Advantages of Corrosion-Resistant Overlay Welding on Steel S355J2N

Abstract

In this paper, the effects of overlay welding of S355J2N steel were studied. We examined how the technological advantages of overlay welding can be taken into account to improve the service lifetime and applicability of components made from traditional S355J2N structural steel during the planning step. Increasing the service life of structures exposed to environmental influences is essential, especially on surfaces exposed to abrasive and chemical corrosion. The direct aim of the investigation was to present a comprehensive picture of technological advantages of the corrosion-resistant overlay welding on steel S355J2N. We mainly analysed experiments with powder-coated wire electrodes which are based on protective gas and robot technology usage. With various mechanical tests, we searched for the minimum number of layers that provides sufficient protection against corrosion. The aim of this paper is to present achieved results during development of a welding technology of a reliably functioning product with increased corrosion resistance.

1. Introduction

Many components are expected to have a tough internal microstructure, while their surfaces should be wear-resistant or corrosion-resistant. Failures, especially in pressure vessels and pipe fittings, occur along boundary conditions, leading to perforation [1,2]. In such cases, stress corrosion or wall thinning caused by internal cavitation can lead to leakage [3]. A local repair of this failure, using overlay welding in the installed environment, can be an effective and environmentally friendly solution. Naturally, research is also underway to explore materials that attempt to eliminate the perforation conditions in tanks and other pipe fittings [4]. Other research directions focus on developing the technological databases of laser overlay welding [5]. This is a relatively new area of usage in industry. In addition, research is also underway to reduce the heat-induced deformations and deflections in the field of various overlay welding processes as well [6]. Given the ongoing energy management challenges of our time, overlay welding interventions are becoming increasingly significant [7,8,9,10]. As the demand for corrosion-resistant materials grows across various applications, there is an ongoing need for innovative solutions that enhance material performance and cost-effectiveness. Thus, choosing appropriate materials and measurements is essential to guarantee safe and reliable operations. Overlay welding is a widely used technique for improving the surface properties of base materials by creating a layer that protects the base metal from corrosion. When overlay-welding different components, a key factor is cost-effectiveness, as the proper choice of welding material and technology can ensure a long service life for the product [11].

Considering that the composition of the deposited layer in the welding of heat-treated steels may differ significantly from that of the base material [12], it is important to select technological parameters that reduce the proportion of the base material in the weld. A well-adjusted mixing ratio in the buffer and cladding layers helps to create the right ratio of alloying elements to achieve the desired corrosion resistance and homogeneous material structure. This requires controlled heat input, adequate seam overlap, and well-designed welding technology and welding speed [13,14]. In our research, we investigated the overlay welding of S355J2N structural steel with a multi-layer, highly alloyed, corrosion-resistant austenitic filler material. Unalloyed structural steels are widely used types of steel, primarily used in welded joints exposed to constant or time-varying mechanical stresses, depending on their application. Their popularity is due to their favourable mechanical properties, low cost, versatility in production, and high formability [15,16]. In industry, it is common to replace more expensive steel types with this steel group when the requirements for mechanical stresses allow doing so, in order to minimize material costs [12].

Since we welded with a filler material of significantly different composition and properties, it was necessary to create an intermediate layer made of an austenitic filler material, which we refer to as the buffer layer [17]. An important objective of our work was to determine at what depth we can speak of the formation of a corrosion-resistant layer. The study employs various testing techniques, including visual testing, macroscopic and microscopic examination, Vickers hardness testing, salt spray testing and electron microscopic examination. The formation of a suitable corrosion-resistant layer depends significantly on the Cr content enriched in the δ-ferrite of the mixing zone. Taking all this into account, we aimed for a δ-ferrite–austenite-type solidified material structure, which is ensured by Cr/Ni equivalent ratios between 1.48–1.95 [18]. At the same time, we had to avoid the appearance of sigma phases as well. Another goal was to develop a method that facilitates the determination of the optimal relationship between the required level of corrosion resistance and the deposited layer thickness for a given welding system.

2. Materials and Methods

2.1. Materials

From a welding perspective, it is very important that welding of structural steels does not necessarily require preheating, due to their low carbon content (C < 0.2%), thus eliminating the risk of hardened weld structures and the associated risk of cracking [19].

For our experiments, S355J2N steel was used as base material. The characteristics of hot-rolled, unalloyed structural steels are summarized in the EN 10025-2 standard [20]. Steel grades marked with the letter ‘S’ have low carbon content and contain a maximum of 0.55% silicon. The chemical composition of the S355J2N steel grade, which is important for the research, can be found in

Table 1. Chemical composition of S355J2N.

Steel Name	Steel Number	C % Max.	Si % Max.	Mn % Max.	P % Max.	S % Max.	Cu % Max.	Other % Max.
S355J2N	1.0577	0.20	0.55	1.60	0.025	0.025	0.55	-
base metal		0.16	0.25	1.18	0.013	0.013	0.01	-

.

The application of each layer during the overlay welding process was carried out based on the professional recommendations of the Kobe Steel Ltd. (Chuo, Japan) Welding Handbook [17]. For the buffer layer (covering layer), we selected the Voestalpine (Düsseldorf, Germany) austenitic stainless rutile flux-cored wire of type T 23 12 2 L P/E309LMo. For the other layers, we used the T 19 12 3 L/E316L type flux-cored wire with D = 1.2 mm (

Table 2. Chemical composition of the electrodes.

Tubular Cored Electrode	C %	Si %	Mn %	P %	S %	Cr %	Mo %	Ni %	Cu %
FOXcore 309LMo-T0	0.02	0.70	1.30	0.024	0.005	22.00	2.60	13.30	0.10
FOXcore 316L-T0	0.02	0.62	1.48	0.025	0.010	18.96	2.82	12.28	0.11

).

The filler materials selected for the experiment are chemically resistant and qualified as ‘highly suitable for welding’. Thanks to their extremely low carbon content, they cannot be hardened, meaning that no significant hardening is expected in the heat-affected zone (HAZ), and there is no expected significant grain coarsening as well.

For shielding gas in the welding process, we used CORGON gas (82% Ar + 18% CO₂) to ensure arc stability, minimize spattering, and provide effective protection against oxidation.

2.2. Welding Process

The welding base system (at the Department of Materials Science and Technology, SZE, Győr, Hungary) consists of an IGM KUKA RTE-4 robotic arm, which features 6 + 2 rotational axes and a TPS 4000 power source (IGM Robot Systems Ltd., Győr, Hungary) (Figure 1).

Welding by the robot was performed using flux-cored wire, which is considered a modern filler material. Compared to solid wire, it offers several advantages, such as high productivity, applicability to a wide range of chemical compositions, lower sensitivity to edge misalignment, a lower average temperature of the weld pool, reduced spattering, and stable arc behaviour; it also keeps the contaminant levels low. Its disadvantages include higher procurement costs and the need for skilled personnel.

There are specific rules for the use of flux-cored wires, which must be followed during application, as the ferrite number is significantly reduced if carbon and nitrogen from the environment enter the weld (such as from the workpiece surface, the shielding gas, the welder’s gloves, etc.). Variations of flux-cored arc welding with shielding gas are well suited for overlay welding purposes, as they can be applied within a wide current range. During the experiments, the chosen welding materials were always used with consideration of a recommended shielding gas for the wire, and these parameters are included in the welding procedure specification. The key parameters of the welding procedure specification relevant to our experiment are included in Figure 2.

2.3. Test Methods

The overlay welded parts were cut by water-jet cutting equipment, and the samples were characterized by different methods. The macro- and microstructure were studied by microscopic investigations after surface preparation. For the macrostructure analysis, a Zeiss Stereo Discovery V20 microscope (Zeiss, Göttingen, Germany) was used, while for microstructure characterization, we applied a Zeiss Axio Imager A1 optical microscope (Zeiss, Göttingen, Germany).

The rough preparation of the samples was done using SiC grinding papers (STARCKE, Melle, Germany) of different grits (80, 180, 400, 600, 1000, 1200, and 2500), and polishing was performed with MD-Dac and MD-Chem polishing pastes (Struers ApS, Ballerup, Denmark). The etching was carried out with 6% nital and ammonium molybdate.

For a detailed analysis of chemical composition, mainly the Cr and Ni content of austenite and δ-ferrite, we used a HITACHI 3400 (Hitachi, Tokyo, Japan) for scanning electron microscopy (SEM), as well as energy-dispersive X-ray analysis (EDX).

The hardness measurement was performed according to the ISO 15614-7 standard [21] at room temperature, using the Vickers method based on the ISO 6507-1 standard [22] with a KB 30-type hardness tester (KB Prüftechnik, Hochdorf-Assenheim, Germany). The load applied was F = 98.1 N (10 kp).

For characterization of the corrosion susceptibility of the welded part, a salt spray test was performed using an Ascott CC1000iP device (Ascott, Tamworth, UK) according to the ISO 9227 NSS standard [23]. The relevant testing parameters were as follows:

• Test duration: 336 h (2 weeks).
• Test chamber temperature: 35 °C.
• Corrosive medium: 5% salt solution (deionized water + salt); the salt quality was Ascott Corro salt.
• pH of the solution: 7.05; pH of the runoff solution: 7.00; average runoff rate: 1.50 mL/h.

3. Experimental

3.1. Theoretical Background of the Weldability of Corrosion-Resistant Steels

From the perspective of weldability, the carbon equivalent (CEV) is of crucial importance, as it helps determine the combined effect of carbon and other chemical elements on the steel’s susceptibility to cold cracking. For good weldability, it is important that the steel grade has a low carbon equivalent, as the heat introduced during welding not only melts the base material and filler material but also significantly heats the surrounding area of the weld.

The equation required to calculate the carbon equivalent, according to the EN 1011-2/A1 standard [24], is as follows:

$$CEV=C+{\displaystyle \frac{Mn+Si}{6}}+{\displaystyle \frac{Cr+Mo+V}{5}}+{\displaystyle \frac{Ni+Cu}{15}}$$

(1)

Using the Graville diagram, structural steels can be categorized into three groups based on their weldability. In zone I, there are easily weldable steels; in zone II, there are steels that should be welded carefully, and in zone III, structural steels that are difficult to weld due to their higher carbon content are found (Figure 3).

Due to its low carbon content and the small proportion of alloying elements, the S355J2N belongs to the carefully weldable category, meaning that controlled heat input can prevent the structure from becoming prone to cracking.

A wide range of corrosion-resistant steels are available in the industry, offering applications in various fields. If the steel contains at least 12% chromium (level I resistance limit) and the corrosion rate in a given medium does not exceed 0.1 mm/year, the steel is considered to be corrosion-resistant [26].

The corrosion resistance of these steel types is due to the thin, pore-free oxide layer that forms on their surface. They generally contain low amounts of carbon and more alloying elements. Minimizing the carbon content is particularly important for improving corrosion resistance.

Austenitic corrosion-resistant steels are among the most well-known and commonly used steel types due to their excellent formability, heat resistance, and weldability. In their microstructure, the dominant phase is austenite, which provides good corrosion resistance against most media. Their strength is roughly comparable to unalloyed steels, with outstanding toughness, and their tear elongation is around 35–45% [26].

A common issue in overlay welds is that the first pass of the heavily alloyed weld, mixed with the base material, results in a brittle, martensitic microstructure. The bond between the two materials with significantly different compositions can only be achieved by creating a buffer layer. Based on the microstructure of the weld, the mechanical properties, welding characteristics, and resistance to corrosion can be well estimated. The most widely used method for determining (or estimating) the microstructure of the weld metal is the Schaeffler diagram, which has been in use since 1949 and shows the phase relationships of heavily alloyed steels [27,28].

During the production of austenitic corrosion-resistant steels, the goal is to achieve a homogeneous austenitic microstructure in order to avoid precipitates, lattice defects running to the surface, irregularities, cracks, and corrosion damage. However, in many cases, other phase elements and precipitates make the microstructure somewhat heterogeneous. In practice, corrosion-resistant steels are not completely austenitic but contain small amounts of δ-ferrite, which is specifically beneficial, as it reduces cracking sensitivity [28]. In addition to the above, it must also be noted that the presence of ferrite above 10% reduces corrosion resistance. Between 800–850 °C, ferrite transforms into the σ-phase, which is highly susceptible to corrosion.

The ferrite content, or ferrite number (FN), can be determined from the chemical composition of the welding seam. Various microstructural diagrams can be used for this purpose. These include the previously mentioned Schaeffler diagram, as well as the WRC-92 diagram [29].

3.2. Theoretical Background of the Welding Process

Applying the appropriate welding conditions is crucial for controlling dilution, especially for the first layer (to avoid hot cracking). Thus, for using the D = 1.2 mm wire electrode, the welding current and welding speed were set within the recommended range, near the upper limit, in order to achieve better welding power [17,30]. This was later taken into account during the fusion of the buffer layer and the cladding layers.

The composition of the filler materials indicates that the first layer contains more chromium and nickel alloying elements than the second and third layers. This is necessary to prevent corrosion from starting at the boundary from the base plate, as the S355J2N steel is strongly deficient in chromium, and we need to ensure corrosion resistance even at the boundary areas. Moreover, the higher alloy content makes the filler material more expensive, so from a cost-effectiveness perspective, it is also advisable to use a cheaper electrode for the second and third layers.

The filler materials selected for the experiment are chemically resistant and fall into the ‘highly suitable for welding’ category. Thanks to their extremely low carbon content, they cannot be hardened, meaning that no hardening occurs in the heat-affected zone (HAZ), and there is no significant grain coarsening. By using the selected shielding gas (82% Ar + 18% CO₂), we managed to ensure arc stability, minimize spattering, and provide effective protection against oxidation.

Certain difficulties may arise from choosing an inappropriate technology, which include the following:

• Sensitization, i.e., a decrease in corrosion resistance due to the formation of chromium carbides (Cr₂₃C₆), so the goal when forming the buffer layer is to achieve a high ferrite number (FN > 5).
• Hot cracking, i.e., the formation of a liquid layer along the grain boundaries in the heat-affected zone. To prevent this, the Cr_(E)/Ni_(E) ratio must be maintained above 1.6 [30].
• Secondary precipitates, such as the formation of the sigma phase, which should be minimized in the second and third weld layers to adjust the amount of maximum allowable FN [31].

These issues can be avoided by selecting the appropriate welding materials and the optimal technology. The delta-ferrite content of the welding seams must be in the 3–15 FN range [30]. This provides adequate resistance to hot cracking, as delta-ferrite better dissolves the impurities (S and P) primarily responsible for crystallization cracking, while not compromising corrosion resistance.

3.3. Mixing Between the Base Material and the First Layer

In the case studied, the mixing with the S355J2N structural steel material must be kept below 35% during the welding of the first layer. If the mixing exceeds this value as a result of the welding process, the ferrite content of the weld must be determined with a calibrated ferrite content meter, or an estimate should be made based on the chemical composition, e.g., using the WRC-92 diagram.

Among the listed issues, grain boundary corrosion is the potential source of failure that required special consideration during the experiments. When selecting the filler material for the first welding seam, we aimed to create a microstructure which, while allowing for the desired mixing, would not produce carbide precipitates, or would do so only in small amounts. The condition for this is that carbon must not be able to withdraw chromium in carbide form at the grain boundaries, thus reducing the corrosion resistance of the base metal below level I (12%, intergranular corrosion) [26]. One way to achieve this is by binding carbon through the previously mentioned titanium and niobium alloying or by increasing the ferrite number, which leads to the formation of carbon-enriched δ-ferrite in the base material. This ferrite cannot harden because there is no austenitic phase during cooling.

Therefore, with the proper technology (controlled cooling process and limited heat input energy), grain boundary corrosion can be excluded.

The first layer is essentially the filler material of the buffer layer. During the process, care had to be taken to ensure that the joint did not heat to temperatures higher than the permissible limit; otherwise, the weld would become inhomogeneous, and cold cracking would occur. Achieving a homogeneous microstructure is a fundamental goal in corrosion-resistant steels, as any inhomogeneity can become a starting point for corrosion attack. In our experiment, the first weld pass served as an intermediate layer, since the composition of the second and third layers significantly differed from that of the low-alloy structural steel, and our goal was to achieve a purely austenitic structure. The appropriate mixing ratio of the buffer layer was determined based on the chromium and nickel equivalencies of the base metal and the E309LMo filler wire, using the Schaeffler diagram as a reference. Here, we took into account that, in the classic case, the bonding surfaces of the filler material and the base metal generally melt at approximately the same rate, and the filler material typically “dilutes” about 20–40% of the molten base metal. Since, in the formation of the buffer layer, in addition to the composition differences, the heat input energy had to be controlled as well, we used the setup data from Figure 3 to consider 70% of the 30% average mixing, taking the technological settings into account. For the deposition of the second and third layers, the 30% mixing ratio was found to be appropriate.

Since we used the Schaeffler diagram to determine the mixing points, we first clarified the methods for calculating Cr_(E) and Ni_(E) according to Schaeffler (2) and WRC-92 (3). The WRC-92 diagram was used to verify the obtained values.

Schaeffler:

$$\begin{gathered} {Cr}_{\left(ES\right)}=Cr\%+Mo\%+1.5Si\%+0.5Nb\%+2Ti\% \\ {Ni}_{\left(ES\right)}=Ni\%+30C\%+0.5Mn\% \end{gathered}$$

(2)

WRC-92:

$$\begin{gathered} {Cr}_{\left(EW\right)}=Cr\%+Mo\%+0.7Nb\% \\ {Ni}_{\left(EW\right)}=Ni\%+35C\%+20N\%+0.25Cu\% \end{gathered}$$

(3)

During the experiments, the electrode used for welding the first layer, mixed with the base material (S355J2N) at approximately 21%, resulted in the weld microstructure shown in Figure 4.

In Figure 4, the horizontal axis of the diagram shows the chromium equivalent containing ferrite-forming alloying elements, while the vertical axis represents the nickel equivalent derived from austenite-forming alloying elements. The multiplication factors of the individual alloying elements (2) indicate their strength of influence on the microstructure. Using this diagram, a good approximation of the material’s expected composition can be determined, which is of crucial importance when selecting welding materials. Based on the diagram, it can be concluded that nearly 9% pure delta ferrite can form in the weld, which is sufficient and necessary to avoid sensitization [32]. The nickel equivalent is approximately 12, while the chromium equivalent is around 21. The Cr_(ES)/Ni_(ES) ratio in the weld is 1.728. These values, when plotted on the WRC-92 diagram in Figure 5, also yield close to 9% delta ferrite. The Cr_(EW)/Ni_(EW) ratio in the weld is 1.711.

3.4. Mixing Between the First and the Second Layer

Based on the Schaeffler diagram shown in Figure 6, it can be concluded that nearly 9.5% pure delta ferrite can form in the weld, which provides a good approximation of the ferrite number shown in Figure 4. The nickel equivalent is approximately 13, and the chromium equivalent is around 22. These values indicate that a nearly homogeneous material structure is expected. The Cr_(ES)/Ni_(ES) ratio in the weld is 1.697.

The ferrite number of the filler material for the second layer (E316L) is 10 FN, according to the WRC-92 calculation method (3). The ferrite content in the overlay welded cross-section is also approximately 10%, as shown on the WRC-92 diagram in Figure 7 [32]. The chromium equivalent number of the second weld is approximately 21%, according to calculation (3), and the nickel equivalent number is closer to ~12.5%. The ferrite number in the welds and diffusion zones did not significantly change, due to the constancy of the heat input energy [33]. The Cr_(EW)/Ni_(EW) ratio in the weld is 1.693.

4. Results

Figure 8 shows a photo of one of the test specimens welded with three layers on a 12 mm thick base plate, in cut condition (water jet). The overlay welding was performed on a 150 × 300 mm S355J2N structural steel plate within a 100 × 200 mm area.

4.1. Macroscopic and Microscopic Examination

Figure 9 shows macro photos of the three-layer samples, selected from the grinding cross-sections perpendicular to the weld, with nearly identical thicknesses, referring to the largest and smallest thicknesses of the deposited layers.

The total thickness of the three layers in the first sample (Figure 9a) is on average 12.3 mm, while in the second sample (Figure 9b) it is 11.8 mm. These measured values correspond to an average thickness of 4.1 mm and 3.9 mm per layer, respectively.

In the following, we will investigate how many layers are necessary to achieve optimal corrosion resistance.

Based on the images in Figure 10, it can be concluded that the weld’s microstructure is austenitic, with delta ferrite. This confirms that a heterogeneous microstructure indeed forms, which represents a kind of compromise, as it allows the corrosion advantages of both ferritic and austenitic structures to be combined.

Homogeneity should primarily not be understood as the uniformity of the microstructure, but rather as the structure’s behaviour as a compact, unified whole against corrosion. Clearly, this compromise requires that these inhomogeneities be carefully controlled through the selection of appropriate technology.

Some sigma-phase precipitation can also be observed in smaller clusters in the second and third layers, as well as dispersed M₂₃C₆-type complex carbides and a negligible amount of slag inclusions in the buffer layer’s mixing zone. The quantity and impact of all these do not reach the threshold for accessibility, so the material structure, according to Figure 10f, can be considered uniform and homogeneous.

4.2. Vickers Hardness Test

When measuring the hardness of welded joints, the hardness (HV10) of the base metal (BM), the heat-affected zone (HAZ), and the weld metal (WM) must be measured in all cases. In the case of overlay welding, according to the ISO 15614-7 standard [21], an indentation series must be made at a 15° angle, as shown in Figure 11.

As shown in Figure 12, due to the low carbon content, no significant hardening occurred in the heat-affected zone, and the hardness of the weld does not differ practically from the base metal. Hence, it can be safely concluded that no significant precipitations were formed.

Insignificant but noticeable differences can be observed in the hardness distribution. The minimum hardness differences are basically determined by the differences between the mechanical properties of the flux-cored wires and the base metal, based on the differences in chemical compositions. We measured an average tensile strength of 530 MPa on the base metal, according to the EN 10025-2 standard [20]. Its converted hardness: 165 HV. According to the technical documents of the flux-cored wires, the tensile strength of the filler material of the first layer (electrode: E309LMo) is 700 MPa (218 HV), and the tensile strength of the same filler material of the 2nd and 3rd layers (E316L) is 560 MPa (175 HV). The data in the technical certificates are certified data; therefore, their determinations by new measurement are not necessary. The difference between the tensile strength of the base metal and the E316L electrode is only 30 MPa, which when converted to Vickers hardness corresponds to a value of 10 HV. Figure 12 shows almost identical values at the end points (point 1: 183 HV, point 18: 180 HV). The explanation for this is that, at point 1, the micro-structure of the material also contains tempered martensite, due to the intensive heat removal from the back side, while at point 18, as we will see later, Cr enrichment occurred. The highest hardness values are found at measurement points 7–8 (211-210 HV). These values are almost identical to the values characteristic of the E309LMo electrode, and their minimal decreases can be explained by the diffusion of Cr into the surface layers. During examinations, no appreciable carbide precipitates were found. Consequently, the material structure can be considered homogeneous.

Furthermore, it can be concluded that the hardness values measured in the base material do not exceed the values specified in the standard (320 HV10).

4.3. Salt Spray Test

In corrosion testing, the goal is to assess the corrosion susceptibility by artificially simulating the effects on the sample.

Figure 13 shows the state of the overlay welding before (Figure 13a) and after (Figure 13b) the salt spray test. The strong corrosive environment caused significant corrosion damage to the S355J2N base plate, but the corrosion-resistant layer appears undamaged to the naked eye. This conclusion is confirmed by markers indicated with red arrows.

In order to compare the corrosion resistance of the welded joints with other steels of similar composition, we examined the resistance to pitting corrosion by determining the PREN (Pitting Resistance Equivalent Number) index [34]. The composition of the E316L wire electrode used in the second and third layers is very similar to the 316L (min. 2.5% Mo) austenitic steel, which has a characteristic PREN value of 25.3–30.7. Therefore, the PREN value of the deposited overlay welding should also fall within this range. The PREN value is determined according to the following Equation (4):

$$PREN=Cr\%+3.3Mo\%+16N\%$$

(4)

4.4. Microscopic Examination After Salt Spray Test

Microscopic images taken after the salt spray test are shown in Figure 14.

Figure 14 clearly shows the corrosive effect on the S355J2N steel (Figure 14a), while no change occurred in the weld microstructure despite the strong corrosive environment (Figure 14b–e).

4.5. Electron Microscopic Examination

In order to obtain a more detailed image of the specimen, we also took images with a scanning electron microscope with a greater depth of field.

The scanning electron microscope creates images from secondary electrons, thus providing information about the chemical distribution as well. The EDX analyses were performed using a Bruker X-ray spectrometer (Bruker AXS Mikroanalysis GmbH, Berlin, Germany) -equipped Hitachi 3400N scanning electron microscope.

The locations of the point analyses are marked with a green cross and a red arrow. Figure 15 shows the results of the first austenite point analysis (FH-WM1 austenite 1) of the first layer, while Figure 16 displays the results of the first delta ferrite point analysis (FH-WM1 delta ferrite 1) of the first layer.

Point analyses were also performed in the base material, the welds deposited with filler materials, and the mixing zones, with two analyses conducted in each area. We were interested in the proportions of alloying elements in the different microstructural phases (austenite, delta ferrite). The images also show smaller carbides, inclusions, and pores, which were examined as well, but their number is so minimal that they do not play a significant role in the evaluation of the microstructure, and therefore, they are not included in the measurement results. The results are summarized in

Table 3. Chemical compositions of point examinations.

	Cr %	Ni %	Mo %	Mn %	Fe %	Si %
WM1 austenite 1	17.84	12.62	2.46	2.27	64.44	0.38
WM1 austenite 2	17.43	12.13	2.04	2.19	65.82	0.40
WM1 δ-ferrite 1	18.82	7.57	2.89	2.09	68.28	0.34
WM1 δ-ferrite 2	20.57	8.77	3.90	1.88	64.35	0.53
WM1-WM2 austenite 1	17.64	11.93	2.29	1.95	65.73	0.46
WM1-WM2 austenite 2	17.62	11.98	2.10	1.82	66.04	0.45
WM1-WM2 δ-ferrite 1	22.45	7.60	3.96	1.82	63.64	0.53
WM1-WM2 δ-ferrite 2	21.62	8.03	3.06	1.92	64.87	0.50
WM2 austenite 1	17.48	13.33	2.92	1.97	63.77	0.52
WM2 austenite 2	17.42	14.42	2.82	1.97	62.90	0.49
WM2 δ-ferrite 1	22.42	7.68	4.60	1.70	63.04	0.55
WM2 δ-ferrite 2	20.52	9.46	3.80	1.70	64.03	0.49
WM2-WM3 austenite 1	18.63	11.73	2.87	1.90	64.39	0.47
WM2-WM3 austenite 2	18.37	11.89	2.94	1.92	64.41	0.47
WM2-WM3 δ-ferrite 1	19.09	11.29	2.52	1.96	64.71	0.43
WM2-WM3 δ-ferrite 1	19.57	10.51	2.47	1.89	65.16	0.40
WM3 austenite 1	17.60	14.28	1.62	1.83	64.01	0.65
WM3 austenite 2	17.67	13.64	2.66	1.81	63.64	0.56
WM3 δ-ferrite 1	23.17	7.19	4.55	1.84	62.74	0.52
WM3 δ-ferrite 2	22.75	7.52	4.84	1.58	62.74	0.56

for easier reference.

In Figure 17, we contextualized the changes in Cr and Ni, which clearly form a directed pattern depending on which microstructural phase they are found in.

The diagram essentially corresponds to the weld’s resistance curve, which, despite fluctuations, shows uniform corrosion resistance well above the first resistance threshold. The curve clearly demonstrates the opposite directional changes in the amount of Cr and Ni alloying elements in the austenite matrix and delta ferrite net. The depicted trend indicates that the austenite matrix contains less chromium than the delta ferrite [29,35], but the austenite is not significantly depleted, which confirms the minimum risk of chromium carbide precipitation. In none of the measurement points did the chromium content fall below the first resistance threshold (12%), ensuring passivity against corrosion.

5. Discussion

Based on the data from the EDX analysis performed with the electron microscope, the following conclusions can be drawn for the buffer layer weld.

• Average alloy content: Cr 18.67%; Ni 10.27%; Mo 2.82%; Mn 2.11%; Si 0.41%.
• Average alloy content in the austenite: Cr 17.64%; Ni 12.38%; Mo 2.25%; Mn 2.23%; Si 0.39%.
• Average alloy content in the delta ferrite: Cr 19.7%; Ni 8.17%; Mo 3.395%; Mn 1.985%; Si 0.435%.
• Measured Cr/Ni ratio in the weld: 1.818.
• PREN value: 27.98.

It can also be concluded that, in the first weld, the Cr content in the austenite is 2.06% lower than in the delta ferrite. At the same time, the Ni content in the delta ferrite is 4.21% lower than in the austenite. The driving force behind these distribution differences lies in the solubility laws, dependent on the microstructure. These key enrichment differences determine the distribution of other elements, particularly Mo, in the austenite and delta ferrite.

Further analysis of the data in Table 3 for the two-layer overlay and their fusion zones leads to the following conclusions:

• Average alloy content: Cr 19.63%; Ni 10.78%; Mo 3.13%; Mn 1.85%; Si 0.50%.
• Average alloy content in the austenite: Cr 17.80%; Ni 12.90%; Mo 2.53%; Mn 1.90%; Si 0.51%.
• Average alloy content in the delta ferrite: Cr 21.45%; Ni 8.66%; Mo 3.73%; Mn 1.80%; Si 0.50%.
• Measured Cr/Ni ratio in the weld: 1.821.
• PREN value: 29.96.

It can also be concluded that, in the overlay welds, the Cr content in the austenite is 3.65% lower than in the delta ferrite. This difference has approximately doubled. This was caused by the accumulation of Cr in the delta ferrite, which diffused from the buffer layer towards the overlay due to the sufficiently long cooling time and high temperature. At the same time, the Ni content in the delta ferrite is 4.24% lower than in the austenite, which is virtually identical to the value measured in the buffer layer. The lack of change can be explained by the higher diffusion activation energy required for nickel compared to chromium [36]. No significant changes in the quantities and ratios of other key alloying elements can be observed compared to those seen in the buffer layer.

Further conclusions can be drawn by examining the total thickness of the deposited layers.

Based on the data in Table 3, the nickel content is lowest in WM3 δ-ferrite 1, at 7.19%, which still ensures an austenitic microstructure [35,37].

• Average alloy content considering the total (buffer layer + cladding layers) layer thickness: Cr 19.42%; Ni 10.67%; Mo 3.02%; Mn 1.90%; Si 0.48%.
• Measured Cr/Ni ratio characteristic of the deposited total layer: 1.82, which indicates a stable, uniform Cr/Ni distribution.
• Characteristic PREN value of the total layer: 29.386. This value shows a good match with the value typical for the 317L stainless steel family [38].

The uniform microstructure of the overlay weld (Figure 9) allows the establishment of a linear relationship between the Cr_(E) equivalent values and the measured Cr_(R) contents. The examination takes place separately in case of the buffer layer and complex cladding layers. The ratio pairs are as follows:

• Schaeffler (2): Cr_(ES)/Cr_(R) =1.133.
• WRC-92 (3): Cr_(EW)/Cr_(R) =1.083.

In the buffer layer (WM1):

$${Cr}_{\left(R\right)}={\displaystyle \frac{{21.13}_{\left(ES\right)}}{1.133}}=18.649\to 18.65\%;\hspace{1em}{Cr}_{\left(R\right)}={\displaystyle \frac{{20.224}_{\left(EW\right)}}{1.083}}=18.674\to 18.67\%$$

(5)

In the cladding layers (WM2-WM3 + fusion zones):

$${Cr}_{\left(R\right)}={\displaystyle \frac{{22.26}_{\left(ES\right)}}{1.133}}=19.647\to 19.65\%;\hspace{1em}{Cr}_{\left(R\right)}={\displaystyle \frac{{21.26}_{\left(EW\right)}}{1.083}}=19.631\to 19.63\%$$

(6)

The results show small deviations regarding the individual layers. The buffer layer already has a sufficiently high PREN value, and the measured Cr content significantly exceeds the first resistance level. Therefore, considering the minimum required layer thickness, the buffer layer could also serve as the overlay layer. Nevertheless, given the significant differences in the properties of the base metal and the deposited layer, a buffer layer is always necessary to create a joint with a stable transition. The required wall thickness, however, is determined by the required thickness of the layer.

Regarding the buffer layer, it is worth revisiting the estimated mixing ratio at the fusion zone, as shown in Figure 4. The heat input was maintained at a relatively high level via the heat input energy, while the electrode speed was at the upper limit [17]. Using a simple calculation method [39,40] and employing Figure 9 and Figure 11, the fusion (mixing) ratio (D) can be determined as follows, using any image editing software:

• We approximate the entire weld cross-section using a circle.
• We determine the original surface of the base metal, which will intersect the weld cross-section.
• We calculate the area of the resulting circular segment (A_BM) and relate it to the total weld cross-sectional area (A_BM + A_FM) (7).

$$D\%={\displaystyle \frac{A_{BM}}{A_{BM}+A_{FM}}}·100$$

(7)

Based on Equation (7), we performed checks at several locations, resulting in the D value fluctuating within the range of 21.5–23.5%, slightly deviating from our estimation. This can be attributed to several simplifying assumptions, and therefore, we believe that Equations (5) and (6) are still valid with a good approximation.

Another question that may arise is the determination of the critical Cr or Ni content at which the phase change occurs. Using the Cr and Ni values obtained from the weld phase analyses, we performed a second-degree regression.

During the investigation, we utilized all measurement results obtained from EDX phase analyses, regardless of the location. The result of the regression analysis is shown in Figure 18.

In Figure 18, the phase-change boundary can be defined such that a Ni content of 11.29% still corresponds to delta ferrite, whereas a Ni content of 11.73% indicates austenite. This can be interpreted as follows: a Ni content of less than 11.29% always corresponds to delta ferrite, while a Ni content greater than 11.73% always corresponds to austenite.

This relationship requires further analysis to determine the exact boundary. Once the relationship is established, we will be able to approximate the Ni content in a given welding seam based on the Cr content, and vice versa. Even in the absence of knowing the exact boundary, the correlation coefficient is 0.975, which is considered an excellent fit.

6. Conclusions

In this study, the primary objective was to investigate the minimum welding seam requirement for a corrosion-resistant overlay in repairing various pressure vessels. Our tests were performed with a less commonly used wire electrode pairing, aiming to minimize welding time and maximize weld quality. During our tests, we concluded that the buffer layer provides a good microstructure transition between simple carbon steels and corrosion-resistant flux-cored wire types; therefore, a buffer layer is always necessary due to the uniform structure of the subsequent overlay layers. Based on our measurements, the thickness of the fusion zone formed between the base metal and the buffer layer did not exceed 10 microns, sharply separating the corroded and intact material structure. Consequently, we were able to determine that the applied layer structure and welding technology are suitable for the application of material-saving overlay welding.

Accordingly, a suitable corrosion-resistant material structure can be created in the buffer layer (3.9–4.1 mm thickness); however, for durability, at least one cladding layer is needed. The minimum layer thickness, already in the first cladding layer, can be set depending on the thickness of the base metal. In the case of our test, the proper cladding layer thickness is at least 3 mm or more compared to the surface of the basic metal. This distance can be found above the buffer layer, in the first cladding layer, near point 11 given in Figure 11.

We have summarized our main observations and the conclusions that can be drawn from them in the following three points:

• In the case of a uniform-overlay welded material structure, a linear relationship can be determined between the Cr equivalents predicted in the mixture and the measured Cr values. All this is an important aspect due to the expected corrosion-resistance level. The novelty of the relationships is that in both the Schaeffler and WRC-92 methods, Equations (5) and (6) are relationships independent of the layer type; therefore, using these relationships, the expected Cr content can be estimated in the buffer and cladding layers very accurately.
• A quadratic regression relationship can be established between the Cr or Ni content and the phase elements. On the basis of the relationship, the limit of phase element changes in a given system can be estimated (but not defined). The functional relationship allows identifying the phases present based on the measurement of the Cr or Ni content. The exact boundary transition so far is missing, but a transition zone can be determined based on our studies, which falls between 11.29–11.73% based on the Ni content. Below it, only the delta ferrite phase is present, and above it, only the austenite phase is present.
• We observed that the Cr content in the delta ferrite shows stronger changes with the increase in the number of layers. We were able to demonstrate all this numerically based on the data in Table 3 and our measurement results. According to the technical certificate, the Cr content of the E316L filler material is 19%, while according to the measurements, the average Cr content of the cladding layer is 19.63%. At the same time, the filler material of the E309LMo buffer weld had a Cr content of 22.9%, but our measurements indicated an average Cr content of only 18.67% in the buffer layer. Based on the above, we can identify that the Cr was enriched in the delta ferrite towards the surface.

All in all, in terms of reducing the environmental load, we presented an effective overlay welding solution in our study, in which we managed to identify further research directions in the topic, which requires a wider expansion of the studies in the future.