This study used 4340M steel, a low alloy steel used in aircraft landing gears and high-strength bolts for marine structures, including wind. Table 1 lists the chemical composition of 4340M steel.

Regarding specimen preparation, the base metal with no treatment (BM specimen) was prepared and subjected to two-stage tempering for three hours at 300 °C after two hours oil quenching at 850 °C (QT specimen). A shot peening treatment was performed after quenching and tempering heat treatment (QT heat treatment; SP specimen). The standard specifications for repairing aircraft landing gears were applied to the shot peening treatment (SAE Standards, 2018). After shot peening, the specimen was subjected to baking heat treatment for four hours at 246 °C (SP-B specimen). A hard Cr-plating treatment was applied after the baking heat treatment to maintain the wear resistance. Plating was performed at a current density of 35 A/dm2 and 56 °C for 20 hours. After plating, baking heat treatment was performed at 246 °C for four hours to prevent hydrogen embrittlement (CRP-B specimen). The plated thickness after Cr-plating was stripped over 11.3 hours (ST specimen). After stripping, Cr-plating was performed again after re-shot peening, followed by a baking heat treatment (RESP-CRP-B specimen). Table 2 lists the specimens used in the Charpy impact test.

The tensile test was conducted using the specimens prepared under the conditions listed in Table 2. In the test, the crosshead displacement speed was set to 2 mm/s in the air at room temperature using a 100 kN universal testing machine (UH-F100A, Shimadzu, Japan). The test was conducted three times for each specimen. In addition, the Charpy impact test was conducted in accordance with ASTM E23 (ASTM Standards, 2024) using a SIM-30 (Sunwoo, Korea) Charpy impact test machine with a 30 kg·m capacity. The Charpy impact test specimens were prepared with a U-notch to enable a shot peening treatment in the notch area. Figs. 1 and 2 show the geometry and dimensions of the specimens (ASTM Standards, 2024). The temperature of the steady temperature and humidity room for the Charpy impact test was controlled at 23 °C ± 5 °C to avoid negative impacts on the ductility of the test material. The Charpy impact test of this study was conducted at five temperatures (−100 °C, −50 °C, 0 °C, 50 °C, or 100 °C) to evaluate the absorbed energy and the percent of brittle fracture according to the temperature for each specimen. The test temperatures were maintained during the impact test by spraying was sprayed to the notched specimen with liquefied nitrogen through a liquefied nitrogen nozzle for more than 30 minutes in the case of low temperatures, and the specimen was installed in the fixing jig as soon as possible. Real-time monitoring was performed using a non-contact infrared thermometer to ensure that the temperature change of the specimen was within ±1 °C. A high-temperature furnace was used to reach the set temperature in the case of high temperatures. Real-time monitoring was also performed using a non-contact infrared thermometer to ensure that the temperature change in the specimen was less than ±1 °C. The impact test was conducted within five seconds as much as possible. After the Charpy impact test, the fracture surface was analyzed by optical microscopy (BX53, Olympus, Japan) and scanning electron microscopy (SEM; EM-30plus, Coxem, Korea).

The tensile test was conducted on the kinds of seven specimens used in this study (Table 2). Table 3 summarizes the results of three times tensile tests conducted on each specimen.

The Charpy impact test was conducted on the kinds of seven specimens at five temperatures, and the obtained absorbed energy and percent of brittle fracture were examined.

Fig. 3 presents the results of the BM specimen, which is the base metal of 4340M steel. For the pure base metal with no treatment, the absorbed energy increased, and the percent of brittle fracture decreased as the temperature increased. From −100 °C to −50 °C, the absorbed energy slightly increased while the percent of brittle fracture remained almost constant. The percent of brittle fracture decreased as the absorbed energy further increased, indicating an increase in the percent of ductile fracture. The absorbed energy, and the percent of brittle fracture crossed at approximately 30 °C.

Fig. 4 shows the results of the QT specimen subjected to the QT heat treatment. The absorbed energy increased gently. The percent of brittle fracture decreased as the temperature was increased to 0 °C, but it tended to increase slightly at higher temperatures. The absorbed energy and the percent of brittle fracture crossed at 0 °C. Over the range 0–50 °C, the percent of brittle fracture increased slightly while the absorbed energy remained almost constant. Over the 50–100 °C range, the absorbed energy increased slightly while the percent of brittle fracture remained almost constant. Despite the gentle increase in the absorbed energy, the percent of brittle fracture did not decrease, and it increased again after 0 °C. The slight increase in the absorbed energy suggests that the toughness was maintained, even though the QT heat treatment achieved high strength and hardness. Nevertheless, the percent of brittle fracture increased because of the increase in tensile stress due to the increase in strength caused by QT heat treatment, which is considered that this is because brittle fracture has become more likely to occur even at high temperatures. It is considered that this point requires reexamination. It is thought that because the tensile strength was greatly improved through QT heat treatment and high strength was achieved, there was no significant change in the value of the absorbed energy depending on the temperature and it showed a tendency to show an almost gentle slope. (Liu et al., 2022).

Fig. 5 shows the results of the SP specimen subjected to the shot peening treatment. The absorbed energy increased as the temperature rose. The percent of brittle fracture decreased as the temperature was increased to 0 °C and increased slightly over the 0–50 °C range and then slightly decreased again over the 50–100 °C range. The absorbed energy and the percent of brittle fracture did not cross. Shot peening significantly improved the strength and hardness of the surface, but the absorbed energy tended to increase as the temperature increased. Nevertheless, the percent of brittle fracture did not decrease despite the increase in the absorbed energy because high strength and high hardness were maintained by the modification effect of the shot peening surface.

Fig. 6 shows the results of the SP-B specimen subjected to baking heat treatment after shot peening. The baking heat treatment after shot peening reduces residual tensile stress. Most metals exhibit a decrease in yield strength when heated. When the yield strength of a material is reduced by heating, yielding or deformation occurs in a material that experiences a residual stress larger than the reduced yield strength. Hence, the residual stress in the heated material is not as high as the yield strength of the material. Therefore, a baking heat treatment is performed after shot peening to prevent the tensile residual stress that involves deformation. The yield strength was improved by a baking heat treatment after shot peening compared to the yield strength after shot peening, as listed in Table 3. In the case of the 4340M steel used in this study, the baking heat treatment was performed at 246 °C (Ahn et al., 2020) to prevent exceeding the temperature at which the maximum hardness is reached. The SP-B specimen showed a similar tendency to the SP specimen, including no crossing between the absorbed energy and the percent of brittle fracture.

Fig. 7 shows the results of the CRP-B specimen subjected to a baking heat treatment after Cr-plating on the SP specimen. Baking heat treatment was performed to prevent hydrogen embrittlement that occurs after Cr-plating. The absorbed energy was lowest at −100 °C, possibly because of the surface hardness increased by Cr-plating. This result was confirmed by the finding that the yield strength of the specimen was slightly higher than other specimens. The percent of brittle fracture at −100 °C was also slightly higher than that of the QT, SP, and SP-B specimens. Overall, the CRP-B specimen exhibited similar tendencies to the SP-B specimen regarding the absorbed energy and percent of brittle fracture, but the changes in the values at temperatures below 0 °C were somewhat larger. The absorbed energy and the percent of brittle fracture crossed at 100 °C, which had no significant impact on the results.

Fig. 8 shows the results of the ST specimen subjected to the stripping of Cr-plating. Cases where Cr-plating was stripped by degradation that may occur during the use of parts, such as wear, friction, and corrosion, were considered. Although the degree of stripping can vary (e.g., partial stripping), this study considered only the most severe cases of complete stripping from the specimens. The absorbed energy and the percent of brittle fracture at −100 °C were generally higher than other specimens, but they exhibited similar tendencies. While the absorbed energy exhibited the gentlest curve, the percent of brittle fracture tended to decrease. The specimen was under the condition of completely stripping Cr-plating, indicating a return to the shot-peened SP specimen before plating. The absorbed energy remained almost constant compared to the case of the shot-peened specimen. This result was attributed to the exposure of the shot-peened surface by stripping the Cr-plating. Nevertheless, the surface hardness by Cr-plating was reduced somewhat by stripping because the Cr-plating completely disappeared, and the strength decreased under the stress relief caused by the electrolysis power applied during stripping. In particular, despite the stripping of Cr-plating across the specimen, the absorbed energy remained relatively constant according to the temperature, and high impact energy values were observed, maintaining the high strength caused by shot peening and exhibiting no sharp reduction in toughness. This appears to be because the stripping of Cr-plating reduced the brittleness somewhat. The absorbed energy and the percent of brittle fracture did not cross.

Fig. 9 shows the results for the RESP-CRP-B specimen, which was subjected to baking heat treatment after Cr-plating after re-shot peening after a stripping. This specimen was prepared to examine whether the parts can be used continuously in the repair and maintenance process after their use. The overall tendencies of the absorbed energy and the percent of brittle fracture were similar to those of the stripped specimen. The absorbed energy at −100 °C was slightly lower than the ST specimen. This result was attributed to the brittleness caused by performing re-shot peening and Cr-plating. No significant reduction in the value of the absorbed energy was observed because of the final baking heat treatment. Stripping was performed once for this specimen, and Cr-plating was applied after re-shot peening. The exposed surface was returned to the initial shot-peened state by removing Cr-plating through stripping. The result is considered that affected by the changes of the dimples formed on the surface and compressive residual stress because shot peening was performed again for the surface. Cr-plating was applied in the same method as that shown in Figs. 7 and 9, but the impact energy was higher because of re-shot peening (Fig. 9). Therefore, future studies should examine the impacts of re-shot peening. When the results both the stripped specimen and the Cr-plated specimen after re-shot peening were compared to the result of the Cr-plated specimen of Fig. 7, however, the absorbed energy at a low temperature of −100 °C increased. This indicates an improvement in impact toughness at low temperatures. The absorbed energy and the percent of brittle fracture did not cross.

Fig. 10 shows the DBTT from the absorbed energy obtained from the Charpy impact test of surface-treated 4340M steel. The absorbed energy of the BM specimen with no surface treatment changed according to the temperature. The absorbed energy of the specimens subjected to a surface treatment (e.g., QT heat treatment, shot peening treatment, Cr-plating, stripping, and re-shot peening) remained nearly constant, and similar results were obtained with no significant change depending on the temperature. On the other hand, slight differences were observed in the absorbed energy value at a low temperature of −100 °C applied in this study. In particular, applying Cr-plating resulted in the lowest absorbed energy value at −100 °C because the surface hardness was increased by plating. Hence, the impact absorption was reduced at low temperatures because of brittleness. By contrast, the absorbed energy was similar when the applied temperature range exceeded −50 °C. The slight change in absorbed energy suggests that little plastic deformation occurred. The base 4340M low alloy steel is classed as high-strength steel. The change in impact energy observed from high alloy steel was relatively constant, showing a straight line because it was modified to high-strength steel with a tensile strength almost twice as high as the base metal through surface treatment, such as QT and shot peening treatment. This led to the surface-treated specimens showing high yield strengths, as shown in Table 3. Metals with high yield strength tend to absorb less impact energy (Dudko et al., 2017; Yang et al., 2017). This result is consistent with the brittle fracture before reaching the stress required for plastic deformation, even at low temperatures, because the yield stress is high (Tong et al., 2018).

The percent of brittle fracture did not decrease below 50%, resulting in no intersection point with the impact value corresponding to 50% of the maximum value of the absorbed energy. The absorbed energy barely changed as the low temperatures changed into high temperatures with respect to approximately −50 °C, showing that it was relatively unaffected by the temperature change. It remained almost constant as the temperature increased above −50 °C. Therefore, the absorbed energy can be estimated because the transition temperature is formed in the low-temperature range below −50 °C. The surface-treated specimens applied in this study, the transition temperature on the absorbed energy can be estimated is formed in the low-temperature range below −50 °C. Despite the high-strength and high-hardness surface modification, the absorbed energy was constant while the transition temperature by impact test was reduced, indicating that the impact toughness was also somewhat maintained at low temperatures. Hence, these results show that the specimens still exhibited high absorbed energy of 100 J or more except for Cr-plating, even though the energy was generally reduced at −100 °C. This helps respond to the fracture of the applied material because of the impacts at low temperatures by making it possible to maintain high toughness while improving the strength and hardness through surface treatment. In addition, the baking heat treatment performed to prevent hydrogen embrittlement reduced the transition temperature because the absorbed energy maintained a constant value with no sharp change.

The fracture appearance transition temperature (FATT) could not be obtained quantitatively because even the BM specimen exhibited approximately 54.2% as the percent of brittle fracture at 100 °C. DBTT in BM was approximately −23 °C. On the other hand, the DBTT was approximately −50 °C for the QT, SP, SP-B, ST, and RESP-CRP-B specimens, which were strengthened through heat treatment and surface treatment by shot peening. The DBTT was approximately − 75 °C for the CRP-B specimen. For these specimens, the absorbed energy showed a constant tendency with no significant change in the range −50–100 °C, and the absorbed energy decreased as the temperature decreased from −50 to −100 °C. This appears to be because the strength and hardness of the surface were increased by the surface treatment applied to the specimens, such as shot peening and Cr-plating, as indicated in the graphs showing the results of each specimen. The tensile strength was similar, even in the cases of stripping as well as re-Cr-plating after re-shot peening (Table 3), which is in good agreement with the results observed from materials with high strength and hardness (Chao et al., 2007).

In the case of the BM specimen, the absorbed energy was relatively constant over the 0–50 °C range, which decreased as the temperature decreased from 0 °C to −50 °C. The absorbed energy changed clearly because the strength of the BM specimen was lower than that of the surface-modified specimens. On the other hand, the change in the absorbed energy because of the high strength and hardness caused by the surface treatment was relatively constant regardless of the temperature within the test range of −50 °C to 100 °C applied in this study. Therefore, the impact resistance was similar regardless of the temperature when higher strength and hardness were maintained by applying surface treatment through shot peening to steels with high strength, hardness, and toughness, such as 4340M steel. Therefore, the surface modification effect can expand the applications and applicable temperature range of the material.

Fig. 11 shows the percent of brittle fracture according to the temperature. Higher tensile stress is required to reach the yield condition because the notch area under stress concentration is in the triaxial stress state. The increase in tensile stress due to the increase in strength caused by shot peening makes the occurrence of brittle fracture easier, even at higher temperatures. Therefore, even if the temperature rises above 0 °C, the percent of brittle fracture does not decrease rapidly but it also indicates to increase slightly. In the case of the BM specimen, the percent of brittle fracture decreased sharply as the temperature increased above 0 °C compared to other specimens, but it did not decrease below 50%. This was attributed to the high tensile strength of the origin material.

Table 4 shows the images of the fracture surfaces of the specimens obtained after the Charpy impact test. The stains observed from some specimens in the images were rust that occurred after the impact fracture, which did not affect the test results. The percent of brittle fracture was generally high for the specimens, including the BM specimen. For the BM specimen, the percent of ductile fracture tended to increase slightly. Nevertheless, the percent of brittle fracture exceeded 50%.

Table 5 shows SEM images of the fracture surfaces of the specimens treated with shot peening and Cr-plating among the fractured specimens after the Charpy impact test shown in Table 4: (a) images of the notch tip of the CRP-B specimen subjected to shot peening, Cr-plating, and baking heat treatment; (b) images of the broken end of the CRP-B specimen; (c) images of the notch tip of the RESP-CRP-B specimen subjected to re-shot peening, Cr-plating, and baking heat treatment after stripping; and (d) images of the break end of the RESP-CRP-B specimen. Overall, the fracture surfaces at each temperature were the cleavage fracture surfaces of the river pattern. As the temperature rose, ductile dimples appeared intermittently along with cleavage fracture surfaces. The percent of brittle fracture did not decrease below 50%, even at 100 °C. In addition, the absorbed energy from the low-temperature range was similar, while DBTT decreased to the low-temperature range. Hence, the durability against impact fracture can be improved by maintaining high toughness and strength by surface treatment.