1. Introduction
For marine plants operating at cryogenic temperatures, high-grade steel is required to withstand external environmental conditions, and weld joints must demonstrate sufficient low-temperature toughness at the operating temperature. While impact toughness is widely used for fracture toughness evaluation, there is an increasing demand for CTOD (Crack Tip Opening Displacement) assessments. Ensuring adequate fracture toughness necessitates considering various factors, including the microstructure of weld joints, welding conditions, welding processes, and diverse groove shapes such as square-type and V-type grooves [1]. Since groove shapes significantly impact productivity, smaller grooves are preferred as long as there are no quality issues with the weld joints.
Numerous studies have investigated the effects of groove shapes on weld joint characteristics. Kim, K.S. et al. [1] studied the impact of V, B, and X-groove shapes on the residual stress distribution of hot-rolled SM490A steel. They found that V and B groove shapes, which exhibit symmetrical thermal cycles, showed symmetrical distributions of residual stress and deformation. In contrast, X-groove shapes exhibited minimal deformation, making them the most favorable in terms of deformation characteristics. Soh, N.H. [2] investigated the effects of single-side and double-side welding on the residual stress distribution of narrow-gap welds in SUS316LN steel. The study revealed that for both welding methods, axial and circumferential residual stresses gradually increased from the base material toward the weld toe, peaking at the boundary between the base material and the weld joint. Single-side welding exhibited higher residual stress compared to double-side welding.
Pandey et al. [3,4] studied the microstructural changes and Charpy impact toughness characteristics of P91 steel under different heat treatment conditions and groove shapes. They found that before PWHT (Post Weld Heat Treatment), the V-groove had higher tensile strength than the Y-groove. However, after PWHT, the V-groove exhibited lower tensile strength than the Y-groove. Sharma and Shahi [5] investigated the mechanical and microstructural properties of weld joints in quenched and tempered (Q&T) steel with a thickness of 15T for Double-Vee, U, and C-groove shapes. The study showed that Double-Vee grooves resulted in higher ultimate strength ratios, yield strength, and maximum tensile strength than other groove shapes. Singh et al. [6] studied the tensile strength and impact toughness of mild steel weld joints with varying groove angles using the SMAW process. They concluded that a groove angle of 50°, smaller than 60° or 70°, produced optimal results. P. Li et al. [7] examined the distribution of intermetallic compound (IMC) layers, microstructure, and mechanical property changes in laser beam spot welds with I, V, and Y-groove shapes. They found that V-grooves produced the most homogeneous IMC layers and exhibited the smallest temperature gradient.
In some studies, research has been conducted not only on the effects of cooling rate and microstructure on the toughness of weld zones but also on the shape of the weld zone [8,9,10]. Notably, Glover, A.G., et al. [8] demonstrated that in the weld zone of a low-alloy steel (0.13 wt%C), as the heat input increases, the microstructure transforms from martensite to upper bainite and acicular ferrite. Additionally, it was reported that with an increase in heat input, MA (Martensite–Austenite constituent) transitions from a mixture of carbides and MA to ultimately forming a carbide-dominant structure.
Furthermore, Ikawa [11] studied the formation of MA under varying cooling rates, reporting that elongated MA forms under rapid cooling conditions, while relatively massive (blocky) MA is formed under slower cooling rates. However, there is a lack of data directly applicable to weld zones produced by modern high-toughness welding materials and high-heat-input welding techniques. Most researchers have focused on evaluating the toughness of the heat-affected zone (HAZ) rather than the weld zone itself using simulation techniques [12,13,14,15]. As a result, while significant research has been conducted on brittle regions such as CGHAZ (coarsened grain heat-affected zone) and ICHAZ (intercritical heat-affected zone) within the heat-affected zone, studies on weld metal are relatively scarce due to its metallurgical heterogeneity. Consequently, research on the fracture toughness characteristics in relation to changes in weld geometry is also limited.
This study aims to evaluate which weld geometries are advantageous for fracture toughness in submerged arc welding (SAW) as part of the development of welding materials suitable for low temperatures (–60 °C), which are in high demand for offshore structures. For this purpose, the microstructure and fracture toughness characteristics of SAW joints were evaluated using both experimental and numerical methods, focusing on K- and X-groove geometries commonly applied to thick plates. The volume fractions of acicular ferrite, grain boundary ferrite, and MA phases were measured to experimentally assess their correlation with fracture toughness. Additionally, the study evaluated the effects of cooling rate on the location of the local brittle zone (LBZ) where brittle cracks initiate, as well as the influence of groove geometry changes through numerical analysis.
2. Experimental Methods
2.1. Base Material and Welding Conditions
The base material used was S460NL, which is widely utilized for marine structural steel plates. The properties of the steel (S460NL) are shown in ASTM A572/A572M-21e1 [16]. The welding materials used were EF–200CT and KD–50, which are KISWEL products (
2.2. Mechanical Property Evaluation
To measure fracture toughness, CTOD testing was conducted at –60°C in accordance with the ISO 15653:2018 standard [17]. The test equipment used included a TSH-200 (Daeyoung M.T.C., Hwaseong Si, Republic of Korea) for fatigue pre-cracking, a DYHU-1000 (Daeyoung M.T.C., Hwaseong Si, Republic of Korea) for 3-point bending, and a COD gauge (RA-15SAL001, Tokyo Sokki Kenkyujo Co., Tokyo, Japan). For both K-groove and X-groove specimens, the NP notch direction was applied. The notch was machined 2 mm away from the fusion line for the K-groove specimens and at the center of the bead for the X-groove specimens. The notch length of the specimen was machined considering a fatigue crack length ratio (a/W) of 0.50. Table 2 provides the information on the CTOD specimens. Tensile testing was performed on one round bar specimen each, extracted from the front and back weld metal sections, in accordance with the DIN EN 10025–4 [18] as Table 3.
2.3. Observation of Welding Microstructure
The weld microstructure was observed after specimen preparation, using a 3% Nital solution (100 mL ethanol + 3 mL nitric acid) for etching for approximately 10 s. The analysis was conducted with an optical microscope at a magnification of ×200. Additionally, the fractions of grain boundary ferrite (GBF) and acicular ferrite (AF) were measured in the columnar zone and reheated zones (coarse-grained, fine-grained, intercritical, and subcritical zones: CG, FG, IC, SC) within the weld using the X-Solution image analysis software (IMT iSolution Lite, Ver. 22.5). As shown in Figure 2, the measurements were taken from three passes in each region: the front weld, the root, and the back weld.
To observe the MA (Martensite–Austenite) constituents, the Alkaline Sodium Picrate method was applied using a mixed solution of 100 mL distilled water and 25 g NaOH. Electrolytic etching was performed at approximately 6 V for 2–3 min. As shown in Figure 2, measurements were taken at three points each from the weld metal and reheated zones of the front weld, root, and back weld sections. The fraction of MA constituents was observed at a magnification of 1000 x and measured using the X-Solution software (IMT iSolution Lite, Ver. 22.5). The fraction of MA constituents was compared based on changes in heat input and interpass temperature.
2.4. Cooling Rate Analysis
To compare the cooling rates of K-groove and X-groove welds, the commercial software Marc 2011 from MSC [19] was used for numerical analysis. Considering the actual welding conditions, a model for heat source analysis was created as shown in Figure 3.
Figure 3a illustrates the analysis model for the X-groove shape. The model dimensions are 50 mm in welding length, 180 mm in width, and 80 mm in thickness, with a three-dimensional moving heat source applied. The minimum element size for the weld metal and heat-affected zone (HAZ) is 0.2 mm × 0.7 mm × 5 mm, which ensures sufficient mesh refinement. Additionally, the model consists of a total of 6350 elements and 7370 nodes. Figure 3b illustrates the analysis model for the K-groove shape. The model dimensions are 50 mm in welding length, 200 mm in width, and 100 mm in thickness, with a three-dimensional moving heat source applied. Numerical analysis was performed using the same method with the K-groove model. The minimum element size for the weld metal and heat-affected zone (HAZ) is 0.2 mm × 0.7 mm × 5 mm, which ensures sufficient mesh refinement. Additionally, the model consists of a total of 9300 elements and 11,532 nodes.
Figure 4 illustrates the boundary conditions applied in the heat transfer analysis. The analysis considers heat flux, thermal conductivity, specific heat, and convective heat transfer by variations with temperature.
The heat source used in the analysis replicated the actual welding conditions to perform a thermal distribution analysis. Additionally, high-temperature material properties for the base material and welding material, varying with temperature, were obtained and applied to the analysis to ensure accurate thermal distribution results.
3. Experimental Results
3.1. CTOD and Mechanical Characteristics
Table 4 shows the CTOD test results for the K-groove and X-groove welds, which were measured as 0.821 mm and 0.1325 mm, respectively. The CTOD value for the X-groove specimen was lower compared to the K-groove specimen. Figure 5 illustrates the fracture surfaces and initiation points of the CTOD test specimens, and Table 5 indicates the locations of fracture initiation. It was observed that most brittle fractures initiated in the reheated zones caused by subsequent weld passes. More fracture initiation points were detected in the X-groove specimens compared to the K-groove specimens.
The mechanical property test results are presented in Table 6, showing that the tensile strength of the X-groove specimens was slightly higher than that of the K-groove specimens.
3.2. Results of Weld Microstructure Observation
Figure 6 shows the microstructures of the K-groove and X-groove specimens etched with a 3% Nital solution, observed using an optical microscope at a magnification of 200×. The results of measuring the fractions of acicular ferrite (AF), grain boundary ferrite (GBF), and Widmanstätten ferrite (WF) are presented in Table 7 and Table 8. The AF fractions for the K-groove and X-groove specimens were 83.14% and 90.65%, respectively, while the GBF fractions were measured as 16.75% and 9.28%, respectively. It was observed that the K-groove specimens had a lower fraction of AF and a higher fraction of GBF compared to the X-groove specimens. This is considered to be due to the slower cooling rate resulting from the higher overall heat input in the K-groove compared to the X-groove.
Figure 7 and Figure 8 show the images, taken at a magnification of 1000× using an optical microscope, of the MA microstructure in the weld metal and reheated zones corroded with the Alkaline Sodium Picrate method, based on the groove geometry. The bright, glowing structures observed are identified as MA phases, which are formed at the ferrite interfaces. The results of the MA phase fractions measured at various locations are presented in Table 9 and Table 10. It was observed that the MA phase is predominantly distributed in the reheated zones’ local brittle zones (LBZs), specifically in the ICHAZ and SCHAZ. Depending on the groove geometry, the MA phase fraction in the ICHAZ of the reheated zones was measured as 0.09% for the K-groove and 0.60% for the X-groove, while in the SCHAZ, it was measured as 0.09% and 0.56%, respectively, showing significantly higher values in the X-groove specimens.
4. Discussion
4.1. Relationship Between Fracture Toughness (CTOD) and Microstructure
The relationship between the AF fraction and CTOD based on groove shape is shown in Figure 9a, while the relationship between the GBF fraction and CTOD is illustrated in Figure 9b. It was observed that as CTOD decreases, the AF fraction increases, demonstrating an inverse relationship, while the GBF fraction decreases, showing a direct relationship.
According to Evans and Bailey [20], AF structures with interlocking configurations significantly enhance low-temperature toughness, whereas GBF with coarse grains and WF structures, which have low crack resistance, typically act as paths for crack propagation, thereby greatly affecting toughness. However, Zhang and Farrar [21] and Dallam et al. [22] reported that when the AF fraction exceeds approximately 90%, it can have a detrimental effect on low-temperature toughness.
Therefore, the lower CTOD of the X-groove compared to the K-groove is likely due to the excessive AF content in the X-groove specimens.
The relationship between the fraction of M–A constituents in each zone of the weld and CTOD is shown in Figure 10. The results indicate that as the amount of M–A increases, CTOD tends to decrease, consistent with previous studies reporting that M–A acts as a source of cracks [23]. Additionally, it was observed that the X-groove specimens, which had lower CTOD values compared to the K-groove specimens, exhibited a relatively higher formation of M–A constituents. This is likely due to the slower cooling rate of the X-groove specimens, which provided sufficient time for M–A formation, identified as the primary factor contributing to the reduced CTOD.
Figure 11 and Figure 12 show the changes in M–A fraction with increasing interpass temperature for the K-groove and X-groove specimens, respectively. Both specimens exhibited a tendency for the M–A fraction to increase at higher interpass temperatures above 200 °C, which is attributed to the influence of slower cooling rates.
Figure 13 illustrates the changes in M–A fraction with increasing heat input for the X-groove specimens. The results indicate that as the heat input increases, the M–A fraction tends to rise, showing a particularly sharp increase at heat inputs above 4 kJ/mm. Generally, it is well known that the M–A fraction increases and toughness decreases at higher heat inputs compared to lower heat inputs, and that M–A in reheated zones has a stronger correlation with toughness characteristics than in solidification zones [23].
Therefore, the increase in heat input for the X-groove specimens likely resulted in a significant rise in the fraction of massive M–A constituents, which had a substantial negative impact on CTOD.
The reheated zone fractions at the notch locations of the K-groove and X-groove macro specimens were measured and compared in Table 11. The reheated zone fraction was 46.6% for the X-groove specimens and 41.7% for the K-groove specimens, indicating that the reheated zone at the notch location of the X-groove was approximately 12% higher than that of the K-groove.
This difference is attributed to the notch line of the X-groove being located at the center of the weld, resulting in overlapping passes that significantly increased the reheated zone compared to the K-groove, where the notch is positioned at the fusion line +2 mm.
Consequently, as shown in Figure 4, the lower CTOD of the X-groove specimens can be attributed to the higher distribution of the weak reheated zones, which serve as primary fracture initiation points.
4.2. Comparison of Cooling Rates Following Improvements
The results of thermal analysis at the same locations relative to thickness are shown in Figure 14 and Figure 15. Figure 14 compares the 7th pass of the K improvement with the 5th pass of the X improvement, revealing that the cooling rate in the 800–500 °C range is faster in the K improvement compared to the X improvement. Figure 15 compares the 23rd pass of the K improvement with the 28th pass of the X improvement. When cooled to approximately 1000 °C, the cooling rate of the K improvement is slightly faster, but the trends are nearly identical at lower temperatures.
In general, welds with high heat input are known to cool more slowly than those with low heat input [21]. However, the cooling rate analysis of the Notch line indicates that the K improvement specimen, with a heat input of 51 kJ/cm, exhibited a faster cooling rate compared to the X improvement specimen, which had a heat input of 38.6 kJ/cm. This is likely due to the location of the notch in the K improvement specimen being relatively close to the base material at a lower temperature, causing a significant portion of the heat to transfer to the base material. In contrast, in the X improvement specimen, heat transfer was delayed due to the higher temperature of surrounding weld passes near the center line, resulting in a slower cooling rate compared to the K improvement.
Therefore, during the slower cooling of the X improvement specimen, there was sufficient time for the formation of M–A constituents, which negatively impact low-temperature toughness. As a result, it is believed that significantly more M–A constituents formed in the X improvement specimen compared to the K improvement specimen. Additionally, the increased formation of M–A constituents appears to have led to a substantially lower CTOD in the X improvement specimen compared to the K improvement specimen.
5. Conclusion Remarks
This study evaluated the microstructure and changes in M–A constituents of welds in S460NL steel according to the K and X improved designs. The results of CTOD tests conducted at −60 °C were compared and evaluated against the fractions of each microstructure. The findings are summarized as follows:
Tensile Strength and Hardness: The tensile strength of the X improvement specimen was measured at 592 MPa, higher than the 574 MPa of the K improvement specimen, showing a similar trend in hardness.
Reheat Affected Zone: Due to the influence of the improved design, the proportion of the reheat affected zone in the X improvement specimen was approximately 12% higher than that in the K improvement specimen. Fracture location analysis revealed that relatively more fractures occurred in the ICHAZ and SCHAZ regions of the X improvement specimen.
Microstructure Observation: The X improvement specimen exhibited a higher proportion of AF (above 90%) compared to the K improvement specimen and a lower distribution of GBF. This excessive AF fraction appears to have negatively impacted toughness.
M–A Constituent Analysis: M–A constituents, which significantly affect transition temperature, were below 0.1% in the K improvement specimen but substantially higher in the X improvement specimen. This was identified as the primary cause of the lower CTOD observed in the X improvement specimen.
Cooling Rate Analysis: The cooling rate of the X improvement specimen was slower than that of the K improvement specimen. This slower cooling allowed sufficient time for the formation of M–A constituents, which is believed to have contributed to the reduced CTOD.
This study considered the effect of groove geometry on fracture toughness for ultra-thick plates exceeding 80 mm. In actual industrial applications, the K-groove in SAW (Submerged Arc Welding) joints has been shown to offer advantages in terms of both welding productivity and fracture toughness. While the X-groove is generally preferred in most SAW joints to prevent defects, the practicality of a method where the root weld is performed using FCAW (Flux-Cored Arc Welding) and the fill and cap layers are welded using SAW is expected to increase for K-groove applications.
Additionally, in cases where the thickness decreases, changes in cooling rates may lead to variations in the microstructure fractions. Based on the findings of this study, a systematic review of the effects of groove geometry on reduced steel thickness will be conducted. Furthermore, future research will focus on the fracture toughness characteristics of weld joints in cryogenic environments, such as those involving liquid hydrogen.
Conceptualization, Y.-T.S. and M.S.; methodology, Y.-T.S.; validation, Y.-T.S., M.S. and C.-J.J.; formal analysis, C.-J.J., M.S. and S.-H.B.; investigation, C.-J.J., M.S. and S.-H.B.; writing—original draft preparation, G.A.; writing—review and editing, Y.-I.P.; visualization, S.-H.B., M.S. and G.A.; supervision, Y.-I.P.; project administration, Y.-I.P.; funding acquisition, Y.-I.P. All authors have read and agreed to the published version of the manuscript.
The original contributions presented in this study are included in the article. Further inquiries can be directed to the corresponding author.
Chang-Ju Jung is currently employed by HD Hyundai Heavy Industries; however, this study was conducted during his time as a graduate student. Myungrak Son is currently employed by KISWEL Co., Ltd. and participated in this study as a collaborator. The remaining authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.
The following abbreviations are used in this study:
| SAW | Submerged Arc Welding |
| CTOD | Crack Tip Opening Displacement |
| MA | Martensite–Austenite |
| GBF | Grain boundary ferrite |
| AF | Accicular ferrite |
| WF | Widmanstätten ferrite |
| CG | Coarse-grained |
| FG | Fine-grained |
| IC | Intercritical |
| ICHAZ | Intercritical heat-affected zone |
| SC | Subcritical |
| SCHAZ | Subcritical heat-affected zone |
| LBZ | Local brittle zone |
Footnotes
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Enlarge this image.Figure 2. Measuring position of microstructure: (a) K-groove; (b) X-groove (unit: mm).
Enlarge this image.Figure 3. Generation of heat source for measuring cooling rate: (a) K-groove; (b) X-groove.
Enlarge this image.Figure 4. Thermal properties according to temperature change for thermal analysis.
Enlarge this image.Figure 6. Microstructure of each specimen’s weld metal: (a) K-groove; (b) X-groove.
Enlarge this image.Figure 9. Comparison of microstructure and CTOD value: (a) AF and CTOD; (b) GBF and CTOD.
Enlarge this image.Figure 11. Change in the fraction of M–A according to the increase in interpass temperature of the K-groove specimen.
Enlarge this image.Figure 12. Change in the fraction of M–A according to the increase in interpass temperature of the X-groove specimen.
Enlarge this image.Figure 13. Change in the fraction of M–A according to the increase in heat input of the X-groove specimen.
Enlarge this image.Figure 14. Comparison of the No. 7 pass of K-groove and No. 5 pass of X-groove: the red line is X-groove, and the blue line is K-groove.
Enlarge this image.Figure 15. Comparison of the No. 23 pass of K-groove and No. 28 pass of X-groove: the red line is X-groove, and the blue line is K-groove.
Welding condition of K- and X-groove specimens.
| Parameter | K-Groove | X-groove |
|---|---|---|
| Thickness [mm] | 80 | 100 |
| Welding position | Flat(1G) | Flat(1G) |
| Heat input [kJ/cm] | 51.0 | 38.6 |
| Preheat and interpass temp. [°C] | Avg. 196 | Avg. 171 |
Dimensions of CTOD test specimen (unit: mm).
| Groove Type | Thickness | Width | Notch Length |
|---|---|---|---|
| K-groove | 74.90 | 74.99 | 33.00 |
| X-groove | 98.00 | 98.00 | 44.10 |
Steel properties of S460NL.
| Steel Grade | Chemical Composition [%] | |||||||||||||
|---|---|---|---|---|---|---|---|---|---|---|---|---|---|---|
| S460NL | C | Si | Mn | P | S | Nb | V | Alges | Ti | Cr | Ni | Mo | Cu | N |
| 0.22 | 0.65 | 0.95–1.80 | 0.030 | 0.025 | 0.06 | 0.22 | 0.015 | 0.06 | 0.35 | 0.85 | 0.13 | 0.60 | 0.027 | |
Results of the CTOD test at –60 °C.
| Groove Type | Temperature [°C] | CTOD, δ [mm] | ||||
|---|---|---|---|---|---|---|
| 1 | 2 | 3 | 4 | Avg. | ||
| K-groove | –60 | 0.58 | 0.48 | 1.41 | - | 0.82 |
| X-groove | 0.07 | 0.18 | 0.16 | 0.12 | 0.13 | |
Initiation position of fracture of CTOD specimens.
| Counts [EA] | IC HAZ | SC HAZ | CG HAZ | |
|---|---|---|---|---|
| K-groove | K-1 | 1 | 0 | 0 |
| K-2 | 0 | 0 | 0 | |
| K-3 | 1 | 1 | 0 | |
| X-groove | X-1 | 1 | 1 | 2 |
| X-2 | 1 | 1 | 1 | |
| X-3 | 2 | 1 | 1 | |
Results of the mechanical property test.
| AWS A5.17 F7A(P)8-EH14 | K-Groove | X-Groove | |
|---|---|---|---|
| Yield strength [MPa] | ≥400 | 550 | 517 |
| Tensile strength [MPa] | 480/660 | 574 | 592 |
| Elongation [%] | ≥22 | 33 | 33 |
Area fraction of microstructures of K-groove specimen weld metal.
| Area Fraction [%] | AF | GBF | WF | |
|---|---|---|---|---|
| F (first weld) | 81.39 | 18.52 | 0.09 | |
| R (root weld) | 91.60 | 8.40 | 0.00 | |
| S (second weld) | 79.76 | 20.24 | 0.00 | |
| Statistics | Mean | 83.14 | 16.75 | 0.11 |
| Max. | 91.60 | 22.32 | 0.57 | |
| Min. | 77.50 | 8.40 | 0.00 | |
| Median | 82.41 | 17.47 | 0.09 | |
| Standard deviation | 3.66 | 3.65 | 0.14 | |
Area fraction of microstructures of X-groove specimen weld metal.
| Area Fraction [%] | AF | GBF | WF | |
|---|---|---|---|---|
| F (first weld) | 85.90 | 13.90 | 0.20 | |
| R (root weld) | 92.80 | 7.12 | 0.08 | |
| S (second weld) | 95.26 | 4.74 | 0.00 | |
| Statistics | Mean | 90.65 | 9.28 | 0.07 |
| Max. | 95.26 | 14.27 | 0.36 | |
| Min. | 85.53 | 4.58 | 0.00 | |
| Median | 91.02 | 8.75 | 0.04 | |
| Standard deviation | 3.16 | 3.16 | 0.10 | |
Results of the M–A constituent area fraction of the K-groove specimen.
| Area Fraction [%] | WN | CG | FG | IC | SC | |
|---|---|---|---|---|---|---|
| F (first weld) | 0.08 | 0.12 | 0.04 | 0.36 | 0.34 | |
| R (root weld) | 0.01 | 0.00 | 0.00 | 0.02 | 0.02 | |
| S (second weld) | 0.02 | 0.01 | 0.00 | 0.01 | 0.01 | |
| Statistics | Mean | 0.03 | 0.02 | 0.02 | 0.09 | 0.09 |
| Max. | 0.08 | 0.17 | 0.12 | 0.53 | 0.59 | |
| Min. | 0.00 | 0.00 | 0.00 | 0.00 | 0.00 | |
| Median | 0.02 | 0.01 | 0.01 | 0.03 | 0.02 | |
| Standard deviation | 0.02 | 0.04 | 0.03 | 0.13 | 0.15 | |
Results of the M–A constituent area fraction of the X-groove specimen.
| Area Fraction [%] | WN | CG | FG | IC | SC | |
|---|---|---|---|---|---|---|
| F (first weld) | 0.38 | 0.14 | 0.21 | 0.45 | 0.68 | |
| R (root weld) | 0.17 | 0.22 | 0.12 | 0.32 | 0.63 | |
| S (second weld) | 0.05 | 0.07 | 0.02 | 0.39 | 0.17 | |
| Statistics | Mean | 0.22 | 0.26 | 0.23 | 0.60 | 0.56 |
| Max. | 1.17 | 1.36 | 1.42 | 2.50 | 1.62 | |
| Min. | 0.01 | 0.01 | 0.02 | 0.09 | 0.11 | |
| Median | 0.14 | 0.13 | 0.12 | 0.36 | 0.41 | |
| Standard deviation | 0.25 | 0.34 | 0.31 | 0.61 | 0.48 | |
Fraction of reheated region at notch line of K- and X-groove specimens.
| Parameter | K-Groove | X-Groove |
|---|---|---|
| Location of notch line | Fusion line + 2 mm | Center of weldment |
| Fraction of reheated region [%] | 41.7 | 46.6 |
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; Chang-Ju, Jung 1 ; Seong-Han Bae 1 ; An, Gyubaek 2
; Son, Myungrak 3 ; Park, Young-Il 1
1 Department of Naval Architecture and Ocean Engineering, Dong-A University, Busan 49201, Republic of Korea;
2 Department of Naval Architecture and Ocean Engineering, Chosun University, Gwangju 61452, Republic of Korea;
3 R&D Center, KISWEL Co., Ltd., Busan 47018, Republic of Korea;