Experimental Study on the Mechanical Degradation of Q420C Steel Welded Joints after Simulated Fire Exposure

Yang, Liguo; Han, Bingfeng; Lu, Yuhao; Su, Zhe

doi:10.1590/1980-5373-MR-2026-0100

Abstract

This study investigates the evolution of toughness in Q420C low-alloy steel after simulated fire exposure. Specimens were heated to 200, 400, 600, and 800°C for 30 minutes, followed by either natural cooling or water quenching. Charpy V-notch tests were conducted over a temperature range of −60 to 20°C. Toughness exhibited a three-stage trend: a significant decline at 400°C, substantial improvement at 600°C (39–53% higher than at 400°C), and sharp deterioration at 800°C The heat-affected zone consistently outperformed the weld zone in toughness, while cruciform butt-welded joints showed inferior low-temperature performance, with the weld zone being most sensitive. Notably, water quenching at 400°C reduced embrittlement, whereas air cooling yielded better toughness at 600°C and 800°C, indicating a temperature-dependent effect of cooling methods on the post-fire fracture resistance of welded joints. Scanning electron microscope analysis revealed ductile dimples in the 600°C group, while the 400°C and 800°C groups were dominated by cleavage features. The cruciform butt weld zone exhibited the most severe cleavage characteristics at low temperatures.

Keywords:
Welded joints; Charpy impact toughness; High-temperature exposure; Cooling method

1. Introduction

High-strength steel, as a key material in high-performance construction, is integral to modern engineering and supports the achievement of global “dual carbon” targets, significantly influencing the steel industry¹. Its superior load-bearing capacity, stiffness, weather resistance, and reliability have led to widespread use in high-rise buildings, long-span bridges, shipbuilding, and industrial facilities²-⁴. However, exposure to high temperatures during fires significantly reduces the yield strength and elastic modulus of high-strength steel, severely compromising structural integrity and increasing the risk of partial or total collapse, thereby endangering lives and property⁵.

To replicate post-fire mechanical property variations in steel, numerous experimental studies have been conducted. Hua et al.⁶ studied the residual mechanical behavior of Q1100 ultra-high strength steel by high temperature (300-1000 °C) treatment, tensile test and metallographic observation. The results show that the metallographic change (cementite precipitation, etc.) at 300-600 °C leads to a decrease in strength and an increase in plasticity, and the strength of cold-formed steel decreases more significantly. Xue et al.⁷ studied the bond behavior between bimetal steel bars and seawater concrete by pull-out test after high temperature with different cooling methods (air / water). At 200-400 °C, the bond strength decreased, and the water cooling decreased more, with a maximum decrease of 63.54% at 400 °C. Wang et al.⁸ studied the mechanical properties of Q1100 ultra-high strength steel by high temperature (300-1000 °C) water cooling, tensile and metallographic analysis. The strength decreases gradually and the plasticity increases significantly at 300-700 °C, and the plasticity reaches the peak at 700 °C. Cold-formed steel has higher strength but lower plasticity. Qiang et al.⁹,¹⁰ investigated S460, S690, and S960 steel, reporting that elastic modulus remained largely unaffected at 600°C but declined at high temperatures, with fire exposure affecting each grade differently. Tan et al.⁵ reviewed the impact toughness of Q235-B and 20# steel treated with CO₂ extinguishing agents, noting increased brittleness due to rapid cooling and reduced toughness caused by internal stress and microstructural heterogeneity when applied after 3 minutes of combustion. Beyond static mechanical properties, researchers have also explored seismic performance¹¹,¹² and fatigue stability¹³-¹⁷.

The preceding discussion highlights the varied effects of fire on the mechanical properties of steel. Among these properties, impact toughness serves as a critical measure of a structure's resistance to dynamic loads under extreme conditions, directly influencing material selection and structural safety. To assess fracture performance in steel components and connections, numerous studies have evaluated impact fracture energy. Lee et al.¹⁸ examined structural steel welds at low and ambient temperatures, while Liu et al.¹⁹ investigated the impact toughness of Q235 steel following salt spray corrosion. Tong et al.²⁰ conducted impact tests on high-strength steel of varying grades, and Li et al.²¹ and Zhang et al.²² performed Charpy V-notch impact tests on Q420 steel under similar temperature conditions. Jia et al.²³ analyzed crack propagation in steel members and connections (JIS SS400, SM490A, SM570TMC) through simply supported beam impact tests at room temperature. On the other hand, regarding the Charpy impact energy of steels at high temperatures, some researchers have conducted Charpy impact tests on SA508-II and MA956 ODS steels, which are commonly used in nuclear power plants, within the temperature range of 0–450 °C²⁴,²⁵.

Q420C steel is extensively used in China's major infrastructure projects, including long-span bridges and super-high-rise buildings. However, exposure to high temperatures and subsequent cooling during fires can severely compromise its mechanical properties, particularly at welded joints, which are inherently vulnerable. The uneven degradation in the weld and heat-affected zones (HAZs) poses a direct risk to structural integrity. Current research lacks comprehensive data support on the post-fire evolution of impact toughness in these regions, limiting its applicability in post-disaster safety evaluations and repair planning. However, the existing literature predominantly focuses on the static tensile properties of post-fire structural steels (e.g., Q235, S690) or the impact toughness of base metals under ambient temperature or standard cooling conditions⁵,⁸. There remains a critical lack of data concerning the low-temperature impact toughness and ductile-to-brittle transition behavior of welded joints—especially complex configurations like cruciform butt welds—subjected to simulated fire exposure and dual-mode cooling (air vs. water quenching). While rapid cooling (like water quenching) has been shown to increase the tensile strength of some steels⁸, its effect on the fracture toughness of heat-affected zones (HAZs) with heterogeneous microstructures is not fully understood. To address this specific knowledge gap, this study systematically investigates

2. Overview of the Experiment

Q420C steel plates were tested using two joint types: butt welds and cruciform butt welds. Standard Charpy V-notch specimens were extracted from the weld and HAZs. Samples were preheated to 200, 400, 600, and 800°C; held for 30 minutes; and then cooled by either air (A) or water (W). Low temperature impact testing was performed at 20, 0, −20, −40, and −60°C, with each condition tested in triplicate and averaged. Scanning electron microscopy was employed to examine fracture surface morphology, enabling a systematic evaluation of the effects of welding configuration, thermal exposure, and cooling method on impact toughness.

2.1. Test specimen

The 8 mm-thick hot-rolled Q420C steel plates used in this study were manufactured by Hualing Lian Yuan Iron & Steel Co., Ltd. The material met the mechanical, chemical, and process requirements specified in GB/T 1591-2018 (High-strength low alloy structural steel). Detailed chemical composition data are provided in Table 1.

Thumbnail

Table 1
Q420C steel plate smelting analysis results.

Welding was performed using the shielded metal arc welding method with SH. J557 fine-grained steel electrodes provided by Shanghai Welding Equipment Co., Ltd. The deposited metal exhibited a yield strength not less than 460 MPa, a tensile strength exceeding 550 MPa, and an elongation after fracture greater than 17%, meeting the requirements for all-position welding under DC reverse polarity and ensuring high strength and ductility in the joints. Specimen preparation followed the standards of JG/T 288-2013 (Test method for impact toughness of welded joints) and GB/T 2650-2008 (Impact test methods for welded joints). Two joint types were designed: butt welds and cruciform butt welds, employing V-groove and single-sided V-groove configurations with full penetration to ensure complete fusion, as illustrated in Figure 1. The welding process parameters are summarized in Table 2. All welds met Grade 1 quality standards and were evaluated using ultrasonic and radiographic non-destructive testing, by GB 50611-2011 (Code for welding of steel structures), to ensure internal defects remained within acceptable limits.

Figure 1
Details of welded joint construction.

Thumbnail

Table 2
Welding process parameters.

V-notch specimens with dimensions of 55 × 10 × 5 mm were prepared per GB/T 229-2020 (Charpy pendulum impact test method for metallic materials). Following surface polishing, the specimens were etched with a 4% nitric acid alcohol solution (Figure 2(a)) and examined under an optical microscope to evaluate surface features and delineate the HAZ. A V-notch was then machined along the rolling direction of the steel plate to ensure alignment of the notch opening with the rolling orientation. Sampling locations for the weld and HAZs are shown in Figure 2(b) and (c), respectively. Based on joint type and sampling region, specimens were categorized into four groups, each tested under four high-temperature conditions, two cooling methods, and five impact temperatures. Three specimens were tested per condition, totaling 480 impact samples, as summarized in Table 3.

Figure 2
Schematic of the test specimen's corrosion morphology and notch location.

Thumbnail

Table 3
Distribution of test specimens.

2.2. Experimental process

The experiment was conducted in two phases: high-temperature treatment and impact testing. Specimens were heated at a constant rate of 10°C/min to target temperatures of 200, 400, 600, and 800°C and then held at the target temperature for 30 minutes to ensure uniform thermal exposure. To simulate a fire suppression scenario, two cooling methods were applied: natural air cooling and water quenching. Following cooling to room temperature, surface morphology changes were observed and recorded.

Impact testing was conducted at five ambient temperatures: 20, 0, −20, −40, and −60°C. To ensure consistency, specimens subjected to the same thermal treatment were cooled simultaneously and held at the target temperature for at least 5 minutes. Tests were performed sequentially from high to low temperatures. During specimen transfer, the contact surfaces of the fixtures and specimens were maintained at consistent temperatures within the cooling medium, with transfer time to 5 seconds. Each specimen was positioned firmly against the anvil, with the notch-centered surface aligned precisely between the anvils to ensure accurate impact by the hammer opposite notch.

2.3. Test equipment

High-temperature treatment was performed using a KSL-1700X box-type sintering furnace (Figure 3(a)), capable of reaching 1700°C with a temperature control accuracy of ±1°C and a programmable heating rate of 0 to −20°C/min, meeting the thermal processing requirements of this study. For low-temperature conditioning during impact testing, a DWC-80 microcomputer-controlled chamber (Figure 3(b)) was employed. Using a three-stage refrigeration system with anhydrous ethanol as the cooling medium, it maintained temperatures from −80 to 20°C with fluctuations within ±0.5°C, ensuring specimens remained at the target temperature for over 5 minutes. Fracture surface morphology was analyzed using a Thermo Fisher APREO 2S scanning electron microscope (Figure 3(c)), equipped with an ultra-high-resolution Schottky field emission electron gun and ETD secondary electron detector, enabling detailed observation of ductile dimples, cleavage steps, and other microstructural features critical for understanding fracture mechanisms under varying treatment conditions.

Figure 3
Test equipment.

3. Results and Analysis

3.1. Appearance of the test piece after different cooling treatments

Post-treatment surface changes in Q420C steel specimens varied with temperature, as illustrated in Figure 4. With increasing temperature, oxidation intensified, gradually altering the surface color from silver-white to deep black. At 200°C, no significant surface changes were observed. Above this temperature, a blue oxide film formed, and the weld zone developed a faint reddish tint. Under natural cooling, surfaces remained relatively smooth with minor rusting, whereas water cooling caused surface layer spalling. At 600°C, a yellowish hue appeared around the weld, accompanied by surface carbide precipitation, which detached after water cooling. At 800°C, extensive carbide formation and a dark black surface were observed, with the carbide layer easily detaching and weld contours becoming prominent. Water quenching at this stage reduced carbide flaking and resulted in a smoother surface than that obtained using natural cooling.

Figure 4
Changes in the appearance of the test piece after different cooling treatments.

Due to thermal expansion during high-temperature treatment, the specimen dimensions slightly exceeded standard specifications. Prior to impact testing, surfaces were ground and polished to restore standard dimensions and surface finish, ensuring data accuracy and consistency.

3.2. Impact energy value

Impact tests were performed at five low-temperature levels using small-sized specimens. Each condition was tested in triplicate, and average values were calculated. To ensure consistency, results were converted to equivalent impact energy values of standard-size specimens. Data in Table 4 are listed as mean values, with the dispersion of data described by coefficient of variation (standard deviation/mean).

Thumbnail

Table 4
Low-temperature impact energy values of materials.

Based on Table 4 and the fitted impact energy curves shown in Fig. 0035, the impact toughness of Q420C steel exhibited distinct variation depending on joint type, sampling location, and high-temperature treatment conditions:

The toughness variation in the weld zone of butt-welded joints followed a three-stage pattern with increasing temperature: initial decline, subsequent improvement, and final deterioration. Except at 400°C, water quenching consistently reduced impact energy by approximately 20% compared with natural cooling. At 200°C, natural cooling slightly enhanced toughness relative to water quenching. At 400°C, toughness decreased markedly, with impact energies falling to 61% (air cooling) and 86% (water cooling) of the values recorded at 200°C. However, water cooling outperformed air cooling overall, yielding an average impact ratio of 1.4 across all test temperatures. Further heating to 600°C led to a notable improvement in toughness, with average impact energies increasing by 39% and 42% under various low-temperature conditions. At −60°C, impact energies reached 82.8 J (air cooling) and 67.1 J (water cooling), demonstrating excellent low-temperature toughness after tempering at 600°C. At 800°C, toughness declined sharply by up to 50%, with water-cooled specimens retaining only 80% of the impact energy observed in naturally cooled counterparts, indicating more severe degradation.
In butt-welded joints, the impact energy trend in the HAZ mirrored that of the weld zone but with consistently higher values, averaging 1.6 times greater. The influence of high-temperature exposure and cooling methods on HAZ toughness was comparatively moderate. At 200°C with natural cooling, the steel demonstrated optimal toughness. At 400°C, impact energy declined 35% under air cooling and 10% under water cooling. Further heating enhanced toughness, with impact energies rising by 30% (air cooling) and 36% (water cooling), though these improvements were slightly lower than those observed in the weld zone. At 800°C, toughness decreased again but less severely, with a reduction of 45% and 48% under air and water cooling, respectively.
In specimens with cruciform butt-welded joints, the impact toughness pattern was similar to that of butt-welded joints but showed greater sensitivity to high-temperature exposure and cooling methods. Across low-temperature conditions, impact energy values declined by approximately 10%, with more significant variations in toughness. At 600°C, the weld zone exhibited substantial improvement, with impact energies increasing by 46% under air cooling and 53% water cooling, both exceeding the corresponding gains in butt-welded joints. In contrast, at 400 and 800°C, toughness declined more sharply. While the HAZ followed similar trends to the weld zone, the magnitude of variation was smaller but still greater than that observed in the HAZ of butt-welded specimens.

The repeatability of the Charpy impact test was evaluated by analyzing the standard deviation of threeparallel samples. The coefficient of variation (COV) of impact energy is mainly between 5% and 15%,indicating good experimental consistency. Especially in the upper and lower platform regions (e.g, 20°Cand -60°C), due to stable fracture mechanisms (dimple or cleavage). Conversely, in the ductile-brittle transition zone (-20°C to -40°C), the fracturepath is highly sensitive to microstructural inhomogeneity (e.g., inclusions, weld/HAZ differences), leadingto higher data dispersion-an inherent feature of welded joint impact testing. Nevertheless, the meanvalues clearly define the transition trend, and data points closely follow the Boltzmann fitting curve(Figure 5), confirming result reliability.

Figure 5
Impact energy value fitting function.

It is noteworthy that the effect of water quenching on impact toughness observed in this study contrasts with its effect on tensile strength reported in literature. Shi et al.²⁶ found that water quenching significantly enhanced the tensile strength of Q235 steel after fire exposure. However, in the present study, water quenching generally reduced the impact energy of Q420C welded joints (except at 400°C). This discrepancy highlights the fundamental difference between static strength and dynamic toughness: while rapid cooling may induce hardening phases that increase tensile strength, it simultaneously introduces quenching stresses and brittle martensitic microstructures in the weld zone, which are detrimental to impact toughness, particularly at sub-zero temperatures.

3.3. Ductile-to-brittle transition temperature

Typically, as temperature decreases, a material's impact toughness declines, with a sharp reduction observed within a specific range. This phenomenon is known as “cold brittleness.” The corresponding temperature range is referred to as the ductile to brittle transition temperature (DBTT). This parameter is essential for assessing a material's performance in low-temperature environments and holds significant value in material design and safety evaluations.

The typical impact energy-temperature curve comprises three regions: an upper plateau, a transition zone, and a lower plateau. The ductile-to-brittle transition temperature is defined as the temperature at which the impact energy reaches 50% of the difference between the upper and lower plateaus. Lower DBTT indicates better low-temperature toughness and reduced susceptibility to brittle fracture. The impact energy exhibited an S-shaped trend with decreasing temperature: initially gradual, then sharply declining, and finally stabilizing. To quantify this behavior, the Boltzmann function was employed for regression analysis, as expressed in the following equation:

y = \frac{A_{1} - A_{2}}{1 + \exp (\frac{x - x_{0}}{d x})} + A_{2}

(1)

This model provides key parameters, including lower plateau energy (A₁), upper plateau energy (A₂), transition temperature (X₀), and transition range (Δ_x), effectively describing the material's transition behavior. The fitting results demonstrated a high degree of accuracy, with a correlation coefficient exceeding 0.99. The data are presented in Table 5. The corresponding regression curve is illustrated in Figure 5¹⁹,²⁷.

Thumbnail

Table 5
Boltzmann fitting function parameters.

While the Boltzmann function provided a good fit with a correlation coefficient (R²) exceeding 0.99, it is important to note that the accuracy of the fitted DBTT (X₀) is inherently constrained by the discrete sampling interval (20°C). Consequently, the reported transition temperatures carry a potential uncertainty of approximately 10°C, corresponding to the maximum distance between an adjacent data pair (e.g. between -20°C and -40°C) where the transition typically occurs. This limitation in sampling resolution is a common characteristic of Charpy impact testing, as it is impractical to test at infinitely fine temperature intervals. However, the high R² values and the consistent S-shaped trends observed across all groups confirm that the five-point sampling strategy is sufficient to capture the general transition behavior and relative comparisons between different heat treatments.

A comparison between the weld zone and HAZ in butt welds revealed that the weld zone exhibited a significantly higher DBTT, indicating greater vulnerability to brittle fracture at relatively higher temperatures. Following the heat treatment at 600°C, the DBTT values of the weld zone were −26.0 (air cooling) and −24.4°C (water cooling), and the respective values for the HAZ under the same conditions were −41.3 and −40.1°C. Additionally, the weld zone displayed lower upper and lower plateau energies and a narrower transition range, confirming its inferior low-temperature toughness relative to the HAZ.
In cruciform butt-welded joints, the HAZ continued to exhibit superior low-temperature toughness compared with the weld zone, with even greater sensitivity to temperature variation. The DBTT (X₀) of the weld zone remained consistently higher than that of the HAZ. For example, after 200°C heat treatment followed by natural cooling, the X₀ of the weld zone was −17.8°C, while the HAZ registered −30.2°C. Furthermore, the weld zone in cruciform butt-welded specimens showed a higher transition temperature than that of the butt welds, which was −22.2°C under the same conditions. Platform energy values further confirmed that the weld zone in cruciform butt welds had the lowest toughness. Specifically, the lower and upper plateau energies of the weld zone were 41.6 and 139.1 J, respectively, which were lower than those of the HAZ (81.4 and 185.2 J) and the butt weld zone (48.9 and 155.7 J). These results identify the weld zone of cruciform butt-welded joints as the most vulnerable region in terms of low-temperature impact resistance.
In both joint types, the HAZ demonstrated better toughness than the weld zone, with the weld zone of cruciform butt-welded joints being most vulnerable to low temperatures. Tempering temperature had a pronounced influence on the DBTT. At 400°C, embrittlement was most severe due to the precipitation of brittle phases, whereas tempering at 600°C lowered the transition temperature through refined microstructural evolution. However, raising the temperature to 800°C resulted in grain coarsening or the formation of hard, brittle phases, significantly reducing toughness. Regarding cooling methods, water quenching mitigated embrittlement at 400°C, while air cooling proved more effective in enhancing toughness at 600°C tempering²⁸.

The non-linear evolution of toughness, characterized by embrittlement at 400°C and recovery at 600°C, provides a specific insight into welded joints that extends the general understanding of post-fire steel. Tan et al.⁵ noted that rapid cooling reduces toughness due to internal stress and microstructural heterogeneity. Our results confirm this at 600°C and 800°C, where air cooling proved more beneficial than water quenching. However, an exception was observed at 400°C, where water quenching mitigated embrittlement. This suggests that at 400°C, the dominant damage mechanism is likely the precipitation of brittle phases (temper embrittlement), and the rapid cooling of water quenching restricts the kinetics of this precipitation, whereas slow air cooling provides sufficient time for detrimental phase aggregation. This distinct behavior at 400°C underscores the necessity of evaluating welded joints across a comprehensive temperature range rather than assuming a monotonic degradation.

3.4. Fracture analysis

To examine the fracture behavior of Q420C steel, both macro- and micro-scale analyses were performed. At the macroscopic level, fracture surfaces were evaluated based on shear lip width, fiber zone ratio, and radiating zone features to quantify ductility variations under different processing conditions. Two representative sample sets were selected for detailed SEM analysis: one comprising four regions (butt weld, butt HAZ, cruciform butt weld, and cruciform butt HAZ) after 200°C tempering with natural cooling; the other including butt weld specimens water-cooled after various tempering temperatures (200–800°C). SEM imaging was used to assess features such as cleavage steps, ductile fracture pits, and secondary cracks, providing insight into fracture mechanisms under differing thermal treatments.

The macroscopic fracture morphology of Q420C steel impact specimens from the weld and HAZs displayed systematic variation with temperature (Figue 6). As the treatment temperature increased, the fibrous zone proportion decreased, while the radiated zone expanded. At 200–400°C, fractures were predominantly fibrous, characterized by a rough, dark gray surface and continuous fibrous distribution in the HAZ. At 600–800°C, the radiated zone dominated, with visible cleavage steps and reduced shear lips, particularly in the weld zone.

Figure 6
Macro appearance of specimen fracture.

Cooling method significantly influenced fracture features: water-cooled specimens showed sharp radiated zone boundaries and narrow shear lips, whereas naturally cooled specimens retained more fibrous zones and broader shear lips with less distinct radiated zones. Decreasing test temperature from 20°C to −60°C further shifted the morphology toward radiated fracture. Below −40°C, the fracture surface was almost entirely radiated, with the weld zone showing this transition as early as −20°C, with fibrous features persisting longer in the HAZ. Overall, elevated temperatures promoted radiated zone formation, natural cooling slowed fibrous zone recession, and the weld zone displayed heightened sensitivity to low temperatures.

Figue 7 presents SEM fracture morphologies of Q420C steel specimens from the HAZ and weld metal (WM) of butt and cross weld after 200°C heat treatment and air cooling, within the DBTT range. At −20°C, the HAZ in butt welds showed deep, uniformly distributed dimples with minimal cleavage features. As the temperature dropped to −40°C, the dimples diminished in size and number, while cleavage planes increased to 30–40%. At −60°C, cleavage features dominated, with clear river patterns.

Figure 7
SEM images taken from different areas.

In the WM of butt welds, dimple degradation was more severe, showing shallower pits and a higher proportion of continuous cleavage planes. In cruciform butt welds, WM retained limited shallow dimples at 0°C. By −20°C, cleavage planes expanded with evident tear ridges, and at −40°C, the surface was almost entirely cleaved, with more cleavage steps than in the butt welds. Cruciform butt-welded WM specimens exhibited greater brittleness at equivalent temperatures due to higher residual stress.

Overall, the HAZ consistently demonstrated better microstructural stability and dimple retention at low temperatures than the WM, owing to its greater distance from the fusion line and reduced thermal stress.

Figue 8 displays the SEM fracture morphologies of the butt weld zone after various high-temperature treatments (200–800°C) followed by water cooling, under impact testing at −20, −40, and −60°C. Among all groups, the 600°C treatment group showed the best ductile behavior, with dense dimples and minimal cleavage features across all temperatures. The 200°C treatment group showed moderate ductility, with dimples decreasing and cleavage steps increasing at temperature drops. The 400°C treatment group exhibited marked brittleness, characterized by numerous cleavage steps and sparse, shallow dimples. The 800°C treatment group showed fully brittle fracture surfaces dominated by continuous cleavage steps and river patterns.

Figure 8
SEM images after exposure to different temperatures.

Across all treatment groups, a consistent embrittlement trend was observed with decreasing test temperature, reflected in reduced dimple density and increased cleavage surface area. The 600°C treatment group maintained the most stable ductility under low-temperature conditions, while the 400°C and 800°C treatment groups demonstrated severe embrittlement.

The three-stage variation in impact toughness-decrease at 400°C, improvement at 600°C, and deterioration at 800°C-is fundamentally governed by microstructural evolution. This is corroborated by Hua et al.⁶, who observed that cementite precipitation in high-strength steel within the 300-600°C range significantly alters mechanical properties. At 400°C, the precipitation and aggregation of brittle carbide phases (such as cementite) along grain boundaries act as stress concentrators, promoting cleavage fracture; this directly corresponds to the sparse dimples and dense cleavage steps observed in our SEM images. As the temperature rises to 600°C, the microstructure undergoes recovery, and the precipitated carbides tend to spheroidize and disperse more uniformly. This refined microstructural evolution relieves internal stress and eliminates the continuous brittle network, thereby enhancing toughness and explaining the dense dimples and minimal cleavage features observed. However, at 800°C, severe grain coarsening occurs in the weld and HAZ, and the subsequent cooling cycle promotes the formation of hard, brittle phases (such as coarse martensite or bainite). This microstructural degradation leads to a complete loss of ductile features, resulting in the fully brittle, continuous cleavage fracture surfaces observed in the SEM. Thus, the nonlinear fracture behavior is intrinsically linked to these temperature-dependent microstructural transformations.

4. Conclusions

The investigation into the impact toughness of steel welded joints following high-temperature cooling yields the following key conclusions:

Butt- and cruciform butt-welded specimens exhibit similar surface changes after high-temperature treatment, with coloration shifting from silver-white to dark black as temperature increases. At 200°C, no visible change occurs; at 400°C, a blue oxide film forms with slight weld discoloration; at 600°C, the weld contour becomes prominent, and surface carbides precipitate and detach upon water cooling; and at 800°C, a dark black carbide layer forms and detaches easily. Natural cooling results in smoother surfaces, while water cooling, particularly 600°C, intensifies surface peeling due to carbide formation.
The impact energy across all four regions follows a nonlinear trend: it decreases at 400°C, peaks at 600°C, and declines sharply at 800°C. The HAZ consistently exhibits higher toughness than the weld zone, with the weld zone of the cruciform butt weld showing the lowest energy. In terms of DBTT, the weld zone is more temperature-sensitive than the HAZ. Transition temperatures are highest at 400°C and lowest at 600°C, and they rise again at 800°C. Within the tested temperature range, the effect of cooling methods on toughness is temperature-dependent: water quenching appears to mitigate embrittlement at 400°C by restricting brittle phase precipitation, whereas air cooling is more beneficial for toughness recovery at 600°C by relieving internal stress.
With increasing treatment temperature, fracture surfaces show a reduced fiber zone and an expanded radiated zone: the fiber zone dominates at 200–400°C, while the radiated zone becomes predominant at 600–800°C, particularly in the weld zone. Water-cooled specimens display sharper radiated zone boundaries and a narrow shear lip, whereas naturally cooled specimens retain more of the fiber zone. At subzero ambient temperatures, the weld zone transitions to a radiated zone at −20°C, while the HAZ retains fiber characteristics down to −40°C. SEM analysis shows ductile dimples in the 200–600°C range but sparse dimples and dense cleavage steps at 400°C; the 800°C treatment group exhibits fully brittle fracture with continuous cleavage. Ductile features deteriorate more rapidly in the weld zone, especially in cruciform butt welds, which display higher cleavage step density. At −60°C, the HAZ retains some ductile features, unlike the weld zone, which shows complete cleavage.

5. Acknowledgments

This research work was supported by the Basic Research Operating Funds for Universities Directly Administered by the Autonomous Region (JY20240006); Inner Mongolia Autonomous Region Natural Science Foundation Project (2023LHMS05041); The Doctoral Research Startup Funding of Inner Mongolia University of Technology (Grant No. DC2300001267).

6. Data Availability

Raw data will be made available upon request.

7. References

¹ Sheng Y, Xu X, Rozelle S. Market structure, resource allocation, and industry productivity growth: firm-level evidence from China’s steel industry. China Econ Rev. 2024;83:102102. https://doi.org/10.1016/j.chieco.2023.102102
» https://doi.org/10.1016/j.chieco.2023.102102
² Ghimire A, Wald F, Vild M, Kabeláč J. Numerical design calculation of the high-strength steel welds. Eng Struct. 2024;300:117201. https://doi.org/10.1016/j.engstruct.2023.117201
» https://doi.org/10.1016/j.engstruct.2023.117201
³ Bjorhovde R. Development and use of high performance steel. J Construct Steel Res. 2004;60(3-5):393-400. https://doi.org/10.1016/S0143-974X(03)00118-4
» https://doi.org/10.1016/S0143-974X(03)00118-4
⁴ Javidan F, Heidarpour A, Zhao X-L, Minkkinen J. Application of high strength and ultra-high strength steel tubes in long hybrid compressive members: experimental and numerical investigation. Thin Walled Struct. 2016;102:273-85. https://doi.org/10.1016/j.tws.2016.02.002
» https://doi.org/10.1016/j.tws.2016.02.002
⁵ Tan J, Wang Z, Feng J, Li F, Li H, Jing J. Mechanical properties of Q235 and 20# steel after CO₂ fire extinguishing treatment. J Construct Steel Res. 2024;220:108815. https://doi.org/10.1016/j.jcsr.2024.108815
» https://doi.org/10.1016/j.jcsr.2024.108815
⁶ Hua J, Ren Z, Wang F, Wu J, Feng D, Xue X. Residual mechanical behavior of Q1100 ultra high strength steel after elevated temperatures. J Construct Steel Res. 2026;236:110011. https://doi.org/10.1016/j.jcsr.2025.110011
» https://doi.org/10.1016/j.jcsr.2025.110011
⁷ Xue X, Wang F, Hua J, Wang N, Yan S. Cooling methods impact on post-fire residual bond properties of bimetallic steel bars in seawater concrete. Fire Saf J. 2026;159:104569. https://doi.org/10.1016/j.firesaf.2025.104569
» https://doi.org/10.1016/j.firesaf.2025.104569
⁸ Wang F, Ren Z, Feng D, Xue X, Wu J, Li S, et al. Mechanical properties deterioration of ultra-high strength steel after elevated temperature exposure. Structures. 2026;84:111026. https://doi.org/10.1016/j.istruc.2025.111026
» https://doi.org/10.1016/j.istruc.2025.111026
⁹ Qiang X, Bijlaard FSK, Kolstein H. Post-fire performance of very high strength steel S960. J Construct Steel Res. 2013;80:235-42. https://doi.org/10.1016/j.jcsr.2012.09.002
» https://doi.org/10.1016/j.jcsr.2012.09.002
¹⁰ Qiang X, Bijlaard FSK, Kolstein H. Post-fire mechanical properties of high stre-ngth structural steels S460 and S690. Eng Struct. 2012;35:1-10. https://doi.org/10.1016/j.engstruct.2011.11.005
» https://doi.org/10.1016/j.engstruct.2011.11.005
¹¹ Wang Y-B, Li G-Q, Cui W, Chen S-W, Sun F-F. Experimental investigation and modeling of cyclic behavior of high strength steel. J Construct Steel Res. 2015;104:37-48. https://doi.org/10.1016/j.jcsr.2014.09.009
» https://doi.org/10.1016/j.jcsr.2014.09.009
¹² Hu F, Shi G, Shi Y. Constitutive model for full-range elasto-plastic behavior of structural steels with yield plateau: calibration and validation. Eng Struct. 2016;118:210-27. https://doi.org/10.1016/j.engstruct.2016.03.060
» https://doi.org/10.1016/j.engstruct.2016.03.060
¹³ Jiao H, Mashiri F, Zhao X-L. Fatigue behavior of very high strength (VHS) circular steel tube to plate T-joints under in-plane bending. Thin Walled Struct. 2013;68:106-12. https://doi.org/10.1016/j.tws.2013.03.006
» https://doi.org/10.1016/j.tws.2013.03.006
¹⁴ Heinilä S, Marquis GB, Björk T. Observations on fatigue crack paths in the corners of cold-formed high-strength steel tubes. Eng Fract Mech. 2008;75(3-4):833-44. https://doi.org/10.1016/j.engfracmech.2007.01.010
» https://doi.org/10.1016/j.engfracmech.2007.01.010
¹⁵ Chiew SP, Zhao MS, Lee CK. Fatigue performance of high strength steel built up box T-joints. J Construct Steel Res. 2015;106:296-310. https://doi.org/10.1016/j.jcsr.2014.12.018
» https://doi.org/10.1016/j.jcsr.2014.12.018
¹⁶ Amraei M, Jiao H, Zhao X-L, Tong L-W. Fatigue testing of butt-welded high strength square hollow sections strengthened with CFRP. Thin Walled Struct. 2017;120:260-8. https://doi.org/10.1016/j.tws.2017.09.004
» https://doi.org/10.1016/j.tws.2017.09.004
¹⁷ Lee JH, Park SH, Kwon HS, Kim GS, Lee CS. Laser, tungsten inert gas, and metal active gas welding of DP780 steel: comparison of hardness, tensile properties and fatigue resistance. Mater Des. 2014;64:559-65. https://doi.org/10.1016/j.matdes.2014.07.065
» https://doi.org/10.1016/j.matdes.2014.07.065
¹⁸ Lee C-H, Shin H-S, Park K-T, Chang K-H. Impact fracture energy of structural steel welds constructed at low ambient temperatures. Constr Build Mater. 2014;50:394-400. https://doi.org/10.1016/j.conbuildmat.2013.09.043
» https://doi.org/10.1016/j.conbuildmat.2013.09.043
¹⁹ Liu H, Zhou Y, Wang F, Lu Y, Chen Z. Experimental study on impact toughness of structural steel and its butt-welded joint at low temperature and corrosion. J Construct Steel Res. 2024;212:108298. https://doi.org/10.1016/j.jcsr.2023.108298
» https://doi.org/10.1016/j.jcsr.2023.108298
²⁰ Tong L, Niu L, Jing S, Ai L, Zhao X-L. Low temperature impact toughness of high strength structural steel. Thin Walled Struct. 2018;132:410-20. https://doi.org/10.1016/j.tws.2018.09.009
» https://doi.org/10.1016/j.tws.2018.09.009
²¹ Li Z, Zhang D, Wu H, Huang F, Hong W, Zang X. Fatigue properties of welded Q420 high strength steel at room and low temperatures. Constr Build Mater. 2018;189:955-66. https://doi.org/10.1016/j.conbuildmat.2018.07.231
» https://doi.org/10.1016/j.conbuildmat.2018.07.231
²² Zhang D, Li Z, Wu H, Huang F. Experimental study on fatigue behavior of Q420 high-strength steel at low temperatures. J Construct Steel Res. 2018;145:116-27. https://doi.org/10.1016/j.jcsr.2018.02.008
» https://doi.org/10.1016/j.jcsr.2018.02.008
²³ Jia L-J, Ikai T, Shinohara K, Ge H. Ductile crack initiation and propagation of structural steels under cyclic combined shear and normal stress loading. Constr Build Mater. 2016;112:69-83. https://doi.org/10.1016/j.conbuildmat.2016.02.171
» https://doi.org/10.1016/j.conbuildmat.2016.02.171
²⁴ Li X, Song Y, Ding Z, Bao S, Gao Z. A modified correlation between KJIC and Charpy V-notch impact energy of Chinese SA508-III steel at the upper shelf. J Nucl Mater. 2018;505:22-9. https://doi.org/10.1016/j.jnucmat.2018.03.056
» https://doi.org/10.1016/j.jnucmat.2018.03.056
²⁵ Száraz Z, Hähner P, Stráská J, Ripplinger S. Effect of phase separation on tensile and Charpy impact properties of MA956 ODS steel. Mater Sci Eng A. 2017;700:425-37. https://doi.org/10.1016/j.msea.2017.06.010
» https://doi.org/10.1016/j.msea.2017.06.010
²⁶ Shi G, Wang S, Chen X, Rong C. Post-fire mechanical properties of base metal and welds of Q235 steel. J Construct Steel Res. 2021;183:106767. https://doi.org/10.1016/j.jcsr.2021.106767
» https://doi.org/10.1016/j.jcsr.2021.106767
²⁷ García-Junceda A, Hernández-Mayoral M, Serrano M. Influence of the microstructure on the tensile and impact properties of a 14Cr ODS steel bar. Mater Sci Eng A. 2012;556:696-703. https://doi.org/10.1016/j.msea.2012.07.051
» https://doi.org/10.1016/j.msea.2012.07.051
²⁸ Ye K, Ozaki F. Impact fracture energy of steel welded connections in fire and post-fire. J Construct Steel Res. 2020;170:106120. https://doi.org/10.1016/j.jcsr.2020.106120
» https://doi.org/10.1016/j.jcsr.2020.106120

*

e-mail: yangliguo@imut.edu.cn

Associate Editor:

Ana Sofia de Oliveira.

Editor-in-Chief:

Luiz Antonio Pessan.

Publication Dates

Publication in this collection
25 Sept 2026
Date of issue
2026

History

Received
02 Apr 2026
Reviewed
06 June 2026
Accepted
13 Aug 2026

This is an Open Access article distributed under the terms of the Creative Commons Attribution license, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

[1] ¹ Sheng Y, Xu X, Rozelle S. Market structure, resource allocation, and industry productivity growth: firm-level evidence from China’s steel industry. China Econ Rev. 2024;83:102102. https://doi.org/10.1016/j.chieco.2023.102102
» https://doi.org/10.1016/j.chieco.2023.102102

[2] ² Ghimire A, Wald F, Vild M, Kabeláč J. Numerical design calculation of the high-strength steel welds. Eng Struct. 2024;300:117201. https://doi.org/10.1016/j.engstruct.2023.117201
» https://doi.org/10.1016/j.engstruct.2023.117201

[3] ³ Bjorhovde R. Development and use of high performance steel. J Construct Steel Res. 2004;60(3-5):393-400. https://doi.org/10.1016/S0143-974X(03)00118-4
» https://doi.org/10.1016/S0143-974X(03)00118-4

[4] ⁴ Javidan F, Heidarpour A, Zhao X-L, Minkkinen J. Application of high strength and ultra-high strength steel tubes in long hybrid compressive members: experimental and numerical investigation. Thin Walled Struct. 2016;102:273-85. https://doi.org/10.1016/j.tws.2016.02.002
» https://doi.org/10.1016/j.tws.2016.02.002

[5] ⁵ Tan J, Wang Z, Feng J, Li F, Li H, Jing J. Mechanical properties of Q235 and 20# steel after CO₂ fire extinguishing treatment. J Construct Steel Res. 2024;220:108815. https://doi.org/10.1016/j.jcsr.2024.108815
» https://doi.org/10.1016/j.jcsr.2024.108815

[6] ⁶ Hua J, Ren Z, Wang F, Wu J, Feng D, Xue X. Residual mechanical behavior of Q1100 ultra high strength steel after elevated temperatures. J Construct Steel Res. 2026;236:110011. https://doi.org/10.1016/j.jcsr.2025.110011
» https://doi.org/10.1016/j.jcsr.2025.110011

[7] ⁷ Xue X, Wang F, Hua J, Wang N, Yan S. Cooling methods impact on post-fire residual bond properties of bimetallic steel bars in seawater concrete. Fire Saf J. 2026;159:104569. https://doi.org/10.1016/j.firesaf.2025.104569
» https://doi.org/10.1016/j.firesaf.2025.104569

[8] ⁸ Wang F, Ren Z, Feng D, Xue X, Wu J, Li S, et al. Mechanical properties deterioration of ultra-high strength steel after elevated temperature exposure. Structures. 2026;84:111026. https://doi.org/10.1016/j.istruc.2025.111026
» https://doi.org/10.1016/j.istruc.2025.111026

[9] ⁹ Qiang X, Bijlaard FSK, Kolstein H. Post-fire performance of very high strength steel S960. J Construct Steel Res. 2013;80:235-42. https://doi.org/10.1016/j.jcsr.2012.09.002
» https://doi.org/10.1016/j.jcsr.2012.09.002

[10] ¹⁰ Qiang X, Bijlaard FSK, Kolstein H. Post-fire mechanical properties of high stre-ngth structural steels S460 and S690. Eng Struct. 2012;35:1-10. https://doi.org/10.1016/j.engstruct.2011.11.005
» https://doi.org/10.1016/j.engstruct.2011.11.005

[11] ¹¹ Wang Y-B, Li G-Q, Cui W, Chen S-W, Sun F-F. Experimental investigation and modeling of cyclic behavior of high strength steel. J Construct Steel Res. 2015;104:37-48. https://doi.org/10.1016/j.jcsr.2014.09.009
» https://doi.org/10.1016/j.jcsr.2014.09.009

[12] ¹² Hu F, Shi G, Shi Y. Constitutive model for full-range elasto-plastic behavior of structural steels with yield plateau: calibration and validation. Eng Struct. 2016;118:210-27. https://doi.org/10.1016/j.engstruct.2016.03.060
» https://doi.org/10.1016/j.engstruct.2016.03.060

[13] ¹³ Jiao H, Mashiri F, Zhao X-L. Fatigue behavior of very high strength (VHS) circular steel tube to plate T-joints under in-plane bending. Thin Walled Struct. 2013;68:106-12. https://doi.org/10.1016/j.tws.2013.03.006
» https://doi.org/10.1016/j.tws.2013.03.006

[14] ¹⁴ Heinilä S, Marquis GB, Björk T. Observations on fatigue crack paths in the corners of cold-formed high-strength steel tubes. Eng Fract Mech. 2008;75(3-4):833-44. https://doi.org/10.1016/j.engfracmech.2007.01.010
» https://doi.org/10.1016/j.engfracmech.2007.01.010

[15] ¹⁵ Chiew SP, Zhao MS, Lee CK. Fatigue performance of high strength steel built up box T-joints. J Construct Steel Res. 2015;106:296-310. https://doi.org/10.1016/j.jcsr.2014.12.018
» https://doi.org/10.1016/j.jcsr.2014.12.018

[16] ¹⁶ Amraei M, Jiao H, Zhao X-L, Tong L-W. Fatigue testing of butt-welded high strength square hollow sections strengthened with CFRP. Thin Walled Struct. 2017;120:260-8. https://doi.org/10.1016/j.tws.2017.09.004
» https://doi.org/10.1016/j.tws.2017.09.004

[17] ¹⁷ Lee JH, Park SH, Kwon HS, Kim GS, Lee CS. Laser, tungsten inert gas, and metal active gas welding of DP780 steel: comparison of hardness, tensile properties and fatigue resistance. Mater Des. 2014;64:559-65. https://doi.org/10.1016/j.matdes.2014.07.065
» https://doi.org/10.1016/j.matdes.2014.07.065

[18] ¹⁸ Lee C-H, Shin H-S, Park K-T, Chang K-H. Impact fracture energy of structural steel welds constructed at low ambient temperatures. Constr Build Mater. 2014;50:394-400. https://doi.org/10.1016/j.conbuildmat.2013.09.043
» https://doi.org/10.1016/j.conbuildmat.2013.09.043

[19] ¹⁹ Liu H, Zhou Y, Wang F, Lu Y, Chen Z. Experimental study on impact toughness of structural steel and its butt-welded joint at low temperature and corrosion. J Construct Steel Res. 2024;212:108298. https://doi.org/10.1016/j.jcsr.2023.108298
» https://doi.org/10.1016/j.jcsr.2023.108298

[20] ²⁰ Tong L, Niu L, Jing S, Ai L, Zhao X-L. Low temperature impact toughness of high strength structural steel. Thin Walled Struct. 2018;132:410-20. https://doi.org/10.1016/j.tws.2018.09.009
» https://doi.org/10.1016/j.tws.2018.09.009

[21] ²¹ Li Z, Zhang D, Wu H, Huang F, Hong W, Zang X. Fatigue properties of welded Q420 high strength steel at room and low temperatures. Constr Build Mater. 2018;189:955-66. https://doi.org/10.1016/j.conbuildmat.2018.07.231
» https://doi.org/10.1016/j.conbuildmat.2018.07.231

[22] ²² Zhang D, Li Z, Wu H, Huang F. Experimental study on fatigue behavior of Q420 high-strength steel at low temperatures. J Construct Steel Res. 2018;145:116-27. https://doi.org/10.1016/j.jcsr.2018.02.008
» https://doi.org/10.1016/j.jcsr.2018.02.008

[23] ²³ Jia L-J, Ikai T, Shinohara K, Ge H. Ductile crack initiation and propagation of structural steels under cyclic combined shear and normal stress loading. Constr Build Mater. 2016;112:69-83. https://doi.org/10.1016/j.conbuildmat.2016.02.171
» https://doi.org/10.1016/j.conbuildmat.2016.02.171

[24] ²⁴ Li X, Song Y, Ding Z, Bao S, Gao Z. A modified correlation between KJIC and Charpy V-notch impact energy of Chinese SA508-III steel at the upper shelf. J Nucl Mater. 2018;505:22-9. https://doi.org/10.1016/j.jnucmat.2018.03.056
» https://doi.org/10.1016/j.jnucmat.2018.03.056

[25] ²⁵ Száraz Z, Hähner P, Stráská J, Ripplinger S. Effect of phase separation on tensile and Charpy impact properties of MA956 ODS steel. Mater Sci Eng A. 2017;700:425-37. https://doi.org/10.1016/j.msea.2017.06.010
» https://doi.org/10.1016/j.msea.2017.06.010

[26] ²⁶ Shi G, Wang S, Chen X, Rong C. Post-fire mechanical properties of base metal and welds of Q235 steel. J Construct Steel Res. 2021;183:106767. https://doi.org/10.1016/j.jcsr.2021.106767
» https://doi.org/10.1016/j.jcsr.2021.106767

[27] ²⁷ García-Junceda A, Hernández-Mayoral M, Serrano M. Influence of the microstructure on the tensile and impact properties of a 14Cr ODS steel bar. Mater Sci Eng A. 2012;556:696-703. https://doi.org/10.1016/j.msea.2012.07.051
» https://doi.org/10.1016/j.msea.2012.07.051

[28] ²⁸ Ye K, Ozaki F. Impact fracture energy of steel welded connections in fire and post-fire. J Construct Steel Res. 2020;170:106120. https://doi.org/10.1016/j.jcsr.2020.106120
» https://doi.org/10.1016/j.jcsr.2020.106120

Indicators	Chemical composition (%)
Indicators	C	Si	Mn	Ni	Cr	Cu	Mo	B	V	Ceq
Actual value	0.140	0.042	0.800	0.007	0.015	0.017	0	—	0	0.278
Factory standard	0.070	0.190	1.080	0.010	0.020	0.030	0.004	—	0	—
Product standard	0.200	0.500	1.700	0.800	0.300	0.300	0.200	—	0.200	0.450

Welding method	Diameter (mm)	Current (A)	Voltage (V)	Speed (cm·min⁻¹)
Manual arc welding	3.2	110	40	20–25

Serial number	Sample group number	Impact location	No. of specimens
1	BW-WZ	Weld metal zone of butt joint	120
2	BW-HAZ	HAZ of butt joint	120
3	CBW-WZ	Weld metal zone of cruciform butt-welded joint	120
4	CBW-HAZ	HAZ of cruciform butt-welded joint	120

Impact region	Temperature (°C)	Impact energy value (J)
Impact region	Temperature (°C)	A200	A400	A600	A800	W200	W400	W600	W800
BW-WZ	20.0	153.7 (8.5)	107.2 (9.8)	190.2 (7.0)	93.8 (12.4)	138.0 (6.2)	123.8 (6.3)	168.3 (7.3)	76.4 (11.5)
	0.0	140.6 (8.0)	93.7 (9.5)	178.3 (11.4)	76.2 (6.7)	120.3 (10.3)	105.6 (12.2)	157.8 (10.9)	60.9 (12.3)
	−20.0	108.4 (12.4)	60.1 (14.0)	154.2 (8.5)	52.4 (9.4)	90.7 (11.2)	76.8 (6.8)	127.6 (13.7)	42.8 (5.6)
	−40.0	66.8 (9.8)	39.5 (11.8)	95.1 (13.0)	32.3 (12.7)	55.6 (13.5)	48.2 (9.0)	82.4 (10.7)	23.8 (9.1)
	−60.0	53.4 (5.1)	26.4 (8.6)	82.8 (6.5)	22.0 (14.7)	39.0 (12.6)	32.1 (14.5)	67.1 (13.0)	16.0 (6.5)
BW-HAZ	20.0	198.7 (9.7)	155.2 (7.3)	239.3 (9.6)	131.6 (5.9)	188.0 (5.2)	171.3 (7.7)	217.2 (12.0)	126.1 (12.3)
	0.0	188.1 (8.3)	138.8 (5.1)	231.0 (13.4)	114.2 (12.4)	178.0 (6.6)	159.3 (10.5)	208.9 (14.8)	109.9 (10.2)
	−20.0	138.7 (13.6)	118.0 (5.7)	217.0 (5.9)	93.4 (13.7)	158.0 (8.3)	136.6 (13.4)	193.7 (5.2)	76.7 (11.5)
	−40.0	138.7 (12.0)	77.3 (8.7)	184.2 (10.2)	65.0 (15.4)	110.0 (14.2)	97.0 (12.1)	161.9 (11.5)	47.4 (5.7)
	−60.0	110.4 (9.4)	51.8 (5.9)	151.9 (9.7)	47.7 (6.7)	74.0 (5.7)	68.3 (10.6)	131.5 (12.0)	30.4 (5.8)
CBW-WZ	20.0	131.1 (7.1)	87.2 (6.3)	185.8 (6.8)	75.2 (9.1)	118.0 (5.3)	102.6 (9.4)	160.1 (11.5)	61.6 (8.8)
CBW-WZ	0.0	117.9 (7.2)	73.4 (8.9)	173.6 (12.1)	61.7 (10.2)	105.9 (11.7)	86.7 (12.6)	149.5 (12.3)	49.0 (9.2)
Impact region	Temperature (°C)	Impact energy value (J)
Impact region	Temperature (°C)	A200	A400	A600	A800	W200	W400	W600	W800
	−20.0	84.8 (9.1)	46.4 (7.5)	129.7 (13.7)	37.8 (6.7)	70.0 (13.5)	57.6 (8.2)	107.9 (5.6)	28.9 (14.1)
CBW-WZ	−40.0	61.6 (8.1)	32.8 (9.7)	88.0 (5.9)	25.7 (9.7)	50.2 (10.3)	39.0 (14.9)	75.2 (13.5)	18.7 (8.8)
CBW-WZ	−60.0	46.2 (5.3)	20.3 (12.4)	68.1 (6.7)	17.2 (10.2)	31.8 (12.1)	27.3 (7.7)	58.7 (6.3)	10.4 (9.2)
CBW-HAZ	20.0	188.1 (7.4)	114.4 (9.7)	246.3 (13.1)	94.1 (9.0)	177.3 (13.4)	155.9 (14.3)	226.1 (10.7)	85.1 (7.7)
	0.0	174.5 (8.5)	106.1 (6.4)	231.3 (6.7)	87.3 (12.9)	162.5 (6.1)	142.9 (12.8)	219.1 (12.3)	78.1 (11.9)
	−20.0	159.3 (14.3)	90.4 (13.7)	211.1 (7.8)	70.7 (10.2)	148.4 (8.7)	130.6 (13.8)	195.5 (5.6)	61.2 (13.5)
	−40.0	109.1 (11.8)	66.4 (5.3)	161.9 (13.5)	52.6 (8.4)	88.3 (9.5)	77.7 (13.7)	141.4 (9.1)	42.4 (9.0)
	−60.0	87.5 (12.2)	53.2 (8.7)	127.5 (6.7)	43.8 (6.1)	74.7 (10.4)	65.7 (14.4)	109.1 (6.5)	35.9 (13.5)

Impact region	Fitting parameters	Function parameters
Impact region	Fitting parameters	A200	A400	A600	A800	W200	W400	W600	W800
BW-WZ	A₁ (J)	48.9	24.1	80.4	15.2	29.3	21.4	62.9	9.3
	A₂ (J)	155.7	114.3	186.9	107.5	145.7	135.8	169.1	88.0
	X₀ (°C)	−22.2	−15.6	−26.0	−12.5	−21.1	−18.7	−24.4	−13.2
	Δx (°C)	11.6	13.8	8.0	18.7	15.9	18.1	10.8	19.2
BW-HAZ	A₁ (J)	84.4	31.6	123.2	28.5	57.0	48.9	104.0	22.6
	A₂ (J)	202.2	160.1	241.2	145.1	188.8	175.2	219.7	135.4
	X₀ (°C)	−37.6	−30.7	−41.3	−23.7	−35.0	−32.5	−40.1	−19.2
	Δx (°C)	17.6	17.1	16.4	22.0	12.2	16.0	16.8	16.0
CBW-WZ	A₁ (J)	41.6	16.5	63.4	15.7	26.8	23.4	56.3	7.8
	A₂ (J)	139.1	98.5	190.9	83.4	127.8	112.7	165.0	72.0
	X₀ (°C)	−17.8	−12.6	−21.5	−10.5	−17.7	−13.4	−19.7	−9.2
	Δx (°C)	15.0	17.2	12.3	15.0	15.8	15.9	11.9	17.4
CBW-HAZ	A₁ (J)	81.4	46.1	106.1	40.4	72.7	64.0	99.5	33.7
	A₂ (J)	185.2	116.8	247.8	96.7	171.6	151	226.6	86.8
	X₀ (°C)	−30.2	−27.0	−34.2	−22.2	−28.2	−28.2	−32.3	−20.6
	Δx (°C)	10.3	14.8	14.8	13.8	7.4	7.4	11.0	12.3