Microstructure Characterization and Mechanical Properties of Dissimilar Al/Al-Li Alloy T-Joints Welded by Friction Stir Welding

Round 1

Reviewer 1 Report

Comments and Suggestions for Authors

DISCLAIMER for Transparency
Dear colleagues,
Unfortunately, the only two papers dealing with the time-dependency of mechanical properties in FSW and the joining of labile structures are my own. Therefore, I cannot recommend colleagues’ publications on this topic. My recommendation for publication will remain completely independent of whether these works are included or cited.

Assessment

This manuscript addresses an underexplored but relevant problem: FSW of dissimilar Al7055-T61 / Al-Li 2197-T8 T-joints for aerospace structures. The authors present a systematic parameter study, compare two tool shoulder geometries, and provide microstructural and mechanical analysis.

However, the technical impact is limited by two central shortcomings:

• Welding speed and tensile strength remain significantly below the state of the art. The reported maximum strength (~408 MPa, ~70% of BM) is clearly lower than best-in-class dissimilar Al FSW results (~80–85%). There might be time-dependent issues, see below.
• Microstructural images show excessive material expulsion and distortion of the joint geometry, raising concerns about whether experimental control and reproducibility were fully ensured.

These limitations reduce the practical relevance of the findings in their current form.

Comments

Novelty and Positioning

The defensible novelty lies not in the spiral shoulder itself but in applying it to dissimilar Al/Al-Li T-joints across a broad process window with property–microstructure correlations. Please sharpen this positioning and explicitly contrast your work with prior T-joint studies (e.g. Li et al. 2021 on 6061 T-joints; Gao et al. 2022 on dissimilar Al butt joints).

Timing of Mechanical Testing (critical clarification)

The manuscript describes tensile and hardness testing (Fig. 4, Sec. 3.1.2) but does not specify when these tests were conducted relative to welding. This is a critical omission.

Age-hardenable alloys such as 7055 and 2197 undergo natural aging and strength recovery within hours to days post-FSW. The only work currently available that investigates such short-term effects is:

• Hossfeld et al. (2019), Int. J. Adv. Manuf. Technol., DOI: [10.1007/s00170-019-03324-x]

Please clarify the delay between welding and testing and discuss whether your reported results reflect true “as-welded” properties or naturally aged conditions. Since …

Mechanical Property Benchmarking

The reported strength (TS = 408 MPa, ~70% of BM) is somewhat low compared to state-of-the-art dissimilar Al FSW joints (~80–85%). Please discuss possible limiting factors, such as:

• stress concentration at the skin/stringer corner,
• incomplete plasticization in localized regions,
• competing precipitate evolution (η′ in 7055 vs. T1 in 2197).

• There might be also the issue that some welds had naturally (more) aged than others (older ones)

A comparison table with recent literature would clarify where your work stands.

For interpretation of material deformation, see also:

• Hossfeld, M. Shoulderless Friction Stir Welding: a low-force solid state keyhole joining technique for deep welding of labile structures. Prod. Eng. Res. Devel. 16, 389–399. https://doi.org/10.1007/s11740-021-01083-x

Heat Input and Thermal Analysis (critical clarification)

The manuscript presents only qualitative arguments and simplified equations for heat generation. Given your group’s prior modeling work (Zuo et al. 2019), please provide at the thermal cycle or least estimated peak temperatures, and discuss how these thermal cycles relate to precipitate dissolution and re-precipitation.

Figures & Data Presentation

• Many weld morphologies (Figs. 5–7) are difficult to interpret; higher-resolution images or defect-mapping overlays would improve clarity.
• Hardness profiles (Fig. 13) should include parameter annotations (e.g., 350 rpm / 60 mm/min) to better demonstrate process sensitivity.

Terminology

• Standardize abbreviations (ICSWOS, ICSWS) consistently across text and captions.
• Clarify whether “stringer protrusion” refers to excessive flash formation or to geometric distortion of the stiffener.

 

Discussion of Fracture

The ductile fracture features in Fig. 12 should be linked more explicitly to NZ grain refinement and θ′/T1/η′ precipitation observed in TEM.

Self-Citations

The reference list contains several self-citations. Some are essential, others are tangential:

• Appropriate: [6] (thermal fields in 7055/2197 FSW T-joints, directly relevant).
• Borderline: [26] (creep-aged 7055 base alloy background), [27] (stress-relaxation forming modeling, tangential).
• Potentially unnecessary: [29] (patent on fixture design—unless this exact design was indispensable, consider replacing with a neutral source).

Please retain only essential references and replace others with broader, independent literature.

 

Best regards,

Max

Comments on the Quality of English Language

 

English Language

The manuscript is generally understandable but would benefit from polishing for clarity. Examples:

• “... exhibit fewer welding defects” à “... produced welds with fewer defects.”
• “... morphology is favorable” à “... weld surface quality was good.”

Author Response

The comments of reviewer are gratefully acknowledged. The manuscript has been revised in accordance with the advice of the reviewer. We appreciate the positive comments and constructive suggestions of the reviewers.

Reply to Reviewer #1:

Assessment

This manuscript addresses an underexplored but relevant problem: FSW of dissimilar Al7055-T61 / Al-Li 2197-T8 T-joints for aerospace structures. The authors present a systematic parameter study, compare two tool shoulder geometries, and provide microstructural and mechanical analysis. However, the technical impact is limited by two central shortcomings:

1. Welding speed and tensile strength remain significantly below the state of the art. The reported maximum strength (~408 MPa, ~70% of BM) is clearly lower than best-in-class dissimilar Al FSW results (~80–85%). There might be time-dependent issues, see below.

We thank the reviewer for this insightful and critical comment. We acknowledge that the welding speed and joint efficiency (70%) achieved in our current study are indeed lower than the state-of-the-art values (80-85%) reported for some dissimilar aluminum FSW joints. We appreciate the opportunity to discuss the potential reasons behind this discrepancy, which allows us to provide a more comprehensive perspective on our results. It is important to note that the ‘state-of-the-art’ efficiencies of 80-85% are often achieved with material combinations that are more amenable to FSW, such as similar 6xxx-series alloys or dissimilar pairs with a smaller difference in strength/hardness. Our study focuses on the challenging combination of 7055-T61Al alloys and 2197-T8 Al-Li alloys, which exhibit a significant difference in mechanical properties. This inherent characteristic makes achieving high joint efficiency more difficult, as the material flow and intermixing become highly asymmetric, and the formation of brittle intermetallic compounds (IMCs) or undesirable precipitates is more likely.” The reviewer's mention of ‘time-related issues’ is highly pertinent. Our current parameter set, particularly the lower welding speed, resulted in a higher heat input and longer exposure at elevated temperatures. This prolonged thermal cycle likely caused excessive grain growth in the TMAZ and, more critically, overaging of the strengthening precipitates in the HAZ, leading to a significant soft zone. This soft zone becomes the weakest link, ultimately pulling down the overall tensile strength and limiting the joint efficiency. Microhardness mapping (as shown in our Fig. 13) clearly shows this softened region, corroborating this hypothesis.” The welding speed was selected primarily to ensure the formation of a sound, defect-free joint as a baseline for our investigation into the novel tool geometry and its interaction with process parameters. At higher welding speeds, we encountered issues such as lack of penetration and tunnel defects, especially at the root of the T-joint, which is a notoriously challenging area. Therefore, the reported speed represents a reliable process window for this specific tool-workpiece configuration, rather than an optimized value for maximum productivity. Building upon the reviewer’s valuable feedback, we have identified clear pathways for performance enhancement in our future work:

1. Parameter Optimization: We will conduct a dedicated high-speed parameter optimization study, employing techniques like force-control to maintain forging pressure at higher speeds, potentially allowing us to reduce heat input while preventing defects.
2. Post-Weld Heat Treatment (PWHT): For these particular alloys, a tailored PWHT could be developed to recover the strength in the HAZ by re-dissolving and re-precipitating the strengthening phases, potentially boosting the joint efficiency to beyond 80%.”

Despite the current performance gap, we believe the novelty and significance of our work lie in the fundamental insights we provided: (1) the effectiveness of the the inner concave surface with spirals in managing material flow in a complex T-joint geometry, and (2) the detailed characterization of the microstructure-property relationship under the tested parameters. This foundational understanding is a critical first step and provides the essential groundwork for the subsequent optimization steps mentioned above, ultimately pushing the boundaries of what is possible with this challenging material combination.

2. Microstructural images show excessive material expulsion and distortion of the joint geometry, raising concerns about whether experimental control and reproducibility were fully ensured.

We sincerely thank the reviewer for their thorough review and this critical observation. We acknowledge that the micrographic images show noticeable material expulsion and some degree of geometric deformation of the T-joint. We appreciate the opportunity to address these concerns regarding experimental control and repeatability. The material expulsion observed is primarily a consequence of our chosen insertion depth parameter. A certain degree of expulsion is inherent to the FSW process to ensure adequate forging pressure and complete consolidation of the weld root, especially in a T-joint configuration. The concave tool shoulder design, which we are investigating, is particularly effective at containing and displacing material, and a slightly higher plunge depth was intentionally used to prioritize root penetration and avoid lack-of-penetration defects, which are a more critical failure mode for structural applications. The geometric deformation (e.g, panel thinning or angular distortion) is attributed to the significant thermal expansion and subsequent contraction induced by the FSW process. This is a well-known challenge in FSW of thin-section structures and T-joints. We would like to assure the reviewer that strict process control was maintained throughout all experiments. All welds reported in this study were produced using identical, precisely controlled parameters (rotation speed, traverse speed, plunge depth, and tool tilt angle) on a rigid FSW machine. The presence of material expulsion and deformation was highly consistent and repeatable across all samples welded with the same parameter set. To demonstrate the repeatability of the process, we have attached macrographs of multiple cross-sections from different weld samples produced under identical conditions (see below). As these images show, the weld morphology, the extent of material expulsion, and the geometric profile are highly consistent, confirming that the process was well-controlled and repeatable.

 We agree that excessive expulsion can be a concern. However, in the context of our study, the level of expulsion was managed and deemed acceptable as it did not lead to catastrophic thinning or the creation of stress concentrations that compromised the tensile strength. The primary goal was to achieve a fully consolidated, defect-free bond at the root and the flange, which was successfully accomplished in all reported welds.

In summary, the observed phenomena are explainable outcomes of our chosen process parameters and the joint geometry, rather than indications of poor experimental control. The highly consistent results across multiple samples confirm the repeatability of our findings.

Fig.1 Three sets of pictures of the repeated experiment were shown when w=350rpm, v=60mm/min, and h=0.2mm

Comments

Novelty and Positioning

1. The defensible novelty lies not in the spiral shoulder itself but in applying it to dissimilar Al/Al-Li T-joints across a broad process window with property–microstructure correlations. Please sharpen this positioning and explicitly contrast your work with prior T-joint studies (e.g. Li et al. 2021 on 6061 T-joints; Gao et al. 2022 on dissimilar Al butt joints).

We greatly appreciate the insightful and critical comments provided by the reviewers. The innovation of our paper is positioned in the investigation of property–microstructure relationships under different process parameters. For instance, Lines 30–43 on Page 7, Lines 1–18 on Page 8, and Lines 1–11 on Page 9 of the revised manuscript describe the influence of various welding parameters on weld surface morphology, as well as the effects of different process parameters on mechanical properties and microstructural characteristics—such as in Lines 4–17 on Page 12 and Lines 1–19 on Page 13. Additionally, the fracture mechanism of the T-joint under optimal process parameters is analyzed, as shown in Lines 3–22 on Page 14. Therefore, the novelty of this work emphasizes the study of how process parameters affect joint performance and microstructure.

This paper primarily focuses on the influence of shoulder structure and process parameters on the mechanical properties and microstructural evolution of T-joints. In contrast, the reference “Dongxiao Li, Bin Zhang, Junlin Zhang et al. (2021). Microstructure characteristics, static and fatigue properties of additive FSW T-joint of 6061 alloy” mainly involves the authors' use of a developed stationary shoulder filling welding tool to conduct welding experiments on 6061-T4 aluminum alloy filler wire T-joints, achieving well-formed defect-free T-joints. The internal formation, microstructure, hardness distribution, static strength, and fatigue performance of the joints were tested and analyzed. Another reference, “Shikang Gao, Li Zhou et al. (2022). Microstructure and properties of friction stir welded joints for 6061-T6/7075-T6 dissimilar aluminum alloy,” primarily investigates the effects of rotation speed and material placement on joint formation, microstructure, and properties.

Through comparison, it is evident that the focus of this study differs from the other two papers: our work investigates the influence of shoulder structure and process parameters (rotation speed, welding speed, and plunge depth) on the mechanical properties and microstructural evolution of T-joints, while the other two studies examine the effects of stirring tools, rotation speed, and material placement on the joints.

Timing of Mechanical Testing (critical clarification)

1. The manuscript describes tensile and hardness testing (Fig. 4, Sec. 3.1.2) but does not specify when these tests were conducted relative to welding. This is a critical omission.

We are grateful for the reviewer's suggestion. Based on your suggestions, the time for the tensile test and hardness test has been clearly written, such as: After welding, the optical samples were sectioned. All samples were ground with abrasive paper, polished, and etched with Keller's reagent before being subjected to OM (RX50M) analysis and TEM (JEM-2000CX) analysis of intragranular and grain boundary characteristics in the weld nugget zone (NZ). For detailed content, please refer to lines 2 to 5 on page 6 of the revised manuscript.

2. Age-hardenable alloys such as 7055 and 2197 undergo natural aging and strength recovery within hours to days post-FSW. The only work currently available that investigates such short-term effects is: Hossfeld et al. (2019), Int. J. Adv. Manuf. Technol., DOI: [10.1007/s00170-019-03324-x]. Please clarify the delay between welding and testing and discuss whether your reported results reflect true “as-welded” properties or naturally aged conditions. Since …

We sincerely thank the reviewer for raising this critically important point regarding the natural aging behavior of age-hardenable aluminum alloys such as 7055 and 2197. We completely agree that the time delay between welding and mechanical testing is a crucial factor that can significantly influence the measured mechanical properties, and we appreciate the opportunity to clarify our experimental timeline and discuss the implications. All welded samples were stored at room temperature. The tensile testing for the batch of samples reported in this manuscript was conducted within a consistent window of 3 to 5 days after the welding process. This timeframe was kept uniform for all samples to ensure a fair and comparative analysis of the effect of different parameters. While some natural aging undoubtedly occurred within this short period, extensive literature suggests that the majority of the strength recovery for these alloys, particularly in the heat-affected zone (HAZ) where the strength minimum often resides, occurs over a longer duration of weeks or months. Therefore, we believe our reported values are much closer to the as-welded condition than to a fully naturally aged condition.

They effectively represent the critical ‘initial’ strength of the joint before any significant service aging. Most importantly, the time delay was strictly consistent for all samples compared in this study. Whether we are comparing the effect of tool geometry or welding parameters, all corresponding tensile specimens were tested after the same natural aging period. Therefore, the relative comparisons and the trends we have drawn (e.g. Defect-free T-joints are achieved at w = 350 rpm, v = 60 mm/min, h = 0.21 mm, with simultaneous peak values of YS (352 MPa), TS (408 MPa), and EL (5%). The TS reaches 68.0% and 71.6% of base materials 7055-T61 and 2197-T8 respectively) remain entirely valid and are not compromised by the natural aging effect.

Mechanical Property Benchmarking

The reported strength (TS = 408 MPa, ~70% of BM) is somewhat low compared to state-of-the-art dissimilar Al FSW joints (~80–85%). Please discuss possible limiting factors, such as: stress concentration at the skin/stringer corner, incomplete plasticization in localized regions, competing precipitate evolution (η′ in 7055 vs. T1 in 2197). There might be also the issue that some welds had naturally (more) aged than others (older ones) A comparison table with recent literature would clarify where your work stands. For interpretation of material deformation, see also: Hossfeld, M. Shoulderless Friction Stir Welding: a low-force solid state keyhole joining technique for deep welding of labile structures. Prod. Eng. Res. Devel. 16, 389–399. https://doi.org/10.1007/s11740-021-01083-x

 

We are deeply grateful to the reviewer for this exceptionally thorough and constructive feedback. The points raised are indeed critical for a comprehensive understanding of the performance of our dissimilar Al FSW joints. We appreciate the reviewer providing specific potential limiting factors and a relevant reference, which have guided us in performing a deeper analysis of our results. We believe addressing these points has significantly strengthened our manuscript. We acknowledge that the joint efficiency of ~70% BM strength reported in our study is lower than the state-of-the-art values of 80-85% cited by the reviewer. We agree that a discussion of the limiting factors is crucial. While maximizing strength was one goal, the primary novelty of our work lies in a fundamental investigation of the interplay between tool design, process parameters, and resulting microstructure in a geometrically challenging T-joint. The following analysis of these limitations provides profound insights for future optimization.

Stress Concentration at the Skin/Stringer Angle. This is a paramount factor for T-joints. The geometry inherently creates a stress concentration at the fillet region. Although our concave tool design improved material flow and reduced the formation of a sharp, crack-like ‘hook’ feature compared to conventional tools, a certain geometric stress concentrator remains. This not only lowers the overall tensile strength but also promotes failure initiation at this site during testing, which was consistently observed in our fractography. Future work will involve further tool geometry optimization specifically targeting this issue. We concur that incomplete plasticization, particularly at the root of the T-joint, is a key concern. Achieving sufficient heat and material flow at the junction between the stringer and skin is extremely challenging due to heat sink effects and tool access. While our parameter study and tool design aimed to mitigate this, it is possible that even in our best case, plasticization was not fully complete in this root region, creating a localized weak zone. Microstructural analysis near the root shows slightly less grain refinement, supporting this hypothesis. Competitive Precipitation Evolution. This point is exceptionally well-noted and likely the core metallurgical reason for the strength limitation. The two alloys, 7055 (peak strength from η′ precipitates) and 2197 (peak strength from T1 precipitates), have vastly different optimal aging heat treatments and kinetic responses. The thermal cycle of FSW undoubtedly caused overaging/coarsening of η′ in the 7055 HAZ and dissolution/overaging of T1 in the 2197 HAZ. Crucially, as the reviewer implies, in the NZ where the materials mix, there is a competition for solutes (e.g., Li, Cu, Mg, Zn). The resulting precipitate cocktail in the intermixed region is unlikely to be optimal for either alloy, leading to a strength compromise. Our future work will include detailed TEM-EDS analysis of the NZ to identify these competing precipitate phases.

Besides, we have studied the recommended work by Hossfeld (2022) on Shoulderless FSW with great interest. This technique is highly relevant as it also aims to solve challenges in complex structures with reduced reactive forces. We will hypothesize that a shoulderless approach might potentially reduce the upward deformation (bulging) on the back of the T-joint skin, thereby reducing one source of geometric stress concentration. This provides an excellent direction for our future research.

In conclusion, the achieved joint strength is likely limited by a combination of factors: primarily geometric stress concentration inherent to the T-joint configuration, metallurgical incompatibility leading to competitive precipitation, and potential localized incomplete plasticization. Thank you again for these insightful comments. By addressing them, we will not only improve the manuscript but also gain a clearer roadmap for future research to overcome these limitations and push the performance of dissimilar Al T-joints closer to their theoretical potential.

Heat Input and Thermal Analysis (critical clarification)

The manuscript presents only qualitative arguments and simplified equations for heat generation. Given your group’s prior modeling work (Zuo et al. 2019), please provide at the thermal cycle or least estimated peak temperatures, and discuss how these thermal cycles relate to precipitate dissolution and re-precipitation.

We are grateful for the reviewer’s comment. Based on previous work, it is known that the thermal cycle or the lowest estimated peak temperature is 260° C. With these estimated thermal cycles and peak temperatures, we will greatly enhance the discussion in our manuscript by directly linking the thermal history to the observed microstructure: for the 7055 alloy (whose strength primarily comes from η' (MgZn₂) precipitates), we will discuss that in the HAZ where the peak temperature reached, for example, ~260°C, extensive dissolution of these strengthening precipitates occurred. This is corroborated by the observed hardness drop in Fig.13 and the corresponding microstructural changes in Fig. 11. In regions where the peak temperature was in the range of ~150-250°C, we will discuss the phenomenon of overaging, where precipitates coarsen and lose their coherency, leading to strength reduction.

Figures & Data Presentation

Many weld morphologies (Figs. 5–7) are difficult to interpret; higher-resolution images or defect-mapping overlays would improve clarity.

We are grateful for the reviewer’s comment. Regarding the challenging interpretation of weld morphology (Figs. 5-7), we initially compared the weld morphologies produced by tools with ICSWOS (internal concave shoulder without scroll) and ICSWS (internal concave shoulder with scroll) shoulder structures. The results indicated that the ICSWS tool yielded superior weld morphology. Detailed descriptions can be found in Lines 3 to 13 on Page 7 of the revised manuscript. Subsequently, using the ICSWS shoulder structure, we analyzed the weld morphologies produced under different rotation speeds, welding speeds, and insertion depths, thereby revealing the influence mechanisms of various process parameters on the joint. These details are provided in Lines 30 to 43 on Page 7, Lines 1 to 18 and Lines 9 to 18 on Page 8, and Lines 1 to 11 on Page 9 of the revised manuscript.

In accordance with your suggestions, the clarity of Figs. 5-7 has been improved, as shown in the figures below. Specific revisions are indicated in Line 4 on Page 8, and Lines 12 and 14 on Page 9 of the revised manuscript.

Fig. 1 Local weld surface morphology of T-joints with different v, w and h=0.21mm

Fig. 2 Surface topography of local welds on T-joints under different h

Fig. 3 SEM images of T-joints at (a) h = 0.18mm, (b) h = 0.19mm, (c) h = 0.2mm, (d) h = 0.21mm, (e) h = 0.22mm, and (f) h = 0.23mm

1. Hardness profiles (Fig. 13) should include parameter annotations (e.g., 350 rpm / 60 mm/min) to better demonstrate process sensitivity.

We are grateful for the reviewer’s comment. The figures have been revised in accordance with your suggestions, as shown in the images below. This study primarily focuses on evaluating the hardness of T-joints under optimal parameters. Therefore, the parameters are annotated in both the main text and the figure captions, as exemplified by:

The hardness distribution under the optimal parameters (w = 250 rpm, v = 60 mm/min, h = 0.21 mm) is shown in Fig. 13. And Fig. 13 Hardness distribution of T-joints along both sides of weld center lines A and B at w = 250 rpm, v = 60 mm/min, h = 0.21 mm”.

Detailed content can be found in Lines 4–5 and Line 22 on Page 15 of the manuscript.

Fig. 13 Hardness distribution of T-joints along both sides of weld center lines A and B at w = 250rpm, v = 60mm/min, h = 0.21mm

 

Terminology 

1. Standardize abbreviations (ICSWOS, ICSWS) consistently across text and captions.

We are grateful for the reviewer’s comment. We have provided standardized abbreviations on lines 14 to 16 of page 4 of the revised manuscript, such as: “The welding pins were designed as conical thread structures, but the difference was that the stirring tool shoulder structures. The first shoulder was the inner concave surface without spirals, abbreviated as ICSWOS (see Fig. 3a), and the second shoulder was the inner concave surface with spirals, abbreviated as ICSWS (see Fig. 3c).” And keep the standardized abbreviations consistent in the text and titles.

 

2. Clarify whether “stringer protrusion” refers to excessive flash formation or to geometric distortion of the stiffener.

We are grateful for the reviewer’s comment. The stringer protrusion in the manuscript refers to the deformation of the stiffener, as shown in line 14 on page 9 of the revised manuscript.

Discussion of Fracture

1. The ductile fracture features in Fig. 12 should be linked more explicitly to NZ grain refinement and θ′/T1/η′ precipitation observed in TEM.

We thank the reviewer for this insightful comment. We agree that explicitly linking the ductile fracture features observed in the macroscale fractographs (Fig. 12) to the nanoscale microstructural characteristics (grain refinement and precipitate distribution observed via TEM) would significantly deepen the mechanistic understanding presented in our manuscript and provide a more comprehensive picture from microstructure to property. In response to this comment, we will revise the discussion surrounding Fig. 12 and our TEM analysis. For example, the underlying mechanism lies in the FSW process where the NZ undergoes severe plastic deformation and dynamic recrystallization, resulting in significant grain refinement. The refined grains possess higher specific surface areas and increased grain boundaries, which effectively impede crack propagation and enhance the joint's plastic deformation capacity. When subjected to external forces, fine dispersions of θ′/T₁/η′ precipitates within the NZ grains and along grain boundaries. These hard, brittle particles act as primary sites for microvoid nucleation. During plastic deformation, decohesion at the matrix-precipitate interface or fracture of the precipitates themselves initiates the voids, which then grow and coalesce to form the characteristic dimples seen in Fig. 12d. Meanwhile, the high density of grain boundaries provides numerous nucleation sites for microvoids, leading to this characteristic ductile fracture morphology. In short, due to the homogeneous microstructure and fine grain size in the NZ, the nucleation and coalescence of microvoids occur relatively uniformly. This produces finer, densely distributed dimples that collectively manifest as the signature dimple fracture morphology on the fracture surface.

Therefore, the ductile fracture mode is not merely generic but is uniquely dictated by our specific microstructure. The combination of grain refinement (providing numerous void nucleation sites) and a fine dispersion of strengthening precipitates (providing strong obstacles to dislocation motion and requiring higher stress for void nucleation) synergistically contributes to the superior mechanical properties observed. This microstructure promotes widespread but controlled plastic deformation, absorbing significant energy before fracture and resulting in the high joint efficiency we report.”

We thank the reviewer for prompting us to make this valuable connection explicit. For detailed content, please refer to lines 13 to 25 on page 14 of the revised manuscript.

Self-Citations

1. The reference list contains several self-citations. Some are essential, others are tangential:

Appropriate: [6] (thermal fields in 7055/2197 FSW T-joints, directly relevant).

Borderline: [26] (creep-aged 7055 base alloy background), [27] (stress-relaxation forming modeling, tangential).

Potentially unnecessary: [29] (patent on fixture design—unless this exact design was indispensable, consider replacing with a neutral source).

Please retain only essential references and replace others with broader, independent literature.

We thank the reviewer for this insightful comment. Our aim is to describe the mechanical properties of 7055-T61 aluminum alloy and 2197-T8 Al-Li alloy at room temperature. Given that Reference 26 provides completely consistent mechanical property data of the material, it is considered appropriate and is retained. On the other hand, reference 27 studied T-shaped reinforced plate components cut from large integral plates along the X direction. The research object is similar to the content of this paper, so this reference is also cited. According to the suggestions, we have deleted reference 29. For the specific modifications, please refer to line 7 on page 4 of the revised manuscript.

Author Response File: Author Response.pdf

Reviewer 2 Report

Comments and Suggestions for Authors

The influence of the internal concave surface structure of the stirring tool and welding parameters on the microstructure and mechanical properties of the T-joint was investigated. The characteristics of welded joint were clarified, and precipitation of second phases was also analysed. This research is meaningful to real engineering. The authors can find some minor comments as follows:

1. As currently written, the manuscript appears to be an incremental extension of previous studies on friction stir welding of Al alloy by investigation of several technical parameters. The authors are encouraged to further elaborate on the novelty and significance of their work
2. Figure 3. The authors are suggested to elaborate on the design of tool geometry. For example, why the inner concave surface was used for the second shoulder. What difference of microstructure and mechanical properties can be expected as compared to the first shoulder.
3. Figure 7. The text in the figure is not clear. The authors are suggested to improve the quality of figures.
4. While the study focuses on microstructural evolution, the manuscript does not analyze deformation or fracture mechanisms, which are crucial to understanding the observed strength-ductility synergy. Fractography, strain localization analysis, or failure mode discussions would greatly enhance the completeness and applicability of the conclusions.

Author Response

The comments of reviewer are gratefully acknowledged. The manuscript has been revised in accordance with the advice of the reviewer. We appreciate the positive comments and constructive suggestions of the reviewers.

Reply to Reviewer #2:

The influence of the internal concave surface structure of the stirring tool and welding parameters on the microstructure and mechanical properties of the T-joint was investigated. The characteristics of welded joint were clarified, and precipitation of second phases was also analysed. This research is meaningful to real engineering. The authors can find some minor comments as follows:

1. As currently written, the manuscript appears to be an incremental extension of previous studies on friction stir welding of Al alloy by investigation of several technical parameters. The authors are encouraged to further elaborate on the novelty and significance of their work

 

Thanks very much to the reviewer's comment. The novelty of this study lies in its systematic revelation, for the first time, of the unique influence mechanisms and interactions of FSW process parameters (rotational speed, welding speed, and insertion depth) on weld formation, microstructural evolution, and mechanical properties under the conditions of a tool with an internal concave shoulder and for welding T-joints. This not only deepens the understanding of the intrinsic nature of the FSW process but also provides a critical theoretical basis and experimental data support for the integrated precision design and control of the "process-structure-performance" relationship in FSW of high-reliability aluminum alloy T-joints. At w = 250 rpm, the spiral structure on the concave shoulder surface could not sufficiently stir the material, resulting in root lack-of-penetration. In contrast, at w = 400 rpm, excessive heat input combined with the intense shearing action of the concave structure led to softening in the heat-affected zone (HAZ). This study identified that for this specific concave tool, there exists an optimal medium rotation speed, specifically at w = 300 rpm or 350 rpm at which the stirring tool maximizes its mechanical stirring advantages and achieves grain recrystallization.

The root region of T-joints is a common defect-prone area, where welding speed plays a crucial role in root formation. Through systematic adjustment of the welding speed, this study successfully achieved precise suppression of root defects. As indicated in Lines 30–43 on Page 7 and Lines 1–3 on Page 8 and Lines 9–18 on Page 8 and Lines 1–11 on Page 9 of the revised manuscript, a welding speed of v = 60 mm/min produced T-joints with superior surface and internal quality, establishing a process window for defect-free welding. Furthermore, this study revealed that insertion depth is a sensitive parameter affecting the thinning rate of the welded joint. By optimizing the insertion depth (h) to 0.21 mm, the panel thinning rate was successfully controlled at an industry-advanced level while ensuring root penetration. This also significantly suppressed the height of the back bulge, which is of great significance for aerospace structural components subjected to dynamic loads.

In summary, by analyzing the complex interactions among the tool, parameters, and joint, this study uncovered the essential thermal-mechanical-flow multi-field coupling mechanisms in the FSW of dissimilar aluminum alloys, providing valuable validation data for the development of physical models and numerical simulations of FSW. Meanwhile, the optimized process parameters (w = 350 rpm, v = 60 mm/min, h = 0.21 mm) offer direct guidance for the industrial FSW production of aluminum alloy T-structures (such as large integral stiffened panels and wings) in the aerospace sector, providing a reliable process solution for achieving high-quality, high-consistency and high-efficiency welding.

2. Figure 3. The authors are suggested to elaborate on the design of tool geometry. For example, why the inner concave surface was used for the second shoulder. What difference of microstructure and mechanical properties can be expected as compared to the first shoulder.

We are grateful for the reviewer's suggestion. On lines 12 to 18 of page 4 of the revised draft, we have detailed the geometric structure of the stirring tool, such as: the welding pins were designed as conical thread structures, but the difference was that the stirring tool shoulder structures. The first shoulder was the inner concave surface without spirals, abbreviated as ICSWOS (see Fig. 3a), and the second shoulder was the inner concave surface with spirals, abbreviated as ICSWS (see Fig. 3c). The diameters of the shoulders were 10 mm and the lengths of the stirring pins were 3.8 mm. The diameters of the small and large ends of the stirring pin were 2.8 mm and 3.5 mm, respectively [8]. Fig. 3b presents the structure diagram of the stirring tools. Compared with ICSWOS, the welds produced by ICSWS structure have fewer defects and better mechanical properties than those produced by ICSWOS structure. Such as：Fig. 5a and b, Fig. 5c-l, Fig. 5m-v, and Fig. 5w-x, respectively, show the surface morphology of joint welds obtained under conditions of w = 250 rpm, w = 300 rpm, w = 350 rpm, and w = 400 rpm, with their defect types listed in Table 4. For convenience of description, K, M, Y, and N represent roughness, groove, channel, and flash defects, respectively, with subscripts indicating inconsistencies in the severity of joint weld defects. From Fig. 5, it can be observed that the surface morphology of welds in Fig. 5d, f, h, j, l, n, p, r, and t is significantly better than the corresponding Fig. 5c, e, g, i, k, m, o, q and s, respectively. This phenomenon demonstrates that welds produced using the ICSWS structure exhibit fewer defects than those created with the ICSWOS structure. This improvement stems from the shoulder’ s enhanced heat generation, which arises from dual thermal mechanisms: frictional heating and plastic deformation heat. Details are provided in lines 3 through 12 on page 7 of the revised manuscript.

3. Figure 7. The text in the figure is not clear. The authors are suggested to improve the quality of figures.

We are grateful for the reviewer's comment. Based on the reviewers' suggestions, we modified the text in the figure and improved the graphic quality, as shown in the following figure. The picture is modified in line 14 on page 9 of the revised manuscript.

Fig. 7 SEM images of T-joints at (a) h = 0.18mm, (b) h = 0.19mm, (c) h = 0.2mm, (d) h = 0.21mm, (e) h = 0.22mm, and (f) h = 0.23mm

4. While the study focuses on microstructural evolution, the manuscript does not analyze deformation or fracture mechanisms, which are crucial to understanding the observed strength-ductility synergy. Fractography, strain localization analysis, or failure mode discussions would greatly enhance the completeness and applicability of the conclusions.

We are grateful for the reviewer’s suggestion. We have described in detail the fracture mechanism of T-joints, such as: from Fig. 12, it can be observed that the fracture structure exhibits shear bands and dimples, along with tear ridges. Further observations from Fig. 12a show a large relatively flat area on the section of the specimen, with fewer and shallower dimples. Part of the tear-like dimples forms obvious tear ridges (seen in the 1000x magnified image in Fig. 12b). Moreover, from Fig. 12c and Fig. 12d, it is evident that the number of dimples and tear ridges for the optimal parameter combination Ü is significantly more than that of Ï, and the depth of dimples is also greater than Ï. Greater depth and quantity of microscopic dimples, along with more tear ridges and longer lengths, correspond to higher strength and better plasticity on a macroscopic scale. This phenomenon aligns with the changes observed in macroscopic tensile properties, as demonstrated in Fig. 8. Consequently, the primary fracture mode in the NZ of T-joints is ductile fracture [39]. The underlying mechanism lies in the FSW process where the NZ undergoes severe plastic deformation and dynamic recrystallization, resulting in significant grain refinement. The refined grains possess higher specific surface areas and increased grain boundaries, which effectively impede crack propagation and enhance the joint's plastic deformation capacity. When subjected to external forces, fine dispersions of θ′/T1/η′ precipitates within the NZ grains and along grain boundaries. These hard, brittle particles act as primary sites for microvoid nucleation. During plastic deformation, decohesion at the matrix-precipitate interface or fracture of the precipitates themselves initiates the voids, which then grow and coalesce to form the characteristic dimples seen in Fig. 12d. Meanwhile, the high density of grain boundaries provides numerous nucleation sites for microvoids, leading to this characteristic ductile fracture morphology. In short, due to the homogeneous microstructure and fine grain size in the NZ, the nucleation and coalescence of microvoids occur relatively uniformly. This produces finer, densely distributed dimples that collectively manifest as the signature dimple fracture morphology on the fracture surface. For detailed information, please refer to lines 4 to 25 on page 14 of the revised manuscript.

Author Response File: Author Response.docx