) case 2, (c) case 3, (d) case 4, (e) case 5. .

Fig. 21. The transition region in case 3: (a) the microstructure and (b) element distribution along A-A’.

der the processing parameters of cases 4 and 5, the characteristic temperature in the two condition exceed the proposed criteria for temperature Tpeak1 and Tinterval , respectively. Therefore it can be predicted that the experimental samples should be of large penetration depth in the two conditions. Experimental results shows that the penetration depth of case 4 and case 5 experimental samples are up to 0.717 and 0.619 mm, respectively, and the dilution rate are 19.2% and 17.1%, respectively. The case 2 sample, which is obtained in the condition of low current intensity, has a penetration depth of only 0.277 mm. The penetration depth of case 1 and 3 are 0.492 and 0.45 mm, respectively. The dilution rate of experi-

mental case 1–3 are all below 15%. The thermal process under the process parameters of case 1–3, as simulated in Section 4.2, satisfies the requirement for the characteristic temperature, as a result, the acceptable dilution rate is obtained. The penetration depth in experiments shows good agreement with that in the numerical simulation, verifying the numerical model. Fig. 23 presents the temperature distribution in the experiment and numerical simulation under the parameter condition of case 1 in Table 4. Observation of the experiment record shows a hightemperature zone in front of the deposited molten metal, as the white solid line points out in Fig. 23. This finding is basically con-

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Fig. 22. Comparison of the penetration depth in experiments and numerical simulations.

sistent with the numerical simulation. A detecting line is defined on the surface of the deposited layer to make a quantitative comparison of the temperature in the experiment and numerical simulation, as shown in Fig. 23. In the experiment record, measurement points N1–N9 with equal intervals in the detecting line are selected. Fig. 24 shows the temperature of the measurement points in the experiment and the temperature distribution of the detecting line in the numerical simulation. The numerical simulation agrees with the experimental measurements, thereby also validating the established numerical model. As mentioned above, penetration depth is determined by the thermal process. The criteria for the characteristic temperature Tpeak1 , Tpeak2 , and Tinterval in fact provides the reference for controlling the thermal process during deposition. Through controlling the thermal process, penetration depth of the deposited layer is limited, ensuring the performance of deposited layer. In order to further verifying the significance of controlling the thermal process

Fig. 24. Temperature of the measurement points in the experiment and temperature distribution along the detecting line in the numerical simulation.

for ensuring the performance of deposited layer, corrosion resistance and microhardness of the experimental samples are tested. The corrosion resistance of the experimental samples is evaluated through a corrosion test, which is performed with the potentiodynamic polarization method. The corrosion test is conducted in a three-electrode cell that uses a platinum plate as the counter electrode, an Ag/AgCl electrode as the reference electrode, and the deposited layer of the experimental samples as the working electrode. During the test, the three electrodes are immersed in a typical 3.5% NaCl electrolyte solution at room temperature. With potentiodynamic polarization equipment, the open-circuit potential (OCP) of the deposited layer is tested for 5400 s. Then, the potentiodynamic polarization curve is plotted via potentiodynamic polarization scanning testing. The scanning parameters are set as follows: scanning rate of 0.33 mV/s and scanning range of −200 to 1200 mV with respect to the OCP of the deposited layer.

Fig. 23. Comparison of temperature distribution in experiment and numerical simulation (t = 12 s).

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Fig. 25. Potentiodynamic polarization curve of the experimental samples. Table 5 Corrosion test data of the experimental samples in 3.5% NaCl solution at room temperature. Case

Ecorr (mV)

Icorr (mA•cm−2 )

1 2 3 4 5

−258.6 −153.8 −249.2 −323.2 −290.3

1.36 7.03 1.17 4.29 1.87

× × × × ×

10−4 10−5 10−4 10−3 10−4

Eb (mV) 495.2 638.6 587.3 354.2 162.4

Fig. 25 shows the potentiodynamic polarization curves of the experimental samples. The corrosion test data of the experimental samples, including corrosion potential (Ecorr ), corrosion current density (Icorr ), and breakdown potential (Eb ), are summarized in Table 5 in accordance with the potentiodynamic polarization curves. A high corrosion potential or a small corrosion current density generally indicates high corrosion resistance of the testing materials. The corrosion test data listed in Table 5 shows that the case 2 sample exhibits the highest corrosion potential (−153.8 mV) and the smallest corrosion current density (7.03 × 10−5 mA/cm2 ), demonstrating high corrosion resistance. The case 1 sample has a corrosion potential of −258.6 mV and a corrosion current density of 1.36 × 10−4 mA/cm2 , which denote a slightly degraded corrosion resistance compared with the case 2 sample. Comparing the corrosion potential and corrosion current density of case 3 sample with case 1 sample shows that the corrosion resistance of case 2 sample is slightly higher than that of case 1 sample. The typical corrosion potential and corrosion current density of wrought Inconel 625 alloy are −264 mV and 4.7 × 10−4 mA/cm2 , respectively [10,36]. Therefore the corrosion resistance of case 1–3 are better than typical wrought Inconel 625 alloy. The experimental samples of cases 4 and 5 exhibit worse corrosion resistance in comparison with case1-3 samples as well as the typical wrought Inconel 625 alloy. As listed in Table 5, the corrosion potentials of the two samples are reduced to −323.2 and −290.3 mV, whereas the corrosion current densities are increased to 4.29 × 10−3 and 1.87 × 10−4 mA/cm2 . In addition to corrosion potential and corrosion current density, breakdown potential is also an indicator for the corrosion resistance of materials. In Fig. 25, each curve has a region where the potential shows an obvious increase with a slight change in current density. This phenomenon indicates that a dense passive film is formed on the surface of the samples, and

19

Fig. 26. Microhardness of the experimental samples.

this film further protects the sample from corrosion. However, with the continuous increase in potential, pitting corrosion occurs, and the passive film is destroyed, which is manifested as a rapid increase in current density with a small variation in potential in the polarization curve. The potential in this condition is defined as the breakdown potential (Eb ). The higher the breakdown potential is, the better the resistance is for the pitting corrosion. Table 5 indicates that case 1–3 samples all have a breakdown potential higher than or close to 500 mV. However, the breakdown potential for the case 4 sample is decreased to 354.2 mV, and case 5 only has 162.4 mV. The corrosion test shows that the case1-3 samples exhibit satisfying corrosion resistance which is better than the typical wrought Inconel 625 alloy. Whereas, the case 4 and 5 samples have the worst corrosion resistance. The experimental samples of cases 4 and 5 are of high penetration depth because there are improper thermal process during deposition. Therefore, the corrosion resistance of the deposited layers is severely degraded compared with the typical wrought Inconel 625 alloy. While the case 1–3 sample, with the controlled thermal process, the corrosion resistance is significantly better than case 4 and 5. The microhardness of the experimental samples is measured by a Vickers hardness tester with a load of 0.3 kg and a dwelling time of 15 s. Microhardness is measured along the depth in the cross section of the experimental samples with a spacing of 0.2 mm between measuring points, as denoted in Fig. 26. The deposited layer of the case 2 sample has an average microhardness of 225.6 HV, and the maximum value is up to 240 HV. The average microhardness values of the deposited layer in cases 1 and 2 are 218.1 and 218.6 HV, respectively. For the case 4 and 5 samples, which are of the high penetration depth due to the improper thermal process in deposition, the average microhardness values of the deposited layer are decreased to 207.4 and 210.1 HV, respectively. The experiments and performance evaluation of the deposited layers prove that improper thermal process in deposition is the critical reason for large penetration depth and performance degradation. Controlling the thermal process is the key for ensuring the performance of the deposited layer. The established numerical model of laser assisted UHF induction deposition provides an efficient tool for understanding the thermal process, and the criteria of temperature Tpeak1 , Tpeak2 and Tinterval in fact provide reference

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for controlling the thermal process. Thus, reasonable performance of the deposited layer can be ensured.

5. Conclusions In this study, a novel deposition method that combines UHF induction heat and laser heat is proposed. This method is numerically and experimentally investigated, and the main conclusions are as follows: 1) A numerical model coupled with electromagnetic and temperature fields is presented to investigate the thermal process during laser assisted UHF induction deposition. Investigations show that eddy current exists in the metal wire and deposited layer adjacent to the coil. The high-temperature zone during deposition has two parts, namely, the fore-part high temperature zone formed by laser heat and the rear-part high temperature zone formed by the deposited molten metal. 2) The influencing mechanism of the two heat sources on the penetration depth of the deposited layer is revealed according to the thermal process under different combination states of UHF induction heat and laser heat. During deposition, UHF induction heat increases the penetration depth by improving the temperature of the deposited molten metal, whereas laser heat raises the penetration depth by providing a high substrate surface temperature for the deposited molten metal. Decreasing the distance between UHF induction heat and laser heat also increases the substrate surface temperature, resulting in an increasing penetration depth of the deposited layer. To avoid a large penetration depth and keep the dilution rate within acceptable range, the criteria for characteristic temperatures Tpeak1 , Tpeak2 , and Tinterval are proposed based on the thermal process analysis as follows: Tpeak1 and Tpeak2 should not exceed 20 0 0 K, and Tinterval should not exceed 1600 K. 3) Laser assisted UHF induction deposition experiments are conducted. Samples with efficient bonding between the deposited layer and substrate are obtained, indicating the engineering feasibility of this method. Comparison of penetration depth and temperature distribution shows good agreement between the experiments and numerical simulations. The performance of the deposited layers, including corrosion resistance and microhardness, are evaluated. The results reveal that improper thermal process in deposition accounts for large penetration depth and performance degradation of the deposited layer. With controlled thermal process, the performance of deposited layer can be ensured. The numerical model of laser-assisted UHF deposition and the criteria for characteristic temperatures provide an efficient way in controlling the thermal process, thereby ensuring the performance of the deposited layer. This paper presents a basic understanding of laser assisted UHF induction deposition. An investigation of the fluid flow of molten metal during deposition is not included in this work. The magnetic force in the molten metal induced by the alternating magnetic field should not be ignored because UHF induction heating is involved in the process. In future work, a numerical model will be coupled with the fluid flow field, and an investigation of fluid flow behavior during laser assisted UHF induction deposition will be conducted.

Declaration of Competing Interests The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

CRediT authorship contribution statement Rui Sun: Conceptualization, Methodology, Software, Formal analysis, Investigation, Writing - original draft, Writing - review & editing. Yongjun Shi: Conceptualization, Writing - review & editing, Project administration, Supervision, Funding acquisition. Xiaogang Wang: Validation, Resources, Investigation. Yankuo Guo: Validation, Resources, Investigation. Xiaoyu Zhou: Validation, Software, Investigation.

Acknowledgments This research was financially supported by PetroChina Innovation Foundation (Grant 2017D-5007-0307), Postgraduate Innovation Engineering Project of China University of Petroleum (East China) (Grant YCX2019052), and Fundamental Research Funds for Central Universities, China (Grant 18CX05004A).)

Supplementary materials Supplementary material associated with this article can be found, in the online version, at doi:10.1016/j.ijheatmasstransfer. 2020.119536.

References [1] M.K. Imran, S.H. Masood, M. Brandt, et al., Thermal fatigue behavior of direct metal deposited H13 tool steel coating on copper alloy substrate, Surf. Coat. Technol. 206 (8–9) (2012) 2572–2580. [2] L. Thivillon, P. Bertrand, B. Laget, et al., Potential of direct metal deposition technology for manufacturing thick functionally graded coatings and parts for reactors components, J. Nucl. Mater. 385 (2) (2009) 236–241. [3] G.P. Dinda, L. Song, J. Mazumder, Fabrication of Ti-6Al-4V Scaffolds by direct metal deposition, Metall. Mater. Trans. A 39 (12) (2008) 2914–2922. [4] H.B. Geng, J.L. Li, J.T. Xiong, et al., Optimization of wire feed for GTAW based additive manufacturing„ J. Mater. Process. Technol. 243 (2017) 40–47. [5] F. Hejripour, F. Binesh, M. Hebel, et al., Thermal modeling and characterization of wire arc additive manufactured duplex stainless steel, J. Mater. Process. Technol. 272 (2019) 58–71. [6] X. Xu, G.Y. Mi, L.D. Xiong, et al., Morphologies, microstructures and properties of TIC particle reinforced inconel 625 coatings obtained by laser cladding with wire, J. Alloy. Compd. 740 (2018) 16–27. [7] S.P. Wei, G. Wang, Y.C. Shin, et al., Comprehensive modeling of transport phenomena in laser hot-wire deposition process, Int. J. Heat Mass Transf. 125 (2018) 1356–1368. [8] A. Kagawa, Y. Ohta, K. Nakayama, Mechanism of crack generation in carbide surface layer of laser clad iron alloy, Mater. Trans. 43 (6) (2002) 1261–1265. [9] L.F.S. Vieira, H.J.C. Voorwald, M.O.H. Cioffi, Fatigue performance of AISI 4340 steel Ni–Cr–b-Si–Fe HVOF thermal spray coated thermal spray coated, Proc. Eng. 114 (2015) 606–612. [10] T.E. Abioye, D.G. Mccartney, A.T. Clare, Laser cladding of inconel 625 wire for corrosion protection, J. Mater. Process. Technol. 217 (2014) 232–240. [11] P. Wen, Z.H. Feng, S.Q. Zheng, Formation quality optimization of laser hot wire cladding for repairing martensite precipitation hardening stainless steel, Opt. Laser Technol. 65 (2015) 180–188. [12] A. Barroi, F. Zimmermann, J. Hermsdorf, Evaluation of the laser assisted double wire with nontransferred arc surfacing process for cladding, J. Laser Appl. 28 (2) (2016) 022306. [13] D.I. Luozzo, M. Fontana, B. Arcondo, Modelling of induction heating of carbon steel tubes: mathematical analysis, numerical simulation and validation, J. Alloy. Compd. 536 (4) (2012) S564–S568. [14] Y. Han, E.L. Yu, T.X. Zhao, Three-dimensional analysis of medium-frequency induction heating of steel pipes subject to motion factor, Int. J. Heat Mass Transf. 101 (2016) 452–460. [15] E. Vega, M. Cabezas, B.N. Muñoz-Sánchez, et al., A novel technique to produce metallic microdrops for additive manufacturing, Int. J. Adv. Manuf. Technol. 70 (5–8) (2014) 1395–1402. [16] J.Y. Hascoet, P. Jérome, P. Mognol, et al., Induction heating in a wire additive manufacturing approach, Weld. World 62 (2) (2018) 249–257. [17] S.F. Zhou, Y.J. Huang, X.Y. Zeng, A study of ni-based wc composite coatings by laser induction hybrid rapid cladding with elliptical spot, Appl. Surf. Sci. 254 (10) (2008) 3110–3119. [18] X.M. Zhou, Q.H. Tian, Y.X. Du, et al., Study of the influence of longitudinal static magnetic field on surface quality and performances of arc welding based additive forming parts, Chin. J. Mech. Eng-EN. 45 (2) (2018) 84–92.

R. Sun, Y. Shi and X. Wang et al. / International Journal of Heat and Mass Transfer 153 (2020) 119536 [19] Y.C. Lim, X. Yu, J.H. Cho, et al., Effect of magnetic stirring on grain structure refinement: part 1 – Autogenous nickel alloy welds, 15 (7) (2010) 583–589 [20] Y.C. Lim, X. Yu, J.H. Cho, et al., Effect of magnetic stirring on grain structure refinement part 2 – Nickel alloy weld overlays, Sci. Technol. Weld. Join. 15 (5) (2010) 400–406. [21] Z.G. Nie, G. Wang, J.D. Mcguffin-Cawley, Experimental study and modeling of H13 steel deposition using laser hot-wire additive manufacturing, J. Mater. Process. Technol. 235 (2016) 171–186. [22] X.W. Bai, H.O. Zhang, G.L. Wang, Modeling of the moving induction heating used as secondary heat source in weld-based additive manufacturing, Int. J. Adv. Manuf. Technol. 77 (1–4) (2015) 717–727. [23] R. Sun, Y.J. Shi, Z.F. Bing, et al., Metal transfer and thermal characteristics in drop-on-demand deposition using ultra-high frequency induction heating technology, Appl. Therm. Eng. 149 (2019) 731–744. [24] P.Z. Ma, Y. Wu, P.J. Zhang, et al., Solidification prediction of laser cladding 316L by the finite element simulation, Int. J. Adv. Manuf. Technol. 103 (2019) 957–969. [25] K.C. Mills, Recommended Values of Thermophysical Properties For Selected Commercial Alloys, 1st ed, Cambrige: Woodhead Publishing, 2002. [26] X.B. Zhang, C. Chen, Y.J. Liu, Numerical analysis and experimental research of triangle induction heating of the rolled plate, Proc. Inst. Mech. Eng. C.- J. Mec. 231 (5) (2017) 844–859. [27] V. Fireteanu, P. Roehr, Finite element analysis of the transverse flux induction heating of moving magnetic steel sheets, The 8th International Symposium on Advanced Topics in Electrical Engineering, 2013.

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[28] A. Artinov, M. Bachmann, M. Rethmeier, Equivalent heat source approach in a 3D transient heat transfer simulation of full-penetration high power laser beam welding of thick metal plates, Int. J. Heat Mass Transf. 122 (2018) 1003–1013. [29] H. Hu, S.A. Argyropoulos, Mathematical modelling of solidification and melting: a review, Model. Simul. Mater. SC. 4 (4) (1999) 371–396. [30] J.C. Heigel, M.F. Gouge, P. Michaleris, et al., Selection of powder or wire feedstock material for the laser cladding of inconel@ 625, J. Mater. Process. Technol. 231 (2016) 357–365. [31] Y.X. Li, P.F. Zhang, P.K. Bai, et al., Analysis of geometrical characteristics and properties of laser cladding 85wt% Ti + 15wt% TiBcN powder on 7075 aluminum alloy substrate, Materials (Basel) 11 (9) (2018) 1551. [32] E. Capello, B. Previtali, The influence of operator skills, process parameters and materials on clad shape in repair using laser cladding by wire, J. Mater. Process. Technol. 174 (1–3) (2006) 223–232. [33] S.S. Sandhu, A.S. Shahi, Metallurgical, wear and fatigue performance of inconel 625 weld claddings, J. Mater. Process. Technol. 233 (2016) 1–8. [34] T.E. Abioye, J. Folkes, A.T. Clare, A parametric study of inconel 625 wire laser deposition, J. Mater. Process. Technol. 213 (12) (2013) 2145–2151. [35] S. Liu, W. Liu, M. Harooni, et al., Real-time monitoring of laser hot-wire cladding of inconel 625, Opt. Laser. Technol. 62 (10) (2014) 124–134. [36] J. Tuominen, P. Vuoristo, T. Mantyla, et al., Improving corrosion properties of high-velocity oxy-fuel sprayed inconel 625 by using a high-power continuous wave neodymium-doped yttrium aluminum garnet laser, J. Therm. Spray. Technol. 9 (4) (20 0 0) 513–519.