Fatigue Crack Growth Behaviour in Welded Joints of Armour Steel

Abstract

Welded joints are widely recognized as the most critical point in structures made of armour steels due to pronounced thermal effects, microstructural heterogeneity, and the degradation of mechanical and fatigue properties. This study investigates the mechanical properties and fatigue crack growth resistance of a welded joint produced on SA 500 armour steel, with the aim of preserving the properties of the base material as much as possible. To achieve this, a welding procedure incorporating a high-strength filler wire and optimized welding parameters was applied. Hardness and tensile testing was conducted to evaluate the extent of property degradation caused by welding. The results demonstrate that the applied welding process effectively limited the reduction in hardness and tensile strength, achieving values reasonably close to those of the base material. In addition, fatigue crack growth behaviour was investigated in accordance with ASTM E647, using both the Paris law and the McEvily law. The obtained fatigue crack growth curves and threshold stress intensity factor (ΔK_th) values indicate the nearly identical fatigue behaviour of the base material and the heat-affected zone, confirming the successful preservation of base material fatigue behaviour in the thermally affected zone. Moreover, the weld metal exhibited superior resistance to fatigue crack initiation and growth. Overall, the results confirm that the proposed welding approach provides favourable mechanical and fatigue performance for welded joints in armour steel applications.

1. Introduction

In engineering practice, knowledge of the mechanisms of fatigue crack initiation and propagation is extremely important, because they directly affect the reliability and service life of the structure. The timely understanding and prediction of a crack’s behaviour allow sudden breaks to be prevented and reduce the risk of expensive repairs or the replacement of structural parts [1,2,3,4,5].

The fatigue crack problem becomes even more important when it comes to armoured steels. Structures such as combat vehicles (tanks and infantry armoured vehicles) are made of armoured steel, and such structures must be designed to preserve their integrity even under extreme load conditions. Any occurrence of uncontrolled breakage can have consequences for the safety and functionality of combat vehicles. The presence of cracks, most often initiated by the impacts of ballistic projectiles, can shorten the service life of the structure and endanger the crew. By testing fatigue crack growth and determining the parameters that can be obtained from this type of testing, particularly the threshold stress intensity factor range (ΔK_th), it is possible to define the loading level below which crack propagation will not occur and assess how rapidly a crack will grow once this threshold is exceeded. Nevertheless, to assess the remaining service life, the measurement of loads under real service conditions must be conducted. In particular, the load defines the stress intensity factor at the crack tip, which is the deciding factor for crack propagation before it reaches the critical length where the crack becomes unstable.

Particularly critical points in such structures are welded joints and all their zones (base material, heat-affected zone and weld metal). Due to their heterogeneous microstructure and residual stresses introduced by the welding process, welded joints often represent initial fracture sites. Therefore, knowledge of fatigue crack initiation and propagation in the welded joints of armour steels is of crucial importance for a reliable assessment of structural safety [6,7,8,9,10].

Numerous studies confirm that extensive experimental studies have been conducted on various types of steels, as well as their welded joints, with the aim of understanding crack behaviour under fatigue loading. In the literature, special attention should be paid to high-strength (HSS) and armour steels, for which the fatigue crack growth rate is stated as crucial for the reliability of the structure and where the behaviour of all zones of the welded joint is often compared. A. Xu et al. [11] investigated the fatigue crack growth rate in high-strength HG785D steel with and without a “soft + hard + soft” composite welded metal, demonstrating that the heterogeneous weld structure significantly influences FCG behaviour and enhances fatigue resistance through a combination of increased strength and toughness, which is particularly important for the reliability and safety assessment of welded structures. Y. A. Moe et al. [12] investigated the fatigue crack growth behaviour of butt-welded ultra-high-strength steels (S700, S960, and S1100), showing that crack propagation resistance varies significantly across the base metal, weld metal, and heat-affected zone, with the welding method playing a key role, where laser welding provides superior fatigue performance compared to gas metal arc welding, highlighting its importance for the accurate fatigue life prediction of welded UHSS structures. Y. Ma et al. [13] investigated the fatigue crack growth behaviour of Q550E high-strength steel welded joints, demonstrating that the stress ratio significantly influences crack growth parameters and fatigue performance, with welded joints showing higher sensitivity and distinct behaviour compared to the base material. Zhang et al. [14] investigated the fatigue crack growth behaviour of weld-repaired high-strength low-alloy steel, showing that the introduction of a buffer layer significantly improves crack growth resistance, with the tri-metal system (base metal + buffer layer + weld metal) exhibiting better fatigue performance than the bi-metal configuration. X. Shen et al. [15] conducted a study of the fatigue crack’s growth behaviour in different zones of the welded joint of the HSS steel EH690 used in shipbuilding. A comparison of experimentally obtained da/dN − ΔK curves on the welded joints (made of GMAW and EBW) of HSS S960QL was made by R.P.S. Sisodia et al. [16]. The research of A. Čabrilo et al. [17] can be highlighted as the closest to this research, since it was conducted on armour steel Protac 500, which belongs to the same hardness class as the steel used in the present study. The researchers used a GMAW welding process using austenitic stainless steel as a filler material. They achieved comparable metallurgical conditions in the welded joint and experimentally determined the parameters of the fatigue crack growth test and drew the necessary conclusions.

The welding of armour steels is an exceptional technological challenge due to their specific chemical composition and complex heat treatment that achieves strong mechanical properties. As the most critical zone of the welded joint, the HAZ stands out precisely because of its excessive heat input during welding, which leads to the degradation of the base material (microstructural changes and reduction in hardness). Special attention, when discussing reduced hardness, is given to the subcritical HAZ (SCHAZ), which is usually not visible in macroscopic observations but is still affected by the heat from the welding process. It exhibits negligible differences in microstructure and reduced hardness compared to the base material. Such a region, typically 6–10 mm in width, is referred to as the degraded zone [18,19]. One such produced steel is the subject of this article, which is armour steel SA 500. It belongs to the group of ultra-high-strength steels in the hardness class 500 HB. The most common welding process for these steels is the GMAW process with the use of austenitic stainless steels as filler materials with the aim of improving the toughness of the welded joint [20,21]. This paper presents a study on welding using high-strength welding wire, which attempted to approximate the mechanical properties of the base material and the welded joint. By optimizing the welding parameters and controlling the welding speed and heat input, we attempted to minimize the degradation of the HAZ. The goal was also to show that by using such a filler material, it is possible to create a high-quality welded joint in armour steel without compromising its primary function, i.e., impenetrability.

In addition to the mechanical tests performed on the welded joint (microstructural analysis, hardness measurement and tensile test), this work also included the testing of fatigue crack growth. For this purpose, specimens of the base material, the HAZ and the weld metal were analyzed, which enabled a comparison of the fatigue crack growth mechanism in different zones of the welded joint. Based on the obtained experimental parameters and the fatigue crack growth curves according to the Paris and McEvily laws, comprehensive insight into the influence of welding using the aforementioned wire on the fatigue behaviour of SA 500 armour steel was provided.

2. Materials and Methods

2.1. Base Material

Welding and experimental analyses were performed on armour steel SA 500 (Swebor Armour^TM 500, Swebor Armour, Luelå, Sweden). This ballistic protection steel with low carbon and manganese content combined with controlled heat treatment provides a good combination of hardness, the R_p0.2/R_m ratio, weldability and ballistic protection. It is widely used in military and civilian purposes (SUVs and trucks as armoured vehicles, security doors and walls, bank safes, etc.). Its chemical composition is shown in

Table 1. Chemical composition of armour steel SA 500.

Element	C	Si	Mn	S + P	Mo	Cr	Cu	Ni	Ti
wt.%	0.30	0.50	0.40	0.02	0.20	0.89	0.03	0.05	0.03

, while the minimum mechanical properties guaranteed by the manufacturer are shown in

Table 2. Minimum mechanical properties of armour steel SA 500.

Yield Strength Rp0.2 (MPa)	Tensile Strength Rm (MPa)	Elongation A5 (%)	Impact Strength KV −40 °C (J)
1200	1670	9.5	25

[22].

2.2. Filler Material

High-strength welding wire ED-FK 1000 (Hermann Fliess & Co. GmbH, Duisburg, Germany) was used as a filler material. This MIG/GMAW wire is normally used for welding high-strength low-alloy steels, and it shows very good impact toughness up to −60 °C. It has a wide range of applications such as in mobile cranes, shipbuilding and the advanced automotive industry [23]. The classification of the wire according to DIN EN ISO and AWS standards, as well as the chemical composition, is presented in

Table 3. Chemical composition of filler material ED-FK 1000.

ED-FK 1000	DIN EN ISO 16834 [24]–A G 89 M21 Mn4Ni2CrMo AWS A5.28 ER120S–G
Element	C	Si	Mn	S + P	Mo	Cr	B	Ni	Ti
wt.%	0.10	0.78	1.79	0.10	0.57	0.36	0.01	2.29	0.08

. The mechanical properties of the wire are shown in

Table 4. Mechanical properties of filler material ED-FK 1000.

Yield Strength Rp0.2 (MPa)	Tensile Strength Rm (MPa)	Elongation A5 (%)	Impact Strength KV −60 °C (J)
960	1000	15.0	47

.

2.3. Welding

The GMAW process was used to weld two plates, dimensions 300 mm × 150 mm and 8 mm thick, made of armour steel SA 500, in shielding gas M21 (82% Ar + 18% CO₂). Welding was performed on an OTC FD-V8L robot (OTC Daihen Europe, Mönchengladbach, Germany), as shown in Figure 1, with a WelBee P500L welding power source (Daihen Varstroj d.d., Lendava, Slovenia). The goal of welding was to reduce the heat input into the welded joint by applying controlled parameters, that is, to degrade the HAZ as little as possible.

The geometry of the prepared butt joint is shown in Figure 2a, the welding pass sequence is presented in Figure 2b, and the welding parameters used are presented in

Table 5. Applied welding parameters.

Pass	Welding Current I (A)	Arc Voltage U (V)	Welding Speed Vw (cm/min)	Heat Input Q (kJ/mm)
1	110	17.2	25	0.45
2	117	16.5	25	0.46
3	193	17.1	36	0.55
4	210	19.4	36	0.68
5	210	19.1	36	0.67
Shielding gas: 18 L/min M21 (82% Ar + 18% CO2)

. The half-V groove was selected in order to obtain the fusion line perpendicular to the surface in order to be able to locate the notch in only one subzone of the HAZ (Figure 3).

2.4. Testing of Mechanical Properties

After welding, the specimens for testing were extracted from the welded plate by water jet cutting. The final specimen dimensions were achieved through additional grinding. Subsequently, the required notches for fatigue crack growth testing were introduced by wire electrical discharge machining (EDM). The position of specimens in the welded plate is illustrated in Figure 3.

From the welded joint, in order to test the mechanical properties, the necessary samples/specimens were obtained to determine the microstructure, measure hardness and test the tensile properties. To test the microstructure and measure hardness, the same, but specially prepared, sample was used. The sample was previously ground and polished and then etched with a composition of 3% nital. Images of the microstructure were acquired from each zone of the welded joint. The hardness measurement was performed according to the ISO 9015-2 standard [25], using the Vickers HV 10 method, with measurements being carried out in three lines along the entire sample: the weld root line, the weld middle line and the weld face line. A Nikon Epiphot 300 microscope (Nikon Metrology Inc., Brighton, MI, USA), as shown in Figure 4a, was used for microstructure analysis, while the hardness measurement was performed on a Roell-Zwick Z600 machine (Zwick, Roell Group, Ulm, Germany), as shown in Figure 4b.

A tensile test was conducted to determine the tensile properties of the base material and the welded joint, namely yield stress, tensile strength, elongation and cross-section contraction. For this experiment, according to EN ISO 15614-1 [26], two specimens with a rectangular cross-section were made from both the base material and the welded joint, with the geometry shown in Figure 5a. The test was carried out according to the EN ISO 6892-1 standard (base metal) [27] and EN ISO 5178 (weld joint) [28] on an Amsler 559/594 tensile machine (Zwick Roell Group, Ulm, Germany), as shown in Figure 5b.

2.5. Fatigue Crack Growth Test

The fatigue crack growth rate is one of the key parameters in fracture mechanics because it directly affects the life of the structure, especially under variable load conditions. In the construction of armour steel, fatigue crack growth is of great importance, because the high reliability, fatigue resistance and controlled behaviour of the material during a long-term exploitation period are expected in such constructions.

Experimental fatigue crack growth testing was performed according to ASTM E647 [29], using an SENB (Single-Edge Notched Bending) specimen. The test was performed on a Rumul Cracktronic testing device (Russenberger Prüfmaschinen AG, Neuhausen, Switzerland), as shown in Figure 6. The geometry of the SENB specimen, which is widely accepted for characterizing the fatigue behaviour of materials under bending, is shown in Figure 7a. To monitor crack growth during the test, RMF-A5 crack gauges (Russenberger Prufmachinen AG, Neuhausen, Switzerland) were used, which enable the reliable control of crack propagation through the registration of the elongation and successive tearing of conductive gauge segments. The use of these crack gauges enables a measurement accuracy of 1 μm, with possible deviations of up to ±0.5 μm during crack growth monitoring due to the data acquisition system and testing conditions. Calibration was performed in accordance with the manufacturer’s guidelines. The crack gauge used on one of the specimens is shown in Figure 7b. This enabled a continuous correlation between the number of cycles and the crack length.

Starting from the initial crack length (a₀), it is possible to determine the current crack length (a_i), which is defined by Equation (1), using a crack gauge. In the interval between two consecutive measurement points, the fatigue crack growth rate is approximated by Equation (2):

$$a_i=a_0+x$$

(1)

$${\left({\displaystyle \frac{da}{dN}}\right)}_i={\displaystyle \frac{a_{i+1}-a_i}{N_{i+1}-N_i}}$$

(2)

where a₀ is the initial crack length; a_i is the current crack length; x is the recorded crack propagation on the crack gauge; da/dN is the fatigue crack growth rate; and N is the number of fatigue loading cycles.

In fatigue crack growth tests, the stress intensity factor range (ΔK) is used, which is defined as the difference between the maximum (K_max) and minimum (K_min) values of the stress intensity factor during one loading cycle, as shown in Equation (3):

$$∆K=K_{max}-K_{min}$$

(3)

For the SENB specimen, the general expression for the stress intensity factor can be written as Equation (4):

$$∆K={\displaystyle \frac{∆F}{B\sqrt{W}}}·f\left({\displaystyle \frac{a_i}{W}}\right)$$

(4)

where ΔF is the load range, B is the specimen thickness, W is the specimen height, and f(a_i/W) is a dimensionless geometry function for the SENB specimen.

The load ratio factor (R) represents the ratio of the minimum and maximum values of the stress intensity factor in one fatigue load cycle and is defined as Equation (5):

$$R={\displaystyle \frac{K_{min}}{K_{max}}}$$

(5)

In this study, the applied load ratio factor was R = 0.1, which means that the minimum load was 10% of the maximum load during each cycle. This value of the R factor is most often used in fatigue testing because it allows for stable crack growth and reduces the possibility of complete crack closure during the cycle. This experiment was conducted in the air, at room temperature (20 °C) with a relative humidity of 50%. Since this experiment was performed on a resonant testing machine, changes in frequency and load are directly dependent on crack growth. As the crack length increases, frequency decreases. The initial frequency for all specimens was 52 Hz; during testing, with increasing crack length, it varied within the range of 45–52 Hz, while the frequency at the final fracture was 40 Hz.

The aim of this test was to describe fatigue crack growth in the form of the Paris and McEvily laws. Paris’s law is shown by Equation (6), while McEvily’s law is shown by Equation (7):

$$\left({\displaystyle \frac{da}{dN}}\right)=C·{\left(∆K\right)}^m$$

(6)

$$\left({\displaystyle \frac{da}{dN}}\right)=C_1·{\left(∆K-{∆K}_{th}\right)}^{m_1}$$

(7)

where da/dN is fatigue crack growth; C and C₁ represent the intersection of the Paris curve with the y-axis; m and m₁ represent the slope of the Paris curve; ΔK_th is the threshold stress intensity factor range.

The results of the fatigue test are represented by the Paris curve in the logarithmic diagram of fatigue crack growth (da/dN) depending on the range of the stress intensity factor (ΔK). Three characteristic regions of crack growth can be clearly defined on the Paris curve: the region near the crack growth threshold (ΔK_th), the region of stable crack growth and the region of unstable crack growth. The region near the crack growth threshold is characterized by very slow or negligible fatigue crack growth, where the ΔK values are not sufficient to maintain continuous crack propagation. The region of stable crack growth is characterized by an almost linear dependence between log(da/dN) and log(ΔK), where the crack grows in a controlled and predictable manner. The region of unstable crack growth occurs at high ΔK values, when there is a sudden acceleration of crack growth and when approaching the final fracture of the material [30,31,32].

Paris’s law is applicable exclusively to the second region, i.e., the region of stable crack growth, while McEvily’s law also takes into account the influence of the crack growth threshold, which enables the description of crack behaviour in the first region as well. The Paris curve with clearly defined regions of crack growth is shown in Figure 8.

For the determination of the Paris curve parameters for both laws, specimens from each zone of the welded joint SA 500 armour steel were used.

3. Results

3.1. Results of Mechanical Property Tests

For a clearer overview of the zones of the welded joint for which the microstructure images are shown, as well as the method of measuring the hardness in three lines (since the microscopic analysis and hardness measurement were performed on the same sample), a macroscopic image of the welded joint is shown in Figure 9. This macroscopic image was taken at 10× magnification and allows for the clear identification of the base material, the heat-affected zone and the welded joint.

Microscopic analysis was carried out at a magnification of 500× and 1000× in order to characterize the microstructure of all zones of the welded joint and measure grain sizes. Figure 9 presents the microstructural details of the base material, HAZs, and weld metal. Detail 1 (Figure 10a and Figure 11a) shows the microstructure of the SA 500 base material, where a clearly visible tempered martensitic microstructure is observed, formed as result of the heat treatment applied to this steel. Detail 2 (Figure 10b and Figure 11b) illustrates the subcritical HAZ (SCHAZ), i.e., the part of the HAZ that is not clearly distinguishable on the macroscopic image. This zone is often referred to as the degraded zone, where the influence of the welding thermal cycle is still reflected through minor microstructural changes. The microstructure in this degraded zone is also tempered martensite but with slight deviation compared to the base material (corresponding to a temperature range below A_c1), where a mild reduction in hardness is expected. Detail 3 (Figure 10c and Figure 11c) presents the intercritical HAZ (ICHAZ), which is characterized by a mixed microstructure. The microstructure consists of freshly formed martensite originating from intercritical austenitic regions and retained tempered martensite from the base material (visible as a contrast difference in colour). This heterogeneous zone is typical for the temperature interval between A_c1 and A_c3. Detail 4 (Figure 10d and Figure 11d) shows the coarse-grained HAZ (CGHAZ), where coarse martensitic grains are present due to the strong influence of welding heat input. This zone is typical for the temperature interval over A_c3. Detail 5 (Figure 10e and Figure 11e) presents the weld metal microstructure taken from the weld face. The microstructure is relatively homogeneous and consists of a martensitic–bainitic structure. Pronounced directional features are observed, resulting from the rapid cooling of the weld and chemical composition of the filler material. Detail 6 (Figure 10f and Figure 11f) presents the microstructure of the weld root. It is located in weld pass No. 5 (the last pass made) and was not reheated by the welding of subsequent passes. The microstructure is very similar to the microstructure of the weld crown.

The hardness measurement results obtained using the Vickers HV 10 method along three measurement lines were converted into Brinell hardness values (HB), as the HB scale is specified by relevant standards for armour steels. The results are presented in the form of a diagram in Figure 12, which clearly illustrates the hardness distribution across all zones of the welded joint.

The average hardness of individual zones is presented in

Table 6. Results of hardness measurements.

Zone of the Welded Joint	Average Hardness (HB)	Minimum Value of Hardness (HB)	Maximum Value of Hardness (HB)	Standard Deviation (HB)
Base material	505	485	513	12.12
Degraded zone	430	304	475	44.36
HAZ	325	266	399	33.77
Weld metal	347	323	390	19.83

.

Table 7. Correlation between microstructure, grain size and hardness of all zones of welded joint.

Zone of the Welded Joint	Microstructure	Grain Size (μm)	Hardness (HB)
Base material	tempered martensite	20–27	505
SCHAZ	tempered martensite	21–26	430
ICHAZ	freshly formed martensite from intercritical austenitic regions + retained tempered martensite from the base material	35–52	339
CGHAZ	coarse-grained freshly formed martensite	140–230	321
Weld metal	martensite–bainite	12–19	347

presents the correlation between the characterized microstructure (including the measured grain size) and the measured hardness for all zones of the welded joint, including all HAZ subregions.

The tensile test results are presented in

Table 8. Results of tensile test.

Specimen	Yield Strength Rp0.2 (MPa)	Tensile Strength Rm (MPa)	Elongation A5 (%)	Cross-Section Contraction Z (%)
TT-BM-1	1427	1578	7.5	21.0
TT-BM-2	1415	1532	7.3	23.0
Average BM	1421	1555	7.4	22.0
TT-WM-1	1022	1063	4.0	13.0
TT-WM-2	904	950	4.4	13.0
Average WM	963	1006	4.2	13.0

, while the fractured specimens are shown in Figure 13.

3.2. Results of Fatigue Crack Growth Test

The fatigue crack growth test results are presented in the form of diagrams, where the first diagram (Figure a) illustrates the Paris law, while the second diagram (Figure b) illustrates the McEvily law. Figure 14 presents the results for the base material, Figure 15 for the HAZ, and Figure 16 for the weld metal. The parameters of the Paris and McEvily laws are given in

Table 9. Results of Paris’s and McEvily’s laws.

Specimen From	Paris’s Law			McEvily’s Law
	C (-)	m (-)	R2	C1 (-)	m1 (-)	ΔKth (MPa∙m1/2)
Base material	2.764 ∙ 10−11	2.651	0.895	8.777 ∙ 10−10	1.594	2.79
HAZ	1.339 ∙ 10−11	3.037	0.842	8.223 ∙ 10−10	1.707	2.80
Weld metal	1.758 ∙ 10−12	3.457	0.910	1.761 ∙ 10−10	2.065	4.21

.

4. Discussion

In this section, the results of the fatigue crack growth test were analyzed, where a comparison was made of both the fatigue crack growth curves and the thresholds for the start of crack propagation (ΔK_th) for the base material (BM), the heat-affected zone (HAZ) and the weld metal (WM). The comparison is shown in Figure 17.

The ΔK_th threshold value for the base material is 2.79 MPa∙m^1/2, while for the HAZ, it is 2.80 MPa∙m^1/2, which indicates almost identical resistance to initiation and early crack growth in these zones. This coincidence represents a very favourable result, considering the fact that the HAZ is often the most critical point of the welded joint due to the microstructural changes caused by the welding thermal cycle. The obtained results confirm that the chosen welding procedure and the filler material enabled the preservation of the fatigue properties of the base material in the HAZ, which was one of the main goals of this research. By comparing the fatigue crack growth curves for the base material and HAZ, very similar behaviour can be observed in the region of stable crack growth, which additionally confirms that the thermal influence of welding did not lead to a significant deterioration of the resistance of the HAZ to crack propagation. In contrast, the weld metal shows a significantly higher value of the threshold ΔK_th of 4.21 MPa∙m^1/2, which indicates better resistance to crack formation compared to the base material and HAZ. This behaviour of the weld metal can be attributed to its relatively homogeneous fine-grained martensitic–bainitic microstructure and corresponding mechanical properties. As shown in Table 7 of this study, the weld metal (i.e., the deposited filler material—high-strength welding wire) exhibits the finest grain size among all zones of the welded joint. Small grain size is known to contribute to higher ΔK_th values, thereby improving resistance to fatigue crack growth. In addition, the filler material, due to its lower alloying level, exhibits lower hardness compared to the base material, which further contributes to an increased ΔK_th. In the comparison between the base material and the HAZ, only minor differences in grain size were observed (between BM and ICHAZ), which explains the nearly identical ΔK_th values obtained for these regions. Residual stresses, which can also influence fatigue crack growth behaviour, were not the primary focus of this study. The welding technology was designed to make residual stress as small as possible. Therefore, all passes were welded with very low heat input (below 0.7 kJ/mm), and double-side welding was applied. However, the residual stresses were measured using three different methods, and only their average values are reported here for completeness. The measured residual stress ranges (difference between longitudinal and transverse stresses) were approximately −50 MPa to 70 MPa in the base material, 115 MPa to 300 MPa in the HAZ, and −50 MPa to 140 MPa in the weld metal. It should be noted that specimen preparation leads to the partial relaxation of residual stresses; therefore, their direct influence on fatigue crack growth behaviour in the tested specimens is considered limited. A more detailed analysis of residual stresses will be presented in a separate publication. By analyzing the fatigue crack growth curves, it can be concluded that the weld metal, in addition to a higher crack initiation threshold, also exhibits more favourable behaviour in the region of stable crack growth (lower curve on the diagram), which additionally confirms that the applied welding procedure did not negatively affect the fatigue resistance of the welded joint. Overall, the results of the fatigue crack growth test clearly indicate that the properties of the base material in the HAZ are successfully preserved, while the weld metal even shows improved resistance to fatigue crack initiation.

A comparison with previous studies, particularly the work of Čabrilo et al. [17] on Protac 500 armour steel, shows that significantly higher ΔK_th values were obtained for the base material (13.4 MPa·m^1/2), indicating a higher resistance to fatigue crack growth compared to the base material investigated in the present study. However, their results also revealed a reduction in ΔK_th in both the HAZ (12.6 MPa·m^1/2) and especially the weld metal (10.1 MPa·m^1/2), which was associated with the use of an austenitic filler material. In contrast, although the absolute ΔK_th values obtained in this study are lower, the applied controlled welding procedure with reduced heat input enabled the preservation of fatigue crack growth resistance in the HAZ, maintaining values comparable to the base material. Moreover, the use of a high-strength filler material resulted in improved ΔK_th in the weld metal, which is particularly important since the weld region is generally considered the most critical part of armour steel structures. These results emphasize the significant influence of welding parameters and filler material selection on fatigue crack growth behaviour, especially in achieving improved performance in the weld metal while maintaining the integrity of the HAZ.

5. Conclusions

According to numerous studies, the welded joint is considered the most critical point in structures made of armour steels, given the pronounced microstructural changes and reduction in mechanical and fatigue properties due to the thermal cycle of welding. For this reason, the presented study was conducted on SA 500 armour steel with the aim of assessing the influence of the applied welding process on mechanical properties and resistance to fatigue crack growth.

The applied welding procedure, which includes the introduction of a high-strength filler wire and optimized welding parameters, showed extremely favourable results in terms of preserving the mechanical properties of the base material. Hardness measurements showed that the base material has a hardness of 505 HB, while in the heat-affected zone (HAZ) and in the weld metal, values of 325 HB and 347 HB were measured, respectively. This corresponds to a hardness reduction of approximately 36% in the HAZ and 31% in the weld metal compared to the base material, which, considering the nature of armour steels and their sensitivity to welding, can be considered a relatively small and technically acceptable decrease.

The tensile test results further confirm the efficiency of the applied welding process. The base material showed a tensile strength of 1555 MPa, while in the welded joint, a tensile strength of 1006 MPa was measured, which corresponds to a reduction of about 35%. Although a drop in tensile strength is inevitable in the welded joints of ultra-high-strength steels, the achieved tensile strengths indicate that the optimized parameters and high-strength wire significantly contributed to the reduction in mechanical property degradation.

Fatigue crack growth analysis showed that the behaviour of the base material in the HAZ was successfully preserved. The threshold values for crack initiation and early growth (ΔK_th) for the base material and the HAZ are 2.79 MPa∙m^1/2 and 2.80 MPa∙m^1/2, respectively, indicating the almost identical behaviour of these zones in terms of resistance to crack initiation. This result represents a very significant achievement, since the HAZ is usually the most sensitive zone of a welded joint.

At the same time, the weld metal shows improved fatigue behaviour, with a significantly higher ΔK_th threshold value of 4.21 MPa∙m^1/2, which indicates a higher resistance to crack formation compared to the base material and HAZ. This can be related to the relatively homogeneous martensitic–bainitic microstructure of the weld metal, as well as the fine grain size measured in the weld metal, which contributes to a higher ΔK_th, as well as the favourable mechanical properties resulting from the chemical composition of the filler material and the cooling conditions.

Based on the obtained results, it can be concluded that the applied welding procedure, with the use of high-strength wire and optimized parameters, enables the preservation of the fatigue behaviour of the base material in the HAZ, as well as the achievement of the even better behaviour of the weld metal in terms of resistance to fatigue crack growth. By employing this welding procedure and the obtained fatigue crack growth parameters, it is possible to evaluate the service life of such a welded joint (including all its zones) in the presence of cracks, which represent the most common type of defect in armour steel structures and typically arise under impact loading in real service conditions. This approach represents a very promising solution for welding armour steel in constructions where safety, reliability and resistance to damage are of crucial importance.