Reinforcement of Tooling Using Residual Stresses Generated by Cladding by Arc Welding

Abstract

Cladding is typically used to protect components from wear and corrosion while also improving the aesthetic value and reliability of the substrate. The cladding process induces significant residual stresses due to the temperature difference between the substrate and the clad layer. However, these residual stresses could be effectively utilized by modifying processes and geometrical parameters. This paper introduces a novel methodology for using the weld-cladding process as a cost-effective alternative to various existing reinforcement techniques. The numerical analyses are performed to maximize the reinforcement of a cylindrical tool. The investigation of how the weld cladding develops compressive stresses on the specimen in response to a change in the weld beads and the welding sequence is presented. For the benchmark shape, experimental verification of the numerical model is performed. The influence of the distance between the weld beads and the effect of the tool diameter is numerically investigated. Furthermore, the variation in compressive stresses due to temperature fluctuations during the extrusion process has been evaluated. The results showed that adequate compressive stresses are generated on the welded parts through the cladding process after cooling. More compressive stresses are induced in the tool as the cross-section of the weld bead is increased. Furthermore, keeping a gap between the adjacent beads improves tool reinforcement. Hence, the targeted reinforcement of the substrate can be achieved by optimizing the welding sequence and process parameters.

1. Introduction

Reinforcements of metal components are used in various applications, e.g., increasing the strength of forming tools in regions where high loads need to be carried, reducing the degradation of components, enhancing the fatigue life, and minimizing distortions [1]. An extrusion tool is a typical application that requires reinforcements, as shown in Figure 1. The reinforcement enhances the tool life and erosion behavior of the tool and reduces elastic deformation.

1.1. Extrusion Tool

Extrusion processes are known for their high dimensional accuracy and good surface quality [2]. The increased production rate makes this process cost-effective [3]. In metal extrusion, the forming forces are applied to one end of the tool, while the formed component is obtained either at the opposite end (forward extrusion) or the same end (reverse extrusion) [4]. In another form of extrusion known as double cup extrusion, the billet is squeezed into the desired shape between the two punches [5]. The movable upper punch is pushed towards the fixed lower punch, and the pressure is applied to the billet, resulting in the free flow of the material around the two punches [5,6]. A double cup extrusion model is shown in Figure 1.

Many Additive Manufacturing (AM) techniques are associated with the extrusion process, such as liquid metal AM, in which the material is melted before extrusion and solidified after cooling [7]. According to ASTM F42 [8], Fused Deposition Modeling (FDM) is also known as material extrusion AM. Nowadays, AM technology is also employed for extrusion-based printing methods, such as Selective Laser Melting (SLM), to produce extrusion tools of diverse materials that would otherwise be difficult to manufacture due to high operating temperatures [9]. As extrusion involves very high pressure, it causes the formation of high tensile stresses in the tool, which can lead to crack formation [10,11]. Tensile stresses develop on the inner side of the tool and limit its load capacity. A reinforcement ring (see Figure 1) is usually used around the tool to increase its strength. The reinforcement ring compensates for tensile stresses and provides the tool with additional strength [12]. Lange [13] reported that using a steel reinforcement ring around the extrusion tool raises the permissible internal pressure of the tool to 1390 MPa. Thus, by adding a single reinforcement ring around the tool, the internal pressure-bearing capacity of the tool can be increased up to 80% relative to the tool without a reinforcement ring. Klocke et al. [12] stated that extrusion dies without reinforcement rings lose 37.5% of their strength during extrusion. Placing two reinforcement rings next to each other enhances the internal pressure bearing capacity of the extrusion tool by 27% [12]. Therefore, to increase the strength of the extrusion tools, either the undesired tensile stresses should be minimized, or reinforcement should be applied [11].

1.2. Reinforcement of Toolings

Numerous metal components are reinforced by applying external pressure on them by various techniques. The press-fit or thermal shrink-fit method has been used to generate compressive stresses on the cylindrical components to reinforce them [14].

A commonly used industrial approach for reinforcing a cylindrical rod using the shrinkage of a hot ring is presented in Figure 2. In this process, the metal ring is heated to a defined temperature, such as 1000 °C [15], followed by mounting on the cylindrical rod and left to cool. The heating causes the thermal expansion of the metal ring and its diameter increases. When the ring cools off, the thermal contraction of the metal develops compressive stresses on the rod due to the close fit between the rod and the ring. These compressive stresses increase the strength and enhance the fatigue life of the metal rod [16].

The method mentioned above for reinforcing the components is very effective and cheap, but it has some disadvantages. Since it requires a tight tolerance between the shaft and the bore (interference fit P7/h6 [17] according to ANSI B4.2 [18]), positioning of the hot ring must be done with advanced tools, such as hydraulic presses, mounting levers and mounting clamps [19].

Although shrink fitting is used to enhance the fatigue life of the components [20], the most significant disadvantage of shrink fitting is that it can only be used on cylindrical components [21]. On the other hand, the hot ring can damage the rod if it dislocates from its axis or if the compressive stresses applied to the ring increase the tensile strength [22]. This method could also significantly change the stress and strain profile and generate considerable tensile stresses on the surface [23]. It may initiate cracks that propagate and eventually lead to specimen failure [14,24]. Shrink-fit requires exact fittings between the adjacent surfaces, leading to complications when dealing with larger components [20].

In addition, to shrink fitting, other techniques, such as thermal autofrettage, mechanical autofrettage, explosives, and part rotation can reinforce the components. The autofrettage process generates a very high pressure within the tube cavity, inducing compressive residual stresses after the pressure is released [25]. According to Thumser et al. [26], the thermal autofrettage of a round, symmetrical part increases its endurance limit and reduces the amount of tensile residual stresses, particularly from the internal surface. Jahromi et al. [27] claimed that the autofrettage of a metal tube with reinforced ceramic particles on its inner surface increases the amount of compressive residual stress. Shufen et al. [28] proposed that locally heating a 1080-steel tube to an elevated temperature using thermal autofrettage followed by its cooling induces compressive stresses at its inner surface. They reported that the development of the compressive residual stresses during cooling increases the fatigue performance of the tube. Kamal et al. [29] imposed high pressure on the inner surface of an SS304 steel tube using thermal autofrettage, which after cooling, developed compressive residual stresses on its surface. They found that the load-carrying capacity and the fatigue strength of the tube enhanced due to compressive stresses. Besides thermal autofrettage, Malik et el. [30] used a mechanical autofrettage method to compress a thick steel cylinder using a conical mandrel. This method develops compressive stresses in the cylinder and strengthens it. Kaplan et al. [31] used an explosive method to reinforce the thick-walled forging dies made of steel and aluminium. They used high-pressure explosive energy in a controlled way to establish plastic flow on the surface of the die-cavity. With the reduction of the pressure from the die, high residual stresses are generated. Zare et al. [32] utilized a rotational method to strengthen a thick-walled cylinder. They rotated the cylinder with a high angular velocity, which developed compressive residual stresses and resulted in its reinforcement. Although the procedure mentioned above offers reinforcement, they primarily create compressive stresses on the internal cavity of the cylindrical specimen. They may also significantly increase the harmful tensile stresses on the cylinder’s exterior surface, leading to cracking [33].

Some other conventional methods to strengthen the components include grain boundary strengthening [34], solid solution strengthening [35], heat treatment [36], precipitation hardening [37], and cold working [38]. Apart from these conventional surface hardening techniques, weld cladding, in which a metal is welded/coated with a layer of another or the same metal, is also an effective method to minimize the tensile residual stresses, hence, improving the mechanical properties of components (up to 3 mm or more) [39]. In the past, various metals have been used as a coating/clad to achieve the desired properties, i.e., nickel, steel alloys, copper alloys, manganese alloys, and composites [40]. Since it is known that welding causes residual stresses, various researchers have sought to minimize the stresses by weld cladding. Pan et al. [41] coated the high strength steel grade 16 Mn on the surface of steel grade 9Cr to enhance the hardness and reduce the residual stresses and corresponding wear and crack propagation. Jiang et al. [42] reported that high thermal loads during cladding create tensile stresses, which might cause cracks and weaken the material. They suggested increasing the number of clad layers on the metal surface to reduce tensile residual stresses. However, they did not mention the optimized number of layers and their pattern. Benghalia et al. [43] used the weld-cladding process on a cylindrical substrate to strengthen it. They revealed that weld-cladding develops more compressive stresses on the substrate than the typical shrink-fit joint. The compression reduces the pre-existing tensile residual stresses on the substrate and enhances the compressive stresses, increasing its strength.

Numerical techniques are becoming an essential tool for estimating temperature, distortion, and residual stresses caused by the welding processes. Several simulation techniques have been presented to compute the thermal behavior during the cladding process in recent years. Weld-cladding simulations are typically performed by thermo-mechanical simulation methods, which predict the global thermo-mechanical behavior of the weld. Incorporating thermal material parameters into the model at the melting point and room temperature simplifies this behavior [44]. Lou et al. [45] developed a three-dimensional finite element model to examine the influence of a heat source on a heat-affected zone. They also observed the effect of the cooling rate on the cladding layer. Gao et al. [46] developed a 3D thermal FEM (Finite Element Method) model to simulate the temperature field in the Fe-based weld cladding. They evaluated different process parameters, i.e., temperature distributions, cooling rate, and solidification rate. Kumar et al. [47] developed a 3D model to estimate various physical phenomena, i.e., heat transfer, fluid flow, and convection during the cladding process. They observed that changing the scanning speed changes the area of the weld pool and the degree of convection. Hosseini et al. [48] revealed that heat input, welding sequence, and substrate geometry influence the thermal cycle and stress state in the wire arc welding processes. Smith et al. [49] evaluated the welding of a single clad layer using more than 40 finite element simulations to assess the influence of different process variables on the expected residual stresses. They observed that the hardening modulus of the material has a significant influence on the developed stresses. However, mesh size, welding efficiency, and thermal boundary conditions are less influential. Karlsson et al. [50] and Lindgren [51] analyzed the effect of different variables involved in the simulations on residual stresses, heat input, and deformation. They found that the material’s heat capacity is one of the parameters that could affect the temperature distribution and resulting stresses. Feroozmehr et al. [52] utilized a finite element approach to investigate the influence of weld deposition patterns on the temperature and stress distribution during the cladding process. However, the weld bead dimensions and the deposition pattern that could generate the maximum compressive stresses on the specimen have not been explained.

From state of the art, it can be concluded that various methods have been used to enhance the strength of the components by cladding. However, no previous study could be found on weld cladding induced targeted residual stresses in a tool. Moreover, there is no information on how multiple clad layers should be welded to maximize the compressive stresses on the component’s surface, where operating stresses are usually the highest. Topics requiring further investigation include: the effect of the number of clad layers, the number of beads in a layer, the size of the beads, the distance between the different layers, and the effect of cavities on the developed stresses. In the present study, weld cladding is introduced to induce compressive stresses in a cylindrical tool in a controlled manner.

1.3. Scope and Structure

The main goals of the present research are:

• To utilize stresses of the weld-cladding process for reinforcing the cylindrical tool.
• To adopt the weld-cladding pattern to maximize the compressive stresses on a cylindrical tool surface.
• To find how the weld-bead size affects the compressive stresses on the tool.
• To analyze the effect of changing the tool’s materials and process parameters on the maximum achievable compressive stresses.

The paper is structured as follows. Section 2 explains the materials and methods. In Section 2.1, the simulation model is introduced. Section 2.4 presents the experimental procedure, i.e., specimen production, welding preparation, and residual-stress measurement technique. In Section 3, the results of a single bead numerical model are validated with the experiments to prove the accuracy of the numerical model, followed by the detailed evaluation of different modeling strategies and process parameters using the numerical analysis. Section 4 comprises a discussion section to correlate the findings of this investigation with existing literature. Finally, Section 5 reports the summary and conclusions drawn from this work.

2. Materials and Methods

2.1. Numerical Procedure

A three-dimensional Finite Element (FE) model of the demonstrator shape was developed to study the thermal history and residual stresses and optimize the welding parameters and trajectories for single and multiple weld beads. A fully coupled thermo-mechanical implicit model having hexahedral solid elements with one-point integration and viscous hourglass control was used for the structural analysis. All simulations were performed with a double-precision solver (ls-dyna_smp_d_R12.0) in LS-DYNA (Ansys, Canonsburg, PA, USA) according to the investigations of [53].

A tied (mechanical) contact was established between the beads from the start of the simulation. While, the time-dependent thermal contacts were enabled between the neighboring beads at a pre-defined time. The boundary conditions for convection and radiation on the tool and beads were linked with the initiation of thermal contact between the connected surfaces. Heat loss via convection and radiation begins from the exposed surfaces of the active beads.

2.1.1. Element Activation Technique

The beads (filler metal) deposition on the substrate (tool) was carried out using ghost material properties, i.e., the quiet element technique. In this technique, all the elements of weld beads are kept quiet/inactivated at first. The inactive elements of the beads remain in the model, but they do not affect the active beads. To prevent structural deformation below the melting point, inactive or dead elements are modeled with a deficient (factor of 0.1 or less of the properties at room temperature [54]) Poisson’s ratio, young’s modulus, and thermal expansion coefficient. At the same time, a very high value of heat capacity is set. The thermal material properties are activated first to ensure the addition of the material in structural analysis. The heat source activates the thermal properties of elements at a proper time step within a temperature range of 1200 °C–1210 °C. In contrast, their mechanical properties are activated above the melting temperature, in the range of 1400 °C to 1450 °C for steel. A temperature range was used to activate both thermal and structural models to avoid convergence issues during the simulation.

The effect of multiple beads in single and multiple layers is investigated. Different weld-bead materials and their influence on generating the compressive stresses are tested using a single bead model. After validating the simulation results with experiments, various welding paths and bead sizes were modeled and analyzed to understand the development of compressive stresses on the tool’s surface after cooling.

2.1.2. Material Model

The thermal analysis was carried out using the material model *MAT_Thermal_CWM T07. It is a material definition with temperature-dependent properties in which the materials are created or activated depending on temperature. The thermo-elastic plastic material *MAT_CWM 270 with kinematic hardening and the von Mises yield criterion were used for the structural analysis. To consider the effects of recrystallization, creep deformation, and transformation plasticity in the simulation, the general phase change model should be used [55]. However, because we address the global behavior of the macro (Type I) residual stresses, these mechanisms are regarded beyond the scope of the present work. The thermo-mechanical material properties of steel grade 309L correspond to the steel used in the experiments. Details of the material parameters and properties of steel 309L are shown in Figure 3.

2.1.3. Heat Source Model

The basic concept of a heat source is to model an isothermal surface of the weld pool, where the heat flow is quite near to reality [56]. In this work, Goldak’s double ellipsoid heat source model is used for thermal analysis, as presented by Goldak et al. [57]. The following equations determine the power density distribution in the front and rear regions:

$${\mathrm{q}}_{\mathrm{f}}\left(\mathrm{x},\mathrm{y},\mathrm{z},\mathrm{t}\right)=\frac{6\sqrt{3}{\mathrm{f}}_{\mathrm{f}}\mathrm{Q}}{\mathsf{\pi }\sqrt{\mathsf{\pi }}{\mathrm{abc}}_{\mathrm{f}}}{\mathrm{e}}^{-3{\left(\frac{\mathrm{x}}{\mathrm{a}}\right)}^2}{ \mathrm{e}}^{-3{\left(\frac{\mathrm{y}}{\mathrm{b}}\right)}^2}{\mathrm{e}}^{-3{\left(\frac{\mathrm{z}+\mathrm{v}\left(\mathsf{\tau }-\mathrm{t}\right)}{{\mathrm{c}}_{\mathrm{f}}}\right)}^2}$$

(1)

$${\mathrm{q}}_{\mathrm{r}}\left(\mathrm{x},\mathrm{y},\mathrm{z},\mathrm{t}\right)=\frac{6\sqrt{3}{\mathrm{f}}_{\mathrm{r}}\mathrm{Q}}{\mathsf{\pi }\sqrt{\mathsf{\pi }}{\mathrm{abc}}_{\mathrm{r}}}{\mathrm{e}}^{-3{\left(\frac{\mathrm{x}}{\mathrm{a}}\right)}^2}{ \mathrm{e}}^{-3{\left(\frac{\mathrm{y}}{\mathrm{b}}\right)}^2}{\mathrm{e}}^{-3{\left(\frac{\mathrm{z}+\mathrm{v}\left(\mathsf{\tau }-\mathrm{t}\right)}{{\mathrm{c}}_{\mathrm{r}}}\right)}^2}$$

(2)

The sum of the forward and rear heat fractions should equal 2.

$${\mathrm{f}}_{\mathrm{f}}+{\mathrm{f}}_{\mathrm{r}}=2$$

(3)

The total heat introduced in the model by the heat source and the total heat loss due to conduction, convection, and radiation is given by:

$$\begin{gathered} {\mathrm{q}}_{\mathrm{Total}}={ \mathrm{q}}_{\mathrm{input}}-{\mathrm{q}}_{\mathrm{lost}} \\ {\mathrm{q}}_{\mathrm{lost}}={ \mathrm{q}}_{\mathrm{conduction}}+{\mathrm{q}}_{\mathrm{convection}}+{\mathrm{q}}_{\mathrm{radiation}} \end{gathered}$$

(4)

Table 1. Goldak’s ellipsoidal heat source parameters for different dimension beads.

Parameter	Symbol	A1	A2	A3	A4
Weld pool width	a	2.25 mm	2.5 mm	3.0 mm	3.75 mm
Weld pool depth	b	3.0 mm	3.3 mm	4.0 mm	5.0 mm
Weld pool (forward direction)	cf	1.5 mm	1.6 mm	2.0 mm	2.5 mm
Weld pool (rear direction)	cr	6.0 mm	6.4 mm	8.0 mm	10.0 mm
Weld pool energy	Q	1840 W	2248 W	3271 W	5111 W

presents the various parameters of Goldak’s heat source used in this study. The parameters n = 1, v = 7.5 mm/s, and f = 1 remain constant for all the models, while the remaining heat source parameters are changed according to the bead shapes, as explained in previous work [53,58,59,60].

The welding parameters for all simulations were kept the same. The weld pool energy was adjusted according to the weld-bead geometry by maintaining a constant welding-energy to bead area ratio.

2.1.4. Meshing and Clamping

The tool was fixed at its bottom (represented by the yellow marks in Figure 4). Only the lower nodes were fixed set, allowing the tool to contract and expand freely under the thermal loads, as in the experiments. The welding time for a single bead on the tool was approximately 46 s, with a welding velocity of 7.5 mm/s. A complete welding cycle was divided into 90-time steps of 0.5 s each. Considering higher temperatures and heat flows along the z-axis, the mesh size in the welding zone was set to 0.5 mm, whereas a coarser mesh size of 2 mm was used outside the welding zone. The mesh size in the circumferential direction was kept constant at 10.45 mm, as shown in Figure 4.

2.2. Finite Element Model (FEM) of a Single-Bead Cladding

The finite element model of a single bead of various cross-sections was constructed on a cylindrical tool (Figure 4). The beads were classified as A1–A4 according to their cross-sections, as shown in Figure 5.

Here, l and d are the length and the diameter of the tool, d_i and d_o are the inner and outer diameters of the beads, w is the width, and h is the height of the beads.

2.2.1. Weld-Bead Position on a Tool with an Internal Cavity

Five numerical models with different bead positions were developed to investigate the influence of an internal feature, i.e., a cavity, in developing compressive stresses on the tool surface (see Figure 6).

The hole diameter at different positions can be seen in Figure 6. The dimensions of the bead and the tool remained constant in all models.

2.2.2. Welding on the Tools of Different Diameters

Six different tool diameters, namely 40 mm, 45 mm, 50 mm, 55 mm, 60 mm, and 65 mm, were modeled to analyze the influence of the tool diameter on the reinforcement. The dimensions of tools and beads can be seen in Figure 7.

The beads were designed with the same inner diameter as the tool’s outer diameter, while the length and thickness of the beads were kept constant in all models.

2.3. Finite Element Model (FEM) for Multiple-Bead Cladding

2.3.1. Single-Layer Parallel Cladding of Multiple Beads with and without a Gap

Models of two and five parallel weld beads in a single layer, with and without a gap, were built to assess whether weld beads parallel to each other influence the compressive stresses on the tool after cooling. Furthermore, the influence of introducing a gap between the beads and the resulting compressive stresses was also evaluated. The placement of 2 and 5 weld beads on the tool is depicted in Figure 8a–c. The welding is performed from the bottom to the top.

The numbers on the left side of Figure 8c represent the welding sequence, i.e., 1 to 5. The symbol l₁ describes the distance between the tool’s base and the edge of the first bead. The remaining symbols are already explained in Figure 5.

2.3.2. Multilayer Parallel Welding of Beads

The dimensions and positioning of the weld beads on the tool in multiple layered parallel welding are shown in Figure 9.

The numbers on the left side of Figure 9c represent the welding sequence in three layers, i.e., 1 to 6. In Figure 9d, d_i represents the smallest diameter of the beads, while d_o denotes the diameter of the outer bead.

Various process parameters were used to categorize the global temperature development and the local temperature history of multilayer welding patterns. According to Rehman et al. [61], time delay has a significant impact on the mechanical characteristics of a material. The time delay between the two consecutive weld beads might influence the global temperature development. Therefore, the effect of introducing a time delay in single and multiple layers was compared to models without a time delay. It is known that the temperature peaks will change when beads are welded in different trajectories [62]. Therefore, the welding direction was reversed to see how the resultant compressive stresses on the tool changes. Figure 10 shows the multilayer cladding models with different trajectories and process parameters, i.e., welding in the same direction and the opposite direction.

Dotted inner circles in Figure 10c,d represent the beads of the first layer, while the solid circles represent the beads of the second and third layers. The numbers above the arrows in Figure 10c,d show the welding sequence, i.e., the two parallel beads were welded in the first layer, followed by the welding of the next two beads in the second layer.

Various welding trajectories of the 2 × 2 pattern can be seen in Figure 11.

The red and blue circles indicate the welding path, while the arrows represent the welding direction of the beads. Inner dotted circles represent the beads of the first layer, while the solid outer circles depict the beads of the second layer. The welding sequence of the beads in the two layers is indicated by the numbers above the circles and the path under each trajectory. For example, in trajectory T1, the first bead was welded in the first layer, followed by the second bead in the first layer. Then the third and fourth beads are welded in the second layer, and so on.

2.4. Experimental Procedure for Single-Bead Cladding

A solid cylinder having a length of 100 mm and a diameter of 100 mm was defined as a sample geometry. Since the experimental investigations were performed for validating the single-bead cladding simulation, the structural steel S-235JR was chosen as a tool material because of its good weldability and low cost. The experimental procedure was divided into two sections: the specimen’s production and the measurement of the residual stresses.

2.4.1. Specimen Production

The tests were carried out using a six-axis FANUC robotic arm with the parameters listed in

Table 2. Welding parameters.

Parameter	Value
Welding current [I]	141 A
Welding voltage [U]	13.2 V
Torch travel speed [TS]	7.5 mm/s
Wire feed [WF]	123 mm/s

.

Solid welding wire ER70S-6 of 1.0 mm diameter was employed as a filler metal. The welding wire ER70S-6 has been used due to its superior welding qualities. It contains significant proportions of manganese and silicon, resulting in high-quality welds even on unclean, greasy, or oxidized surfaces. The silicon also enhances the fluidity of the molten weld pool, resulting in a smoother bead appearance and, therefore, less post-weld finishing [63].

Table 3. Chemical composition of S235 and ER70S-6.

Material	Element (Wt %)
	C	Mn	Si	P	S	Cr + Mo + Ni
ER70 S-6	0.06–0.15	1.4–1.85	0.8–1.15	0.025	0.035 max	<0.63
S235	0.17 max	1.40 max	-	0.045 max	0.045 max	-

shows the chemical composition of the filler wire and the substrate according to American welding standards AWS [64] and Deutsches Institut für Normung DIN EN 10025-2 [65]. A shielding gas composed of 80% Argon and 20% CO₂ was used at a flow rate of 15.0 l/min to avoid oxidation and contamination of the weld pool [66].

The initial temperature of the specimen was 25 °C. The specimen’s movement was limited during the welding process by clamping it at eight different positions, as shown in Figure 12.

Beads were welded on the substrate using the parameters described in Table 2. During the welding process, the robotic arm with the attached welding torch positions vertical to the substrate. The machine-table starts rotating in a pre-defined path following an NC-code developed in Matlab and Rhino-3D, as explained by Nguyen et al. [67]. The welding time for a single bead was approximately 46 s, after which the tool was left to cool for the next 300 s without changing its position. A constant emissivity value of 0.9 was chosen [68]. During welding and cooling processes, the temperature was captured using a thermal infrared camera (VarioCAM, Infratec, Dresden, Germany) having an image resolution of 1.024 × 768 pixels at 1 frame/s. The height and thickness of the beads were 7.5 mm and 5 mm.

2.4.2. Residual Stress Measurement

In this work, the hole-drilling method, a semi-destructive method for measuring residual stress, was used to validate the simulation results. The hole-drilling method is fast, inexpensive, and can be used for a variety of materials. Although the hole drilling method does not always generate accurate results, the purpose of using this method is to obtain a rough estimate of the stress distribution by having some qualitative statements without a detailed analysis.

During the drilling procedure, few residual stresses are relieved, and comparable strains on the surface of the specimen are evaluated using the strain gauges. The three strain gauges (also known as rosette strain gauges) are attached to the surface of the specimen to determine the strains around the drilled holes (see Figure 13a). Based on the strains, residual stresses are calculated using the calibration constants. For the stress analysis, a disc-shaped specimen was cut-off from the middle of the weld bead along section A-A. A hole-drilling method along with the position P of the hole in the specimen is shown in Figure 13.

3. Results and Discussions

3.1. Single Bead Weld-Cladding Simulation

The heat source melts the metal wire during the welding process and builds up a bead. The molten metal causes thermal expansion of the material on the surface of the substrate. On cooling, the shrinkage of the metal develops stress on the weld bead and the substrate material near the weld, as shown in Figure 14. After cooling, the thermal expansion creates more visible stresses on the tool’s surface. Increasing compressive stresses on the tool surface are governed by the thermal loads and the shrinkage of the weld beads during solidification. Temperature and radial stress development in the numerical model of the demonstrator is shown in Figure 14.

The temperature development can be seen along the cross-section in Figure 14a, while Figure 14b,c shows the radial stresses in the tool. The thermal loads without any influence of the bead are demonstrated in Figure 14b. Figure 14c shows both thermal and mechanical stresses from the weld bead. Most of the stresses in the tool arise from thermal loads during welding, as can be seen in Figure 14b. During welding, at time t = 3.5 s (start of welding), compressive residual stresses are generated in the heat-affected area of the tool, surrounded by tensile residual stresses, as can be seen in Figure 14b,c. When the heat source moves away, cooling begins on the tool’s surface. Due to the shrinkage of the weld bead, the compressive stresses are generated on the outer surface of the tool (see Figure 14c). Due to the thermal shrinkage of the material at t = 53 s (at the end of the welding phase), the tensile stresses in the middle of the tool decrease due to global cooling, and the compressive stresses dominate (see Figure 14b,c) After the sample has cooled to room temperature (at 118 s), the compressive residual stresses dominate, except in the start/stop region. The mechanical contact of the beads with the tool generated approximately 40–60% of the internal stresses, while the thermal loads caused the rest. Therefore, it can be concluded that the combined compressive stresses of thermal loads and the shrinkage of the weld bead build up compressive stresses and strengthen the tool.

The mean surface temperature and the equivalent stress for an element (having a size of 10.45 mm × 0.5 mm) are shown in Figure 15. As the heat source moves away from the reference element, the liquid metal solidifies, causing thermal shrinkage, visible in the sloping curve. After cooling to around 70 °C, the stress curve reveals a constant value of approximately 160 MPa on the tool surface.

3.2. Experimental Validation of Single Bead Simulation

The temperature distribution was recorded, and the outcomes were compared at different instances, i.e., at the beginning of the welding (0°), after welding half bead (180°), and after the complete bead (360°), as shown in Figure 16.

Results indicated that the temperature development in both the experiment and the simulation followed the same trend. The heat distribution pattern on the tool’s surface can also estimate the welding path. The thermal history of two points at different locations on the tool has been compared, as shown in Figure 17. Temperature peaks indicate the moment when the heat source moves over the element. As point P1 was close to the welding bead, a higher peak was detected than P2, relatively far from the welding bead.

After cooling, a 7 mm disk-shaped specimen was cut out of the tool with a metal-cutting saw and finished for stress measurement using a hole-drilling machine PROTEC (Protec GmbH, Ottobrunn, Germany). A hole diameter of 1.7 mm and a hole depth of 0.6 mm were achieved during the drilling, as detailed in Section 2.3.2. The residual stresses along the radial and theta directions were measured at point P of the specimen (see Figure 18).

A comparison of experimental and simulation results for radial and theta stresses along point P is presented in Figure 19.

The radial stresses σ_r in the tool along point P were mainly compressive, while the stresses in theta direction σ_θ were tensile (see Figure 19). After the tool cooled, more compressive stresses were observed, which caused the reinforcement of the tool.

For further analysis, only the simulation of the various geometric models will be performed, and results will be presented. In the post-simulation analysis of LS-Dyna, the pressure exerted on the tool’s surface by the molten metal is assumed to be equivalent to mean stress [69], so in the following text, the mean stress σ_m on a particular element of the tool will approximate the reinforcement. After the part has cooled to room temperature, the stress history of the specified elements will be compared to analyze the results.

3.3. Optimization of a Single-Layer, Multiple-Beads Cladding

More than one bead is welded parallel to each other in a layer to observe their influence on each other and the resulting stresses they induce on the tool surface. It should be noted that the multiple beads are welded continuously without giving a pause time to cool the previous bead to room temperature. The effect of adding pause time between the beads will be presented in the following section.

3.3.1. Influence of the Numbers of Beads and Their Dimensions

The number of beads and their cross-section could affect the stress distribution on the tool surface, as also demonstrated by Jousten et al. [70]. Figure 20 compares the stress histories of 2-beads and 5-beads models of different dimensions.

The position of the referenced element on the tool surface is highlighted in yellow. The FEA models have been developed for two and five parallel weld beads of various sizes, namely A1, A2, A3, and A4 (see Figure 5). Here the welding direction is from the bottom to the top, i.e., the bead at the bottom is welded first, followed by the bead at the top. It can be seen that as the heat source gets closer to the referred point, which is on the top bead, the material expands, and tension develops on the elements underneath the beads. As the referenced element lies under the top bead, a higher stress peak was observed when the heat source passed over it, as shown in Figure 20. The beads solidified and shrank as the temperature dropped, revealing the compressive stresses on the tool surface.

The compressive stresses experienced by the tool also vary with the size of the beads. Smaller beads, such as A1 and A2, induced less compressive stresses on the tool. The compressive stresses on the tool enhanced as the beads became larger. More heat flow between the beads and the tool has been observed in bigger beads A3 and A4. As a result of the increased heat transfer, more thermal expansion and shrinkage were counteracted by the colder material around it, resulting in more compressive stresses, which produced more compressive stresses on the tool [71].

The difference in the stress histories of the 2 and 5 beads models can also be seen in Figure 20. Compared to the 2-beads model, the 5-beads model introduces more heat into the model, resulting in an overall rise in temperature. Due to conduction, heat from the first few beads was dissipated in the tool, and its temperature was increased. As the tool was reheated with multiple beads, the overall temperature of the model began to rise, and the only medium of cooling was direct cooling by air. The reheating caused the 5-bead model to cool slowly, and the beads induced less compressive stresses on the tool than 2-beads. Therefore, for the same-sized beads, the compressive stresses induced on the tool surface by the 2-beads model were greater than that caused by the 5-beads model. As a result, the tool with fewer beads appeared more robust than the tool with more beads, as also stated by Hu et al. [72] and Ganguly et al. [73]. In the case of steel 309L, the maximum compressive stress of σ_m = −128 MPa was generated by 2-beads with dimensions A4.

3.3.2. Influence of a Gap between the Beads

As previously noted, most of the compressive stresses on the tool surface were located directly underneath the weld beads. The number of beads in weld cladding also affects the compressive stresses induced on the tool. The maximum compressive stresses on the tool decrease as the number of parallel beads increase (see Section 3.3.1). In the multiple beads model, the reheating limits the maximum achievable compressive stresses of the individual bead. The reheating causes the conversion of the compressive stresses back into tensile stresses. With the 100 mm diameter tool and beads of dimension A3 (w = 6 mm, h = 4 mm), the smallest or critical gap necessary to minimize the reheating effect in multiple parallel beads is determined in Figure 21.

If there is no or a small gap (up to 11 mm) between the two parallel beads, the maximum compressive stresses generated on the tool surface by the first bead are less than that of the second bead. As the distance between the two parallel beads increases, the compressive stress on the tool by the first bead increases. After a critical gap of 11 mm between the adjacent beads, their effect on each other becomes almost negligible. The maximum compressive stress generated by both the beads becomes the same. The maximum compressive stress on the tool remains unchanged as the distance between the beads is increased further.

If the distance between the two beads is above 15 mm, there is no influence of the temperature on the compressive stress generated by each of the beads. It can be inferred that the minimum gap between the two parallel weld beads should be equal to the critical gap of 11 mm to maximize the resultant compressive stresses on the tool surface.

3.3.3. Influence of the Internal Hole in the Tool

In this sub-section, the tools with a tapered hole along the length of the cylinder have been considered. The stiffness of the tool is weakened in the cross-section as the hole becomes larger. Thus, with a large diameter hole, the tool can deform elastically, and the reinforcement effect due to shrinkage of the weld beads is reduced. Five models were simulated with the beads welded on the tool at different positions (see Figure 6 and Figure 22a). It is to be noted here that only one bead is welded at one of the five positions in one model. The maximum compressive stress produced by a bead welded at various positions on the tool surface is depicted in Figure 22b. In contrast, the stresses inside the tool are shown in Figure 23.

At position 1, where the tool has an inner hole diameter of 57.7 mm underneath the bead, developed the maximum compressive stress of σ_{m, max} = 162 MPa. The maximum compressive stress increased from σ_{m, max} = 182.45 MPa to σ_{m, max} = 230.47 MPa as the hole diameter reduced, i.e., from position 2 to position 4. The maximum compressive stress has been achieved on the tool with a bead at position 5 and accounts for approximately σ_{m, max} = 239.1 MPa. Therefore, it can be concluded that the hole diameter and the maximum compressive on the tool surface are inversely proportional to each other. The bead’s position on the tool without a hole underneath resulted in more compressive stress; therefore, more reinforcement can be achieved there on the tool after cooling.

In each of the models with different bead positions, the stress distribution in the tool after cooling and unclamping can be seen in Figure 23. Analogous to the increase in stiffness in the tool, an increase in compressive stress is also visible in the cross-section of the tool underneath the bead. The difference in the stress distribution on the left and right sides of the part is just due to the start and endpoint of the trajectory. The compressive region was smaller in models 1 and 2 because the hole diameter under the bead was bigger. As the hole diameter decreased, more compressive stress was induced on the tool surface (under the bead), and more tensile stress was produced within the tool to balance it, as shown in models 3 and 4. The bead position in model 5 caused more compressive stress on the tool surface, which is visible as a wider blue region on the surface.

The beads welded at positions 1 and 2 resulted in rapid cooling through conduction and convection, resulting in a smaller tensile region in the middle of the tool. As the hole diameter reduced, cooling became slower, and a larger tensile area emerged inside the tool, as shown in models 3 and 4. In the case of model 5, most of the heat exchange was carried out by convection with the external surface of the tool, as there was no hole under the bead. Therefore, a wider compressive region appeared along the surface.

3.3.4. Influence of a Time Delay between the Weld-Beads

The addition of pause time or delay between the weld beads can influence the compressive stress state on the tool surface after cooling. A time delay of 20 s was kept before welding the next bead until all the beads were welded. A comparison of the stress histories of the tool elements in two parallel-bead models with and without time delay is shown in Figure 24. The addition of a time delay of up to 20 s in the 2-beads model has not shown any noticeable difference. However, the delay resulted in relatively less compressive stresses on the tool surface in the 5-beads model. Compared to the model without delay, the tool cooled more and was reheated to a lower peak during the welding of the next bead. Hence, the beads in the model, without delay, on cooling, shrunk more and generated more compressive stresses on the tool. The model cooled more quickly as a result of the delay. Hence, the heat did not accumulate, and comparatively less shrinkage was observed on cooling, as also described by Lanxess [71]. The continuous welding of the 5-beads without delay induced approximately 5–15 MPa more compressive stresses on the referenced element than the model with delay and minimized substrate distortion, as reported by Mughal et al. [74]. The delay between the weld beads only increases production time and decreases the reinforcement. Therefore, fewer beads and welding strategies without time delay are preferred.

3.3.5. Influence of the Welding Direction

The direction of the adjacent bead was reversed to investigate the influence of welding direction on induced compressive stresses, as shown in Figure 25. The thermo-mechanical history of the model with the reversed welding direction (R) was found to be different from that of the model with the same-direction welding. However, because input energy and welding time were the same in both models, comparable compressive stresses on the tool after cooling were observed. Again, the same relation was found: 2-beads developed more compressive stress on the toll than 5-beads, and bigger beads compressed the tool more than smaller beads.

3.4. Optimization of the Multiple-Layered Weld-Cladding Patterns

Tool reinforcement improves load-bearing capacity and protects operators from unexpected tool failure. This section concerns thick or multiple layered weld beads. Following the previous findings, the two parallel beads welded individually in a single layer induced more tool compressive stresses than the five parallel beads. Therefore, in this section, only two parallel beads will be tested in multiple layers (two and three layers), and their outcomes will be compared and optimized to enhance the compressive stresses in the tool. In the latter text following acronyms will be used:

➢

2 × 1 = A single layer of 2 beads welded parallel to each other.

➢

2 × 2 = Two welding layers, each layer having two parallel beads.

➢

2 × 3 = Three welding layers, each layer having two parallel beads.

The 2 × 1 pattern has already been discussed in Section 2.2.1. A schematic illustration of the 2 × 2 and the 2 × 3 pattern is shown in Figure 9.

3.4.1. Influence of Different Trajectories

As described in Section 3.3.5, the thermal history of the weld-cladding process depends on the welding trajectory. In this section, different welding trajectories (T1–T5) (see Figure 11) in the 2 × 2 pattern were compared to determine the best welding trajectory to generate the maximum compressive stress in the tool. A comprehensive comparison of the stress histories of the aforementioned trajectories welded with beads of different dimensions, i.e., A1–A4, is shown in Figure 26.

Among all the trajectories T1–T5, the trajectory T1 induced the maximum compressive stress on the tool. The maximum compressive stress of σ_{m, max} = −113 MPa, has been induced on the tool using A4 beads. Trajectory T2 also appeared to be a good trajectory with cross-sections A1 and A2 (see Figure 26). With beads ranging in size from A1 to A3, the trajectory T4 produced maximum compressive stress of about σ_{m, max} = −108 MPa with A3 beads. T1 is regarded as a better and more optimal trajectory than the other trajectories, since the beads in this trajectory generated the most considerable compressive stresses on the tool after cooling.

3.4.2. Influence of Adding Multiple Layers

A thick layer of reinforcement in a 2 × 3 (three-layer) pattern has been simulated using trajectory T1 to avoid tool breakage owing to overload. A detailed comparison of two parallel beads in one, two, and three layers is shown in Figure 27.

It can be noticed that the beads in a single layer (2 × 1 pattern) developed the maximum compressive stress on the tool surface. The maximum compressive stress in the tool was generated by the larger beads (A4) with its maximum value as σ_{m, max} = −128 MPa. The A4-beads developed less compressive stresses in the tool with a 2 × 3 pattern than the 2 × 2 or 2 × 1 pattern. The reheating of the layers added more heat to the tool, resulting in slow cooling and variations in the microstructure and mechanical properties. Consequently, the beads do not shrink as much as laid down in a single layer.

3.5. Maximum Compressive Stresses in Different Practical Scenarios

3.5.1. Reinforcement of the Tool under Operating Conditions

At higher operating temperatures, residual stresses relax over time [75,76]. The behavior of the extrusion tool reinforced by the weld-cladding process and conventional method is examined at different temperatures achieved during extrusion to evaluate the effect of extreme conditions on the tool’s reinforcement. The reinforced tools are heated to higher temperatures, i.e., from 25 °C to 800 °C, corresponding to the operating temperatures during extrusion. After cooling from various operating temperatures, the compressive stresses (reinforcement) on the tool surface (under the bead) were reduced. Compressive stress decreased following the same pattern in weld-cladding and conventional processes. During the weld-cladding process, the reinforcement on the extrusion tool declines at a rate of 8% for every 100 °C increase in temperature from 25 °C to 400 °C. At higher working temperatures from 400 °C to 700 °C, the tool’s reinforcement deteriorated at a rate of about 40%. At 800 °C, the tool’s reinforcement decreased to approximately 98% of its starting value, reaching just 4 MPa after cooling. In contrast to the weld-cladding process, the tool reinforced by the conventional method lost its strength rapidly up to 500 °C. As operating temperatures climb from 600 °C to 800 °C, the decline in tool reinforcement with temperature becomes the same with both methods, as shown in Figure 28.

Since the same material was used for the computational analysis of weld-cladding and conventional methods, a similar trend was observed under different operating conditions. As a result, it may be anticipated that the material’s thermo-mechanical properties significantly influence the reinforcement. Hence, weld-cladding can be employed as a low-cost alternative to more expensive traditional procedures for providing enough reinforcement on the components.

3.5.2. Influence of the Tool’s Diameter on the Maximum Compressive Stresses under Different Operating Conditions

Weld cladding on several tool diameters, i.e., 40 mm, 45 mm, 50 mm, 55 mm, 60 mm, and 65 mm (see Figure 7), was numerically tested to determine if the tool diameter changes the compressive stresses on the tool surface. The width and height for each bead were kept constant, i.e., 8 mm and 5 mm, while the inner diameter of the beads was adjusted to fit the outer diameter of the tool. Simulations were carried out at three different operating conditions as mentioned:

(i)

Constant velocity: The welding velocity and heat input remained unchanged, while the welding time was adjusted based on the length of the welding path.

(ii)

Constant welding time: The welding time was constant, while the welding velocity and heat input were adjusted based on the cross-sectional area of the weld-bead.

(iii)

Maximal reinforcement: The entire bead was heated beyond the melting point and then cooled to room temperature using the efficient method.

In the first case, as the tool’s diameter increases, so does the volume and the welding path; thus, extra welding time is required to weld the bead. In the second case, the effect of a larger diameter is compensated for by changing the welding velocity and power. The total heat input on a unit volume remains constant while maintaining the welding time. The third case considers an efficient situation in which local heating and cooling effects are eliminated by heating the bead above the melting temperature, i.e., 1700 °C, followed by air cooling to room temperature.

The average of maximum-compressive stresses of all the tool’s elements under the bead at different operating conditions is shown in Figure 29.

The heat loss to the air intensified as the tool’s diameter increased. In all cases, the compressive stress developed by the bead on the smaller diameter tool was less than on the larger diameter tools. When the tool diameter increased, so did the length of the weld-bead and its subsequent heat input, resulting in enhanced compressive stress on the tool surface after cooling. When the tool diameter exceeds 50 mm, the curves of cases 1 and 2 become nearly identical. Since the behavior of compressive stresses with increasing tool diameter is consistent in all situations, it also confirms the validity of the weld-cladding approach used in this work.

3.5.3. Influence of Heat Source Velocity on Maximum Compressive Stresses

Since the welding velocity influences the residual stresses of the substrate [77,78], different welding velocities have been tested using the standard weld-cladding numerical procedure and material properties of steel 309L. The dimensions of the tool and the bead were kept consistent for all models. The heat input in all the models was kept the same by keeping a constant ratio between the input energy and the heat-source velocity. After welding, the models were allowed to cool for the next 400 s before being analyzed. The following heat source velocities were tested i.e., 7.5 mm/s, 15 mm/s, 20 mm/s, 25 mm/s, 30 mm/s, and 35 mm/s. The average of the maximum compressive stresses on all the elements of the tool surface under the bead at different heat-source velocities is shown in Figure 30.

As the heat source travel speed increased, the maximum compressive stresses on the tool surface increased up to a velocity of 20 mm/s, where the maximum compressive stress of σ_{m, max} = 201 MPa was reached. With the further increase in heat source travel speed, the maximum compressive stresses on the tool surface remained almost unchanged up to 35 mm/s. Due to insufficient heating of the bead and tool, the compressive stresses in the tool decreased as the heat source travel speed increased beyond 35 mm/s. Hence, the optimal velocity range for welding the tool with the given dimensions is between 20 mm/s and 35 mm/s.

4. Discussion

In this paper, the reinforcement using the weld cladding process has been discussed by undergoing various numerical and experimental investigations. Various geometrical and process parameters that can influence the residual stresses were investigated based on the existing literature and own observations. The concept of reinforcement has been utilized to enhance the effective compressive stresses on the extrusion tool by varying various parameters. As proposed by Jousten et al. [70], the number of beads and their cross-section could affect the stress distribution on the tool surface; hence the weld beads of various cross-sections have been analyzed. It has been found that the compressive residual stresses on the tool vary with the size as well as the number of beads. Small beads of size A1 induced less compressive stresses on the tool as compared to the larger beads, i.e., A4. Similarly, the number of beads on the tool also affects the resulting compressive stresses, as also reported by Hu et al. [72]. The tool with two parallel weld beads developed more compressive stresses on the surface of the tool under the beads as compared to the five beads due to repetitive heat input and thermal softening effect. The repetitive heating of the substrate in multi-pass arc welding influences the fatigue strength of the material, as also informed by Ganguly et al. [73]. Therefore, to compensate for this effect and enhance the reinforcement of the tool even with multiple beads, a critical gap of 11 mm between the two consecutive weld beads was found effective. According to Hosseini et al. [48], heat input, the geometry of the substrate, and the welding sequence influence the resulting thermal cycle and the stress state. The effect of geometry on the residual stresses has been discussed in this work. The results indicated that the residual stresses on the tool surface after weld cladding depend on the geometry of the tool. For example, more compressive stresses are produced in the welding zone after cooling if there is no or small hole in the geometry. Similarly, the different welding sequences/trajectories have been tested. The results indicated that the welding sequence significantly influences the residual stress development on the tool surface. The welding trajectory T1, in this work, appeared to be the best strategy which developed maximum compressive stresses on the tool’s surface with different bead cross-sections. According to Rehman et al. [61], time delay has a great influence on the mechanical properties of the material. Adding more delay time between the welding adds tensile stresses on the substrate, which could deteriorate the mechanical properties of the specimen. In this work, a time delay of 20 s has been kept between the two consecutive weld beads. The results indicated that the weld beads without delay developed comparatively more compressive stresses on the tool’s surface than beads with a time delay between them.

As the residual stresses relax with time at a higher operating temperature [75,76], the reduction of the reinforcement under the actual operating conditions during the extrusion was investigated. The results indicated that with the increasing temperature during the extrusion process, the amount of compressive residual stresses on the tool reduces, resulting in the reduction of the mechanical strength of the tool. Similarly, in the case of a cylindrical tool, the influence of the tool’s diameter under different operating conditions has been investigated. The welding velocity also influences the resulting residual stresses [77,78] of the substrate. Therefore, the variation in the compressive residual stresses welded at different welding velocities has been presented.

5. Summary and Conclusions

In the present research, a novel approach is presented to utilize the weld-cladding process for the targeted generation of desired compressive residual stresses with the example of a demonstrator tool. During welding, thermal expansion and contraction of weld beads have been effectively used to develop compressive stresses near the tool’s surface. The experimental results validated the temperature distribution and residual stresses of the numerical investigations. Simulations with various welding and geometrical parameters such as continuous welding (with and without delay) and welding in different trajectories were performed to obtain the optimized process parameters and path for tool reinforcement. A comparison of weld-cladding and conventional methods was performed to validate the efficiency of the weld-cladding in various practical scenarios.

The conclusions and suggestions that can be drawn from this work are as follows:

• Approximately 40–60% of residual stresses in the tool surface are induced by mechanical contact of beads to the tool, while the rest developed due to repeated thermal cycles.
• The thermal contraction of the weld bead and the thermal expansion and contraction of the tool during the wire-arc welding process can be used to strengthen the tools.
• Repeated heating and cooling of beads accumulate heat within the model, resulting in slow cooling and reduced tool reinforcement.
• In the multiple bead welding process, reheating reduces the compressive stresses generated by the previous beads on the tool. With the 100 mm diameter of a tool, and beads of dimension A3 (h = 4 mm, w = 6 mm), establishing a minimum gap Δx ≥ 11 mm reduces the thermal influence of the adjacent beads.
• During welding, the weld bead with a larger cross-section should be preferred. The beads of dimension A4 (7.5 mm × 5 mm) introduced the maximum compressive stresses in the tool in almost all the welding patterns. Therefore, they can be adapted to reinforce the components more effectively.
• The weld-cladding of multiple beads (5-beads) in parallel sequence without delay adds more reinforcement to the tool and minimizes distortion compared to the welding with a time delay between each bead.
• Due to heat accumulation, adding each new layer in a multilayer weld-cladding process mitigates the compressive stresses in the tool. For example, when welding the tool with A4 beads, the 2-layered (2 × 2) pattern produced approximately 11–14% less compressive stresses than the 1-layered (2 × 1) pattern. Similarly, the 3-layered (2 × 3) pattern induced even less compressive stresses in the tool.
• The tool’s reinforcement is reduced to 98% of its initial value during an extrusion operation at an operating temperature of 800 °C.
• Internal features such as a hole or a notch reduce the tool’s stiffness.
• Increasing the tool and bead diameters improves heat flow and increases the maximum compressive stresses on the tool surface after cooling.

Based on the investigated rules, an optimized weld-cladding approach with better parameters can be designed to achieve low-cost reinforcement of different components. Overall, reinforcement of tooling by weld-cladding could be a low-cost alternative that incorporates considerable compressive stresses with a single weld bead or thin layer. If thick reinforcements are required, multiple layers should be welded together. Future research would incorporate the effect of recrystallization, creep deformation, and transformation plasticity in the simulation model for detailed micro and sub-micro stress analysis.