Research on the Properties and Low Cycle Fatigue of Sc-Modified AA2519-T62 FSW Joint

The aim of this research was to examine the mechanical and fatigue properties of friction stir welded Sc-modified 5 mm thick AA2519-T62 extrusion. The joint was obtained using the following parameters: 800 rpm tool rotation speed, 100 mm/min tool traverse speed, 17 kN axial, and MX Triflute as a tool. The investigation has involved microstructure observations, microhardness distribution analysis, tensile test with digital image correlation technique, observations of the fracture surface, measurements of residual stresses, low cycle fatigue testing, and fractography. It was stated that the obtained weld is defect-free and has joint efficiency of 83%. The failure in the tensile test occurred at the boundary of the thermo-mechanically affected zone and stir zone on the advancing side of the weld. The residual stress measurements have revealed that the highest values of longitudinal stress are localized at the distance of 10 mm from the joint line with their values of 124 MPa (the retreating side) and 159 MPa (the advancing side). The results of low cycle fatigue testing have allowed establishing of the values of the cyclic strength coefficient (k′ = 504.37 MPa) and cyclic strain hardening exponent (n′ = 0.0068) as well as the factors of the Manson–Coffin–Basquin equation: the fatigue strength coefficient σ′f = 462.4 MPa, the fatigue strength exponent b = −0.066, the fatigue ductility coefficient ε′f = 0.4212, and the fatigue ductility exponent c = −0.911.


Introduction
Friction stir welding (FSW) is a very efficient technology in the production of aluminum alloy joints, as it provides a number of advantages over traditional fusion welding such as the lower temperature of the joining process [1][2][3]. This particular factor is crucial when it comes to welding of high-strength aluminum alloys, which are mainly precipitated-hardening materials (2XXX and 7XXX series) and their specific strength is an effect of thermally unstable precipitates (e.g., Guinier-Preston zones, θ phase) [4][5][6][7]. Considering the production of efficient welds of these materials, the losses in hardening must be taken into account due to the thermal affection of the joining process on the workpiece. In this paper, AA2519-T62 armor grade aluminum is taken under investigation in terms of its friction stir-welded joint properties. This alloy contains 5.3-6.4% copper and is subjected to the precipitation hardening process, and thus, acquires high specific strength, which makes it a very desirable material in terms of aerospace, automotive, and military applications [8][9][10]. It is noteworthy that the investigated alloy is a modification of AA2519 characterized by the addition of scandium and zirconium, which improves its mechanical properties and resistance to elevated temperature in the form of higher recrystallization temperature [9,[11][12][13]. Although friction stir welding is a suitable  Table 2. Mechanical properties of AA2519-T62 [34].

Young Modulus (E) Yield Strength (R e0.2 ) Tensile Strength (R m ) Elongation (A)
78 GPa 312 MPa 469 MPa 19% Prior to the joining process, the workpieces to be welded were ground and cleaned with isopropanol. The FSW was conducted with an ESAB FSW Legio 4UT (ESAB, Gothenburg, Sweden) machine with the applied axial force of 17 kN. The tool rotation speed and tool traverse speed were equal to 800 rpm and 100 mm/min, respectively. The type of tool used was an MX Triflute (ESAB, Gothenburg, Sweden) and the tilt angle was set to 2 • . These parameters were selected based on the previous study performed by authors [34]. The obtained joint was sectioned perpendicular to the welding direction. The metallographic preparation involved cutting a sample from the joint, mounting it in resin, grinding it with abrasive papers of 80, 320, 500, 800, 1200, 2400, and 4000 gradations, and polishing with diamond pastes (gradations of 3 and 1 µm). The prepared sample was etched with Keller reagent (20 mL H 2 O + 5 mL 63% HNO 3 + 1 mL 38% HCl + one drop of 40% HF) for about 10 s. The microstructure observations have been conducted on the Olympus LEXT OLS 4100 digital light microscope (Olympus, Tokyo, Japan). The Vickers microhardness (Struers Copenhagen, Denmark) of the weld was measured on its polished cross-section by applying 0.98 N load in accordance with the EN ISO 6507 standard. The distribution of microhardness was obtained for the top, middle, and bottom parts of the cross-section of the weld: 0.8, 2.5, and 4.4 mm from the face of the joint, respectively. The basic mechanical properties of the joint were examined by tensile testing according to ASTM standard E8/E8M-13a [38]. Tensile tests were carried out on an Instron 8802 MTL universal testing machine (Instron, Norwood, MA, USA) with WaveMatrix computer software (Instron, Norwood, MA, USA). Additionally, the tensile test was supported by digital image correlation (DIC) in order to examine the local strain in the welded joint zone. For DIC, the Dantec Q-400 optical system (Dantec Dynamics GmbH, Ulm, Germany) was used, and the obtained data were processed with Istra4D 4.4.7 software (version 4.4.7). The fatigue testing was conducted on an Instron 8802 servohydraulic fatigue testing system according to ASTM E606/E606M standard. To measure the value of strain during testing a 2520-603 dynamic extensometer (Instron, Norwood, MA, USA) with a gauge length of 25 mm was used. The fatigue tests were carried out on five various levels of total strain amplitude: 0.35%, 0.4%, 0.5%, 0.6%, and 0.8% with strain ratio R = 0.1. For each level, three samples were examined. The fracture surfaces of tensile and fatigue samples were analyzed on the scanning electron microscope (SEM) Jeol JSM-6610 (Jeol, Tokyo, Japan). The samples for tensile and fatigue testing were prepared with the geometry presented in Figure 1.

Young Modulus (E) Yield Strength (Re0.2) Tensile Strength (Rm) Elongation (A)
78 GPa 312 MPa 469 MPa 19% Prior to the joining process, the workpieces to be welded were ground and cleaned with isopropanol. The FSW was conducted with an ESAB FSW Legio 4UT (ESAB, Gothenburg, Sweden) machine with the applied axial force of 17 kN. The tool rotation speed and tool traverse speed were equal to 800 rpm and 100 mm/min, respectively. The type of tool used was an MX Triflute (ESAB, Gothenburg, Sweden) and the tilt angle was set to 2°. These parameters were selected based on the previous study performed by authors [34]. The obtained joint was sectioned perpendicular to the welding direction. The metallographic preparation involved cutting a sample from the joint, mounting it in resin, grinding it with abrasive papers of 80, 320, 500, 800, 1200, 2400, and 4000 gradations, and polishing with diamond pastes (gradations of 3 and 1 μm). The prepared sample was etched with Keller reagent (20 mL H2O + 5 mL 63% HNO3 + 1 mL 38% HCl + one drop of 40% HF) for about 10 s. The microstructure observations have been conducted on the Olympus LEXT OLS 4100 digital light microscope (Olympus, Tokyo, Japan). The Vickers microhardness (Struers Copenhagen, Denmark) of the weld was measured on its polished cross-section by applying 0.98 N load in accordance with the EN ISO 6507 standard. The distribution of microhardness was obtained for the top, middle, and bottom parts of the cross-section of the weld: 0.8, 2.5, and 4.4 mm from the face of the joint, respectively. The basic mechanical properties of the joint were examined by tensile testing according to ASTM standard E8/E8M-13a [38]. Tensile tests were carried out on an Instron 8802 MTL universal testing machine (Instron, Norwood, MA, USA) with WaveMatrix computer software (Instron, Norwood, MA, USA). Additionally, the tensile test was supported by digital image correlation (DIC) in order to examine the local strain in the welded joint zone. For DIC, the Dantec Q-400 optical system (Dantec Dynamics GmbH, Ulm, Germany) was used, and the obtained data were processed with Istra4D 4.4.7 software (version 4.4.7). The fatigue testing was conducted on an Instron 8802 servohydraulic fatigue testing system according to ASTM E606/E606M standard. To measure the value of strain during testing a 2520-603 dynamic extensometer (Instron, Norwood, MA, USA) with a gauge length of 25 mm was used. The fatigue tests were carried out on five various levels of total strain amplitude: 0.35%, 0.4%, 0.5%, 0.6%, and 0.8% with strain ratio R = 0.1. For each level, three samples were examined. The fracture surfaces of tensile and fatigue samples were analyzed on the scanning electron microscope (SEM) Jeol JSM-6610 (Jeol, Tokyo, Japan). The samples for tensile and fatigue testing were prepared with the geometry presented in Figure 1.  The values of residual stresses were established by the hole-drilling method which is described in ASTM E 837 in detail [39]. Because of the thickness of the plate (5 mm), the blind-hole procedure was used. In this approach, the hole is drilled at the depth of 2 mm and the released strains are measured by the specially designed rosette. The accurate location of the hole in the center point of the rosette was ensured by using the VPG Micro-Measurement System RS-200 milling guide (Micro-Measurements, Raleigh, NC, USA), which is presented in Figure 2a. A constant value of Young's modulus was used during the analysis, which equals 78 GPa. The measurements were made in 10 points located at different distances of the joint axis: two in the joint axis, two at the weld line, and six in the vicinity of the joint, both at the advancing side and retreating side. The schema of measuring point locations is presented in Figure 2b. Three measuring points had to be shifted from others during mounting in order to make the possibility of proper installation of rosettes. The values of residual stresses were established by the hole-drilling method which is described in ASTM E 837 in detail [39]. Because of the thickness of the plate (5 mm), the blind-hole procedure was used. In this approach, the hole is drilled at the depth of 2 mm and the released strains are measured by the specially designed rosette. The accurate location of the hole in the center point of the rosette was ensured by using the VPG Micro-Measurement System RS-200 milling guide (Micro-Measurements, Raleigh, NC, USA), which is presented in Figure 2a. A constant value of Young's modulus was used during the analysis, which equals 78 GPa. The measurements were made in 10 points located at different distances of the joint axis: two in the joint axis, two at the weld line, and six in the vicinity of the joint, both at the advancing side and retreating side. The schema of measuring point locations is presented in Figure 2b. Three measuring points had to be shifted from others during mounting in order to make the possibility of proper installation of rosettes. Strains revealed during drilling and measured at the individual gauges of the rosette enabled determining of the values of principle stresses σmax and σmin together with their directions. Obtained results were used to calculate the values of stresses, longitudinal σ11, and transverse σ22, connected with the FSW joint. The error of computed values of residual stresses was estimated at ±(8-10) MPa in the whole range, which is acceptable. All tests mentioned in the experimental part were carried out 4 weeks after the joining process in order to stabilize the properties of the weld [2].

Macroscopic Observations
The weld face of the obtained joint is presented in Figure 3. For FSW joints, two sides of the weld can be distinguished: the retreating side (the direction of tool rotation is opposite to the welding direction) and the advancing side (the direction of tool rotation is concordant with the welding direction) [2]. The weld face is characterized by a typical, regular weld track, formed by the tool shoulder. As can be observed, no visible defects are present and the flash is localized mainly on the retreating side of the weld. Strains revealed during drilling and measured at the individual gauges of the rosette enabled determining of the values of principle stresses σ max and σ min together with their directions. Obtained results were used to calculate the values of stresses, longitudinal σ 11 , and transverse σ 22 , connected with the FSW joint. The error of computed values of residual stresses was estimated at ±(8-10) MPa in the whole range, which is acceptable. All tests mentioned in the experimental part were carried out 4 weeks after the joining process in order to stabilize the properties of the weld [2].

Macroscopic Observations
The weld face of the obtained joint is presented in Figure 3. For FSW joints, two sides of the weld can be distinguished: the retreating side (the direction of tool rotation is opposite to the welding direction) and the advancing side (the direction of tool rotation is concordant with the welding direction) [2]. The weld face is characterized by a typical, regular weld track, formed by the tool shoulder. As can be observed, no visible defects are present and the flash is localized mainly on the retreating side of the weld. The light microscopy image of the macroscopic view of the joint cross-section is presented in Figure 4. Lack of visible defects confirms that parameters used are appropriate for friction stir welding of this alloy. The width of the stir zone is about 8.5 mm. The welding process has reduced the thickness of the workpiece from 5 mm to 4.5 mm in the center of the weld. Generally, in FSW joint the two zones and one subzone can be distinguished: thermo-mechanically affected zone (TMAZ) formed due to effects of temperature and severe plastic deformation, stir zone (SZ), a specific subregion of TMAZ localized in the central part of a joint, characterized by the presence of ultrafine, dynamically recrystallized grains, and heat-affected zone (HAZ) affected only by the heat of the process. The observations revealed the presence of onion ring patterns in the stir zone, visible mainly at the advancing side of the weld. The differences in the macrostructure of the TMAZ/SZ boundaries at the advancing and retreating sides have been taken under investigation during microstructure analysis.

Microstructure Analysis and Microhardness Measurements
The microstructure of the TMAZ/SZ interface at the advancing side of the joint is characterized by the clear border between dynamically recrystallized grains of the stir zone and elongated grains of the thermo-mechanically affected zone, which visualize the flow of the material during the stirring process ( Figure 5a). Concurrently, at the retreating side, the border between SZ and TMAZ is of transition nature (Figure 5b). The microstructure of TMAZ in close proximity to SZ can be identified as bands of recovered grains irregularly separated by recrystallized grains. These differences between each side have their source in the unsymmetrical distribution of the temperature during the FSW process. Studies revealed that the retreating side is characterized by higher effect of the temperature [40,41]. During welding, the generated heat influences the plastically deformed microstructure by promoting heat-activated phenomena such as grain recovery and recrystallization. As can be seen, the dynamic recovery has a predominant role while fine, dynamically recrystallized grains are localized mainly between large, deformed grains in the form of long bands (Figure 5b). This suggests that the grains which undergo dynamic recrystallization are compressed and intensely elongated by the larger grains in the stirring process and the quantity of heat allows them to rebuild their severely deformed microstructures. This phenomenon has not been observed at the advancing side ( Figure  5a). The light microscopy image of the macroscopic view of the joint cross-section is presented in Figure 4. Lack of visible defects confirms that parameters used are appropriate for friction stir welding of this alloy. The width of the stir zone is about 8.5 mm. The welding process has reduced the thickness of the workpiece from 5 mm to 4.5 mm in the center of the weld. Generally, in FSW joint the two zones and one subzone can be distinguished: thermo-mechanically affected zone (TMAZ) formed due to effects of temperature and severe plastic deformation, stir zone (SZ), a specific subregion of TMAZ localized in the central part of a joint, characterized by the presence of ultrafine, dynamically recrystallized grains, and heat-affected zone (HAZ) affected only by the heat of the process. The observations revealed the presence of onion ring patterns in the stir zone, visible mainly at the advancing side of the weld. The differences in the macrostructure of the TMAZ/SZ boundaries at the advancing and retreating sides have been taken under investigation during microstructure analysis.  The light microscopy image of the macroscopic view of the joint cross-section is presented in Figure 4. Lack of visible defects confirms that parameters used are appropriate for friction stir welding of this alloy. The width of the stir zone is about 8.5 mm. The welding process has reduced the thickness of the workpiece from 5 mm to 4.5 mm in the center of the weld. Generally, in FSW joint the two zones and one subzone can be distinguished: thermo-mechanically affected zone (TMAZ) formed due to effects of temperature and severe plastic deformation, stir zone (SZ), a specific subregion of TMAZ localized in the central part of a joint, characterized by the presence of ultrafine, dynamically recrystallized grains, and heat-affected zone (HAZ) affected only by the heat of the process. The observations revealed the presence of onion ring patterns in the stir zone, visible mainly at the advancing side of the weld. The differences in the macrostructure of the TMAZ/SZ boundaries at the advancing and retreating sides have been taken under investigation during microstructure analysis.

Microstructure Analysis and Microhardness Measurements
The microstructure of the TMAZ/SZ interface at the advancing side of the joint is characterized by the clear border between dynamically recrystallized grains of the stir zone and elongated grains of the thermo-mechanically affected zone, which visualize the flow of the material during the stirring process ( Figure 5a). Concurrently, at the retreating side, the border between SZ and TMAZ is of transition nature (Figure 5b). The microstructure of TMAZ in close proximity to SZ can be identified as bands of recovered grains irregularly separated by recrystallized grains. These differences between each side have their source in the unsymmetrical distribution of the temperature during the FSW process. Studies revealed that the retreating side is characterized by higher effect of the temperature [40,41]. During welding, the generated heat influences the plastically deformed microstructure by promoting heat-activated phenomena such as grain recovery and recrystallization. As can be seen, the dynamic recovery has a predominant role while fine, dynamically recrystallized grains are localized mainly between large, deformed grains in the form of long bands (Figure 5b). This suggests that the grains which undergo dynamic recrystallization are compressed and intensely elongated by the larger grains in the stirring process and the quantity of heat allows them to rebuild their severely deformed microstructures. This phenomenon has not been observed at the advancing side ( Figure  5a).

Microstructure Analysis and Microhardness Measurements
The microstructure of the TMAZ/SZ interface at the advancing side of the joint is characterized by the clear border between dynamically recrystallized grains of the stir zone and elongated grains of the thermo-mechanically affected zone, which visualize the flow of the material during the stirring process ( Figure 5a). Concurrently, at the retreating side, the border between SZ and TMAZ is of transition nature ( Figure 5b). The microstructure of TMAZ in close proximity to SZ can be identified as bands of recovered grains irregularly separated by recrystallized grains. These differences between each side have their source in the unsymmetrical distribution of the temperature during the FSW process. Studies revealed that the retreating side is characterized by higher effect of the temperature [40,41]. During welding, the generated heat influences the plastically deformed microstructure by promoting heat-activated phenomena such as grain recovery and recrystallization. As can be seen, the dynamic recovery has a predominant role while fine, dynamically recrystallized grains are localized mainly between large, deformed grains in the form of long bands (Figure 5b). This suggests that the grains which undergo dynamic recrystallization are compressed and intensely elongated by the larger grains in the stirring process and the quantity of heat allows them to rebuild their severely deformed microstructures. This phenomenon has not been observed at the advancing side (Figure 5a). The microhardness distribution of the join is presented in Figure 6. The obtained microhardness distribution allows observation of a typical "W"-shaped hardness curve. In the case of the advancing side, the lowest value of microhardness (108 HV0.1) was reported at the distance of 6 mm on the bottom part of the joint. Simultaneously, on the retreating side, the significant softening of welded material is reported. The low hardness zone is localized at the distance of 8 to 10 mm from the joint line, where the material was heated without being strain-hardened by plastic deformation. Generally, it can be stated that the reduction in microhardness predominantly affects the retreating side of the joint (about 25%), which confirms the higher effect of heat during welding in this area. Due to dynamic recrystallization in the stir zone leading to the formation of fine-grain microstructure, the microhardness slightly increases compared to TMAZ and HAZ. As can be seen, the highest value of microhardness in the stir zone is obtained for the upper part of this zone, where the effect of the tool shoulder severely refined the grains.

Tensile Test Results
The obtained tensile curves for AA2519-T62 base material and its friction stir welded joint are presented in Figure 7. The welding process has caused a significant reduction in elongation to failure of the material-from the value of 19% to 7%. The tensile strength of the tested joint was established as 389 MPa, which represents 83% joint efficiency. Comparing this value with the results of the studies [16,18], it is very high for high-strength aluminum alloy of 2XXX series, which can be partly The microhardness distribution of the join is presented in Figure 6. The obtained microhardness distribution allows observation of a typical "W"-shaped hardness curve. In the case of the advancing side, the lowest value of microhardness (108 HV0.1) was reported at the distance of 6 mm on the bottom part of the joint. Simultaneously, on the retreating side, the significant softening of welded material is reported. The low hardness zone is localized at the distance of 8 to 10 mm from the joint line, where the material was heated without being strain-hardened by plastic deformation. Generally, it can be stated that the reduction in microhardness predominantly affects the retreating side of the joint (about 25%), which confirms the higher effect of heat during welding in this area. Due to dynamic recrystallization in the stir zone leading to the formation of fine-grain microstructure, the microhardness slightly increases compared to TMAZ and HAZ. As can be seen, the highest value of microhardness in the stir zone is obtained for the upper part of this zone, where the effect of the tool shoulder severely refined the grains. The microhardness distribution of the join is presented in Figure 6. The obtained microhardness distribution allows observation of a typical "W"-shaped hardness curve. In the case of the advancing side, the lowest value of microhardness (108 HV0.1) was reported at the distance of 6 mm on the bottom part of the joint. Simultaneously, on the retreating side, the significant softening of welded material is reported. The low hardness zone is localized at the distance of 8 to 10 mm from the joint line, where the material was heated without being strain-hardened by plastic deformation. Generally, it can be stated that the reduction in microhardness predominantly affects the retreating side of the joint (about 25%), which confirms the higher effect of heat during welding in this area. Due to dynamic recrystallization in the stir zone leading to the formation of fine-grain microstructure, the microhardness slightly increases compared to TMAZ and HAZ. As can be seen, the highest value of microhardness in the stir zone is obtained for the upper part of this zone, where the effect of the tool shoulder severely refined the grains.

Tensile Test Results
The obtained tensile curves for AA2519-T62 base material and its friction stir welded joint are presented in Figure 7. The welding process has caused a significant reduction in elongation to failure of the material-from the value of 19% to 7%. The tensile strength of the tested joint was established as 389 MPa, which represents 83% joint efficiency. Comparing this value with the results of the studies [16,18], it is very high for high-strength aluminum alloy of 2XXX series, which can be partly

Tensile Test Results
The obtained tensile curves for AA2519-T62 base material and its friction stir welded joint are presented in Figure 7. The welding process has caused a significant reduction in elongation to failure of the material-from the value of 19% to 7%. The tensile strength of the tested joint was established as 389 MPa, which represents 83% joint efficiency. Comparing this value with the results of the studies [16,18], it is very high for high-strength aluminum alloy of 2XXX series, which can be partly explained by the high thickness of the workpiece (5 mm), which often results in the highest values of joint efficiency [2].  The digital image correlation technique provides more details about welded specimen behavior during the tensile test [42,43]. The obtained strain maps resulting from the DIC measurement for five points marked on the curve (Figure 7) are presented in Figure 8. The visualization of strain distribution allows observing of its uneven nature. The SZ and TMAZ are characterized by the highest strain, localized mainly in the central part of the SZ and advancing side of TMAZ. Of note is the fact that despite the retreating site the TMAZ has received a higher amount of heat during the welding process; its behavior in the tensile test reveals the spread of the The digital image correlation technique provides more details about welded specimen behavior during the tensile test [42,43]. The obtained strain maps resulting from the DIC measurement for five points marked on the curve (Figure 7) are presented in Figure 8.  The digital image correlation technique provides more details about welded specimen behavior during the tensile test [42,43]. The obtained strain maps resulting from the DIC measurement for five points marked on the curve (Figure 7) are presented in Figure 8. The visualization of strain distribution allows observing of its uneven nature. The SZ and TMAZ are characterized by the highest strain, localized mainly in the central part of the SZ and advancing side of TMAZ. Of note is the fact that despite the retreating site the TMAZ has received a higher amount of heat during the welding process; its behavior in the tensile test reveals the spread of the The visualization of strain distribution allows observing of its uneven nature. The SZ and TMAZ are characterized by the highest strain, localized mainly in the central part of the SZ and advancing side of TMAZ. Of note is the fact that despite the retreating site the TMAZ has received a higher amount of heat during the welding process; its behavior in the tensile test reveals the spread of the strain without significant concentration in the specific area. The failure in the tensile test occurred in the TMAZ/SZ interface on the advancing side of the weld. The fracture surface of the specimen was subjected to scanning electron microscope observations (Figure 9).
Materials 2020, 13, x FOR PEER REVIEW 8 of 18 strain without significant concentration in the specific area. The failure in the tensile test occurred in the TMAZ/SZ interface on the advancing side of the weld. The fracture surface of the specimen was subjected to scanning electron microscope observations (Figure 9). The observations have revealed that decohesion took place on the boundary between SZ and TMAZ. This area was identified in the microstructure analysis part (Figure 5a) as consisting of severely deformed, elongated grains of TMAZ adjoining the dynamically recrystallized grains of SZ. Such differences in the microstructure promote decohesion and it can explain the better coherency of the SZ/TMAZ interface on the retreating side, where the boundary between each zone is not clear and TMAZ has partly recrystallized microstructure.

Residual Stress Measurements
In order to compute the residual stresses in the FSW joint the following equations were used [44], Equations (1)-(3): where: σmax and σmin-the principle stresses; ε1, ε2, and ε3-the strains measured at the gauges of rosette no. 1, 2 and 3, respectively ( Figure 2b); A and B-the coefficients dependent on material properties, type of rosette, and dimensions of drilled hole; and α-the angle between gauge no. 1 and the direction of the nearest principle stress. The obtained results are presented in graphical form in Figure 10. Red arrows indicate the directions of tension stresses and blue arrows indicate the directions of compression stresses. The observations have revealed that decohesion took place on the boundary between SZ and TMAZ. This area was identified in the microstructure analysis part (Figure 5a) as consisting of severely deformed, elongated grains of TMAZ adjoining the dynamically recrystallized grains of SZ. Such differences in the microstructure promote decohesion and it can explain the better coherency of the SZ/TMAZ interface on the retreating side, where the boundary between each zone is not clear and TMAZ has partly recrystallized microstructure.

Residual Stress Measurements
In order to compute the residual stresses in the FSW joint the following equations were used [44], Equations (1)-(3): where: σ max and σ min -the principle stresses; ε 1 , ε 2 , and ε 3 -the strains measured at the gauges of rosette no. 1, 2 and 3, respectively ( Figure 2b); A and B-the coefficients dependent on material properties, type of rosette, and dimensions of drilled hole; and α-the angle between gauge no. 1 and the direction of the nearest principle stress. The obtained results are presented in graphical form in Figure 10. Red arrows indicate the directions of tension stresses and blue arrows indicate the directions of compression stresses. Results obtained by hole-drilling method were used to compute the values of longitudinal stresses σ11 and transverse stresses σ22. Nevertheless, in most cases, the values were obtained directly during the measurements because the directions of the main gauges (tensometers nos. 1 and 3) were convergent with the directions of searched stresses. The Equation (4) was used in other cases.
where α is an angle between the directions of searched stress σ11 (22) and known principle stress σmax(min). The results are presented in Figure 11. The results shown in Figure 11 indicate that tension stresses are present in the joint and its nearest vicinity. Greater values of stresses are in a longitudinal direction than transverse; additionally, they are greater on the advancing side than the retreating side. As the retreating side is characterized by higher heat input, it promotes stress relaxation and can explain the lower value of residual stress in this zone of the weld [40,41]. The highest stresses are at the weld line on the advancing side (159 MPa). The stresses in a joint center reach of 33-77 MPa and are similar in all directions. At the distance of approximately 10 mm from the weld joint, the stresses have very low values, near to zero. It should be noted that at these points the direction of principle stresses is the most rotated of all. It is probably caused by the material flow during the stirring process resulting in Results obtained by hole-drilling method were used to compute the values of longitudinal stresses σ 11 and transverse stresses σ 22 . Nevertheless, in most cases, the values were obtained directly during the measurements because the directions of the main gauges (tensometers nos. 1 and 3) were convergent with the directions of searched stresses. The Equation (4) was used in other cases.
where α is an angle between the directions of searched stress σ 11 (22) and known principle stress σ max(min) . The results are presented in Figure 11. Results obtained by hole-drilling method were used to compute the values of longitudinal stresses σ11 and transverse stresses σ22. Nevertheless, in most cases, the values were obtained directly during the measurements because the directions of the main gauges (tensometers nos. 1 and 3) were convergent with the directions of searched stresses. The Equation (4) was used in other cases.
where α is an angle between the directions of searched stress σ11 (22) and known principle stress σmax(min). The results are presented in Figure 11. The results shown in Figure 11 indicate that tension stresses are present in the joint and its nearest vicinity. Greater values of stresses are in a longitudinal direction than transverse; additionally, they are greater on the advancing side than the retreating side. As the retreating side is characterized by higher heat input, it promotes stress relaxation and can explain the lower value of residual stress in this zone of the weld [40,41]. The highest stresses are at the weld line on the advancing side (159 MPa). The stresses in a joint center reach of 33-77 MPa and are similar in all directions. At the distance of approximately 10 mm from the weld joint, the stresses have very low values, near to zero. It should be noted that at these points the direction of principle stresses is the most rotated of all. It is probably caused by the material flow during the stirring process resulting in The results shown in Figure 11 indicate that tension stresses are present in the joint and its nearest vicinity. Greater values of stresses are in a longitudinal direction than transverse; additionally, they are greater on the advancing side than the retreating side. As the retreating side is characterized by higher heat input, it promotes stress relaxation and can explain the lower value of residual stress in this zone of the weld [40,41]. The highest stresses are at the weld line on the advancing side (159 MPa). The stresses in a joint center reach of 33-77 MPa and are similar in all directions. At the distance of approximately 10 mm from the weld joint, the stresses have very low values, near to zero. It should be noted that at these points the direction of principle stresses is the most rotated of all. It is probably caused by the material flow during the stirring process resulting in the plastic deformation of less heated areas in the surrounding area. With increasing distance from the joint, the stresses change to compressive which are significantly greater in a longitudinal direction contrary to transverse. The obtained values of residual stresses and their nature of changes are similar to those available in other publications [45][46][47][48][49][50]. In all cases, the residual stresses are the highest at the weld line and lower at the center of joints. The values of stresses significantly decrease and change from tension stresses into compression. Moreover, the maximum value of measured residual stress equals approximately 50% of the yield stress of paternal material AA2519-T62. It should be noted that during the welding using either TIG or laser method the residual stresses can even reach the yield stress [2,3,51].

Low Cycle Fatigue Properties and Fracture Surface Observations
The variations of stress and plastic strain amplitudes with the number of cycles are shown in Figure 12a,b respectively. Regardless of the strain amplitude, three periods of fatigue life can be distinguished: cyclic hardening, cyclic stabilization, and final rapid drop in the value of stress amplitude until failure, which corresponds to the cyclic properties of the base material [52]. The length of each period differs depending on the used strain amplitude. Especially, for the highest values of strain amplitude (0.5, 0.6, and, 0.8%), it is difficult to identify the cyclic stabilization period and the samples undergo cyclic hardening until rapid failure. Compared to the base material, the obtained values of stabilized stress amplitude (Figure 12a) are lower for about 25% and also the reduction of fatigue life was reported [52]. At the same time, the registered plastic strain amplitudes of the AA2519-T62 FSW joint are much higher (Figure 12b). These differences can be partly explained by the lower ductility of the obtained joint compared to the base material (Figure 7).  [45][46][47][48][49][50]. In all cases, the residual stresses are the highest at the weld line and lower at the center of joints. The values of stresses significantly decrease and change from tension stresses into compression. Moreover, the maximum value of measured residual stress equals approximately 50% of the yield stress of paternal material AA2519-T62. It should be noted that during the welding using either TIG or laser method the residual stresses can even reach the yield stress [2,3,51].

Low Cycle Fatigue Properties and Fracture Surface Observations
The variations of stress and plastic strain amplitudes with the number of cycles are shown in Figure 12a,b respectively. Regardless of the strain amplitude, three periods of fatigue life can be distinguished: cyclic hardening, cyclic stabilization, and final rapid drop in the value of stress amplitude until failure, which corresponds to the cyclic properties of the base material [52]. The length of each period differs depending on the used strain amplitude. Especially, for the highest values of strain amplitude (0.5, 0.6, and, 0.8%), it is difficult to identify the cyclic stabilization period and the samples undergo cyclic hardening until rapid failure. Compared to the base material, the obtained values of stabilized stress amplitude (Figure 12a) are lower for about 25% and also the reduction of fatigue life was reported [52]. At the same time, the registered plastic strain amplitudes of the AA2519-T62 FSW joint are much higher (Figure 12b). These differences can be partly explained by the lower ductility of the obtained joint compared to the base material (Figure 7).  The hysteresis stress-strain loops of examined welded joint for various levels of strain amplitude are presented in Figure 13a-e and the stabilized loops are compared in Figure 13f.
Materials 2020, 13, x FOR PEER REVIEW 10 of 18 the plastic deformation of less heated areas in the surrounding area. With increasing distance from the joint, the stresses change to compressive which are significantly greater in a longitudinal direction contrary to transverse. The obtained values of residual stresses and their nature of changes are similar to those available in other publications [45][46][47][48][49][50]. In all cases, the residual stresses are the highest at the weld line and lower at the center of joints. The values of stresses significantly decrease and change from tension stresses into compression. Moreover, the maximum value of measured residual stress equals approximately 50% of the yield stress of paternal material AA2519-T62. It should be noted that during the welding using either TIG or laser method the residual stresses can even reach the yield stress [2,3,51].

Low Cycle Fatigue Properties and Fracture Surface Observations
The variations of stress and plastic strain amplitudes with the number of cycles are shown in Figure 12a,b respectively. Regardless of the strain amplitude, three periods of fatigue life can be distinguished: cyclic hardening, cyclic stabilization, and final rapid drop in the value of stress amplitude until failure, which corresponds to the cyclic properties of the base material [52]. The length of each period differs depending on the used strain amplitude. Especially, for the highest values of strain amplitude (0.5, 0.6, and, 0.8%), it is difficult to identify the cyclic stabilization period and the samples undergo cyclic hardening until rapid failure. Compared to the base material, the obtained values of stabilized stress amplitude (Figure 12a) are lower for about 25% and also the reduction of fatigue life was reported [52]. At the same time, the registered plastic strain amplitudes of the AA2519-T62 FSW joint are much higher (Figure 12b). These differences can be partly explained by the lower ductility of the obtained joint compared to the base material ( Figure 7).  The obtained loops reflect the fatigue properties observed in Figure 12a,b. In the first cycles of fatigue life, cyclic hardening occurs, which is much more visible in the samples tested with the strain amplitudes of 0.5, 0.6, and 0.8% (Figure 13c-e). The same strain amplitudes are characterized by a noticeable dissipation of strain energy manifested in a relatively high width of the mid-life hysteresis loops. The dissipation, caused predominantly by the plastic deformation leading to the development of microcracks [53], increases significantly together with the applied strain amplitude (Figure 13f). The parameters of the stabilized loops allowed establishing of the relationship of stress amplitude versus plastic strain amplitude, presented in Figure 14. The obtained curve can be described by the following equation [54]: The obtained loops reflect the fatigue properties observed in Figure 12a,b. In the first cycles of fatigue life, cyclic hardening occurs, which is much more visible in the samples tested with the strain amplitudes of 0.5, 0.6, and 0.8% (Figure 13c-e). The same strain amplitudes are characterized by a noticeable dissipation of strain energy manifested in a relatively high width of the mid-life hysteresis loops. The dissipation, caused predominantly by the plastic deformation leading to the development of microcracks [53], increases significantly together with the applied strain amplitude (Figure 13f). The parameters of the stabilized loops allowed establishing of the relationship of stress amplitude versus plastic strain amplitude, presented in Figure 14. The obtained loops reflect the fatigue properties observed in Figure 12a,b. In the first cycles of fatigue life, cyclic hardening occurs, which is much more visible in the samples tested with the strain amplitudes of 0.5, 0.6, and 0.8% (Figure 13c-e). The same strain amplitudes are characterized by a noticeable dissipation of strain energy manifested in a relatively high width of the mid-life hysteresis loops. The dissipation, caused predominantly by the plastic deformation leading to the development of microcracks [53], increases significantly together with the applied strain amplitude (Figure 13f). The parameters of the stabilized loops allowed establishing of the relationship of stress amplitude versus plastic strain amplitude, presented in Figure 14. The obtained curve can be described by the following equation [54]: The obtained curve can be described by the following equation [54]: where: σ a -stress amplitude (MPa), ε p -plastic strain amplitude [mm/mm], k -cyclic strength coefficient (MPa), and n -cyclic strain hardening exponent. The values of k' and n' have been taken directly from the function describing the plot in Figure 14: The obtained value of the cyclic strength coefficient (504.37 MPa) is almost three times lower than for the base material (1518.1 MPa) [52]. Although, the decrease of cyclic strength coefficient is expected for welded joints, in this case, the reported drop is significant and the obtained value is similar to the values reported by research [31,36,[55][56][57]. In the next step, the parameters of the stabilized loops were used for establishing the plots of elastic and plastic strain amplitudes vs. number of reversals, presented in Figure 15. where: σa-stress amplitude (MPa), εp-plastic strain amplitude [mm/mm], k′-cyclic strength coefficient (MPa), and n′-cyclic strain hardening exponent. The values of k' and n' have been taken directly from the function describing the plot in Figure 14: The obtained value of the cyclic strength coefficient (504.37 MPa) is almost three times lower than for the base material (1518.1 MPa) [52]. Although, the decrease of cyclic strength coefficient is expected for welded joints, in this case, the reported drop is significant and the obtained value is similar to the values reported by research [31,36,[55][56][57]. In the next step, the parameters of the stabilized loops were used for establishing the plots of elastic and plastic strain amplitudes vs. number of reversals, presented in Figure 15. The Manson-Coffin-Basquin equation is described by the following formula [54]: where: σ′f-fatigue strength coefficient (MPa), E-Young modulus (MPa); b-fatigue strength exponent; ε′f-fatigue ductility coefficient; and c-fatigue ductility exponent. The value of Young modulus of the welded joint was established in the tensile test and it is equal to 68 GPa. The specific values have been taken from the functions describing the plots in Figure 15: In terms of the plastic strain component, it is very close to the parameters set up for base material [52], while the elastic strain component deviates from it noticeably.
The tested samples tend to fail in the HAZ on the retreating side of the joint and for the lowest value of strain amplitude (0.35%) and in the case of the higher values of amplitude the failure occurred at the boundary of TMAZ/SZ on the advancing side of the weld. This zone was characterized in the microstructure analysis part (Figure 5a) and it is also a place of failure in the tensile test (Figure 9). For the fracture surface observations, the three representative samples were selected and tested at the amplitudes of 0.35%, 0.5%, and 0.8%. The fracture surface of the first sample is presented in Figure 16a-c. The Manson-Coffin-Basquin equation is described by the following formula [54]: where: σ f -fatigue strength coefficient (MPa), E-Young modulus (MPa); b-fatigue strength exponent; ε f -fatigue ductility coefficient; and c-fatigue ductility exponent. The value of Young modulus of the welded joint was established in the tensile test and it is equal to 68 GPa. The specific values have been taken from the functions describing the plots in Figure 15: In terms of the plastic strain component, it is very close to the parameters set up for base material [52], while the elastic strain component deviates from it noticeably.
The tested samples tend to fail in the HAZ on the retreating side of the joint and for the lowest value of strain amplitude (0.35%) and in the case of the higher values of amplitude the failure occurred at the boundary of TMAZ/SZ on the advancing side of the weld. This zone was characterized in the microstructure analysis part (Figure 5a) and it is also a place of failure in the tensile test (Figure 9).
For the fracture surface observations, the three representative samples were selected and tested at the amplitudes of 0.35%, 0.5%, and 0.8%. The fracture surface of the first sample is presented in Figure 16a-c. The fracture surface has a typical structure of fatigue fracture, consisting of an initiation site, fatigue crack propagation region, and rapid fracture area (Figure 16a). The initiation of fatigue crack in localized in the top corner of the investigated sample, the surface is free of visible defects, smooth, and contains river-like patterns (Figure 16b). The fatigue crack propagation area is characterized by the presence of fatigue striations with noticeable participation of secondary cracks (marked with yellow arrows) indicating a local intensification of material decohesion (Figure 16c). The observed fractured surface is of mixed ductile and brittle character. The fracture surface of the sample tested with the strain amplitude of 0.5% is presented in Figure 17a-c. The fracture surface has a typical structure of fatigue fracture, consisting of an initiation site, fatigue crack propagation region, and rapid fracture area (Figure 16a). The initiation of fatigue crack in localized in the top corner of the investigated sample, the surface is free of visible defects, smooth, and contains river-like patterns (Figure 16b). The fatigue crack propagation area is characterized by the presence of fatigue striations with noticeable participation of secondary cracks (marked with yellow arrows) indicating a local intensification of material decohesion (Figure 16c). The observed fractured surface is of mixed ductile and brittle character. The fracture surface of the sample tested with the strain amplitude of 0.5% is presented in Figure 17a-c. The fracture surface has a typical structure of fatigue fracture, consisting of an initiation site, fatigue crack propagation region, and rapid fracture area (Figure 16a). The initiation of fatigue crack in localized in the top corner of the investigated sample, the surface is free of visible defects, smooth, and contains river-like patterns (Figure 16b). The fatigue crack propagation area is characterized by the presence of fatigue striations with noticeable participation of secondary cracks (marked with yellow arrows) indicating a local intensification of material decohesion (Figure 16c). The observed fractured surface is of mixed ductile and brittle character. The fracture surface of the sample tested with the strain amplitude of 0.5% is presented in Figure 17a-c. The surface is more similar to the tensile sample and no areas typical for fatigue fracture have been distinguished under macroscopic observation. The observed surface contains part of the SZ/TMAZ interface with a visible texture demonstrating plastic flow in the stirring process ( Figure  17a). At this interface, it is possible to identify vertical "cliffs" suggesting the grain-boundary separation ( Figure 17b). However, the higher magnification revealed the fatigue striations and occurrence of numerous, very deep secondary cracks (Figure 17c). Not only the specific microstructure (Figure 5a) but also the high values of residual stresses ( Figure 11) can promote decohesion in this area. The near-surface stresses directly influence the initiation of a fatigue crack. This phenomenon is of great importance in the cases of relatively low plastic strain amplitudes, and for the large plastic deformation, the residual stresses tend to relax.
No noticeable presence of secondary cracks between the fatigue striations was reported. The fracture surface of the sample tested with the strain amplitude of 0.8% is presented in Figure 18a-c.  The surface is more similar to the tensile sample and no areas typical for fatigue fracture have been distinguished under macroscopic observation. The observed surface contains part of the SZ/TMAZ interface with a visible texture demonstrating plastic flow in the stirring process (Figure 17a). At this interface, it is possible to identify vertical "cliffs" suggesting the grain-boundary separation ( Figure 17b). However, the higher magnification revealed the fatigue striations and occurrence of numerous, very deep secondary cracks (Figure 17c). Not only the specific microstructure (Figure 5a) but also the high values of residual stresses ( Figure 11) can promote decohesion in this area. The near-surface stresses directly influence the initiation of a fatigue crack. This phenomenon is of great importance in the cases of relatively low plastic strain amplitudes, and for the large plastic deformation, the residual stresses tend to relax.
No noticeable presence of secondary cracks between the fatigue striations was reported. The fracture surface of the sample tested with the strain amplitude of 0.8% is presented in Figure 18a-c. The surface is more similar to the tensile sample and no areas typical for fatigue fracture have been distinguished under macroscopic observation. The observed surface contains part of the SZ/TMAZ interface with a visible texture demonstrating plastic flow in the stirring process ( Figure  17a). At this interface, it is possible to identify vertical "cliffs" suggesting the grain-boundary separation ( Figure 17b). However, the higher magnification revealed the fatigue striations and occurrence of numerous, very deep secondary cracks (Figure 17c). Not only the specific microstructure (Figure 5a) but also the high values of residual stresses ( Figure 11) can promote decohesion in this area. The near-surface stresses directly influence the initiation of a fatigue crack. This phenomenon is of great importance in the cases of relatively low plastic strain amplitudes, and for the large plastic deformation, the residual stresses tend to relax.
No noticeable presence of secondary cracks between the fatigue striations was reported. The fracture surface of the sample tested with the strain amplitude of 0.8% is presented in Figure 18a-c.