
==== Front
Heliyon
Heliyon
Heliyon
2405-8440
Elsevier

S2405-8440(24)12664-3
10.1016/j.heliyon.2024.e36633
e36633
Research Article
Study of cracks in the last-stage rotor blade of a steam turbine and the corrosion fatigue properties of its materials
Gao Jiashun ab
Tang Zhuolin c
Guo Bicheng c
Xu Zhilong zhilong.xu@163.com
ac⁎
Liu Ming mingliu@fzu.edu.cn
d⁎⁎
Sun Wenbing e
Zhao Zhenye c
a School of Marine Engineering, Jimei University, Xiamen, 361000, PR China
b Xiamen Ocean Vocational College, Xiamen, 361000, Fujian, PR China
c School of Marine Equipment and Mechanical Engineering, Jimei University, Xiamen, 361000, Fujian, PR China
d School of Mechanical Engineering, Fuzhou University, Fuzhou, 350000, Fujian, PR China
e Fujian Hongshan Cogeneration Power Co., Ltd, Quanzhou, 362000, Fujian, PR China
⁎ Corresponding author. School of Marine Engineering, Jimei University, Xiamen 361000, PR China. zhilong.xu@163.com
⁎⁎ Corresponding author. mingliu@fzu.edu.cn
20 8 2024
15 9 2024
20 8 2024
10 17 e366331 7 2024
19 8 2024
20 8 2024
© 2024 The Authors
2024
https://creativecommons.org/licenses/by-nc/4.0/ This is an open access article under the CC BY-NC license (http://creativecommons.org/licenses/by-nc/4.0/).
In this study, to clarify the failure mechanism of the last-stage rotor blade in the low-pressure cylinder of a steam turbine, the peculiarities of crack initiation and propagation on the inlet side of the last-stage rotor blade at a distance of 125–165 mm were analyzed, along with the corrosion fatigue properties of its materials. The results showed that crack initiation occurred at the tip of the pit due to a combination of factors: stress concentration at the tip of the pit, corrosion of the Cr-poor area near the prior austenite grain boundary, centrifugal tensile stress, and steam bending stress. The crack propagation could be divided into the initial intergranular and late transgranular propagation stages. The main reason for the initial intergranular propagation was stress corrosion, and the main reason for the later transgranular propagation was corrosion fatigue. High-frequency induction quenching technology can improve the microhardness of the blade's surface material and enhance the blade's resistance to water erosion, but it may also reduce the corrosion fatigue resistance of the blade material. The rotary bending corrosion fatigue test can effectively simulate the crack propagation process of the blade. These results are of great significance for the safe operation of the last-stage rotor blade in the low-pressure cylinder of a steam turbine.

Graphical abstract

Image 1

Highlights

• Crack propagation includes initial intergranular propagation and late transgranular propagation.

• The main reason for the change in crack propagation mode is the change of the material microstructure.

• The main reason for intergranular propagation is stress corrosion, and the main reason for transgranular propagation is corrosion fatigue.

• The rotary bending corrosion fatigue test can effectively simulate the crack propagation process of a blade.

• An improper high-frequency induction quenching process reduces the corrosion fatigue properties of the material.

Keywords

Last-stage rotor blade
Martensitic stainless steel
High-frequency induction quenching
Crack
Corrosion fatigue test
==== Body
pmc1 Introduction

The steam turbine is a critical piece of power equipment in electric power, large ships, chemical industries, and other fields. During the operation of a steam turbine, due to the combined effects of the centrifugal force, the steam load, water erosion, and corrosion, rotor blades are prone to water erosion pits, stress corrosion, and fatigue corrosion [1], and then cracking or even fracture, which poses a significant threat to the safety of the unit. The last-stage rotor blade (LSRB) is the key component of the low-pressure cylinder of a steam turbine (LPST), and it is larger than other rotor blades and more prone to failure [2,3]. Therefore, it is of great practical significance for the safe operation of a steam turbine to analyze the causes of crack initiation and propagation in the LSRB of the LPST in complex environments and propose corresponding solutions.

In recent years, scholars have conducted fruitful research on the cracking of LSRBs. It was found that the impact and erosion of particles in wet steam [[4], [5], [6], [7], [8]], defects in the material microstructure and heat treatment process [7,[9], [10], [11]], fretting wear [[12], [13], [14]], stress corrosion [[15], [16], [17]], and corrosion fatigue [4,[18], [19], [20], [21]] lead to crack initiation and propagation. In a weakly corrosive wet steam environment, the LSRB is subjected to a large centrifugal force, which often causes typical intergranular fracture. Stress corrosion may cause intergranular fracture of the material [16]. In addition to stress corrosion, intergranular fracture may also be caused by hydrogen embrittlement [[22], [23], [24]], internal stress, and an uneven grain size distribution caused by heat treatment process defects [10,[25], [26]]. Therefore, the specific causes of intergranular fracture should be combined with the material heat treatment process and service environment for comprehensive analysis [27,28].

Due to the large size of the LSRB, the alternating steam load [29,30], and the difference in the material microstructure [31,32], the fracture form of the crack may change from intergranular fracture to transgranular fracture after the crack has extended a certain distance [[33], [34], [35]]. Since plastic deformation of the material rarely occurs, and the stress concentration sensitivity of the crack tip is large, resulting in a high crack propagation rate during intergranular fracture [36,37]. However, plastic deformation is the main form of material deformation at the crack tip when a transgranular fracture occurs. Due to the plastic slip of the material of the crack tip, the stress concentration sensitivity is reduced, and the crack propagation speed is low [38]. Although many scholars have studied the mechanisms of intergranular fracture and transgranular fracture, the cause of the transition from intergranular fracture to transgranular fracture has not been studied. Since the propagation rates of intergranular fracture and transgranular fracture are different [39,40], it is essential to analyze the transformation causes of crack propagation to accurately calculate the overall crack propagation rate and predict the blade life.

After six years of operation, cracks were found in several of the LSRBs of a 600 MW steam turbine [41]. After another year of operation, cracks were also found on four LSRBs that had not failed originally. Four cracks were found on the inlet sides of the LSRBs 125–165 mm from the shroud. Some of these cracks had the characteristics of intergranular fracture and transgranular fracture. The structure of the LSRB is shown in Fig. 1(a), and four cracks are shown in Fig. 1(b)–(e). These four cracks were new cracks that appeared in the past year, and the fracture morphologies of the cracks were not damaged. Two of the cracks were shorter than those in the literature [41], suggesting a need for further analysis of the location of crack initiation. In contrast, the other two cracks were longer than those reported in the literature [41], suggesting a similar urgent need for further analysis of the crack propagation characteristics. Therefore, it would be beneficial to perform additional studies based on previous failure analyses [41].Fig. 1 Last-stage rotor blade (LSRB) of a 600 MW supercritical steam turbine and its cracks. (a) Structure of the LSRB; the blue dotted area is the location of the crack, the yellow dotted area is the pit area, and the red area is the high-frequency induction quenching area. (b)–(e) Cracks of different lengths: (b) 100 μm, (c) 1.5 mm, (d) 5.2 mm, and (e) 13.5 mm. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.)

Fig. 1

Previous failure analyses did not involve testing samples from failed blades to determine the corrosion fatigue properties, and there are no equivalent experimental methods for accelerating the simulation of crack initiation and propagation on the inlet side of the LSRB. Therefore, this study first conducted a comparative analysis of four cracks to examine their locations, the causes of crack initiation, and the characteristics and mechanism of crack propagation. Then, samples were extracted from a failed LSRB, and rotary bending corrosion fatigue tests were carried out to analyze the corrosion fatigue properties of the materials. Finally, the results of the rotary bending corrosion fatigue tests were compared with the crack analysis findings to verify their reliability and demonstrate the feasibility of using rotary bending corrosion fatigue tests to accelerate the simulation of the crack propagation process on the inlet side of the LSRB.

2 Materials and methods

2.1 Materials

The LSRB of the steam turbine was made of martensitic stainless steel. The grade of the material was 1Cr12Ni3Mo2VN, its composition and mechanical properties met the standard requirements, and its Cr content was 11.55 %, as shown in Fig. 2(a). Furthermore, the heat treatment of the material was divided into two steps:① Overall heat treatment before machining. The material was quenched at 1010 °C for 40 min and cooled with oil. After quenching, the material was tempered at 590 °C twice for 120 min, and its mechanical properties are shown in Fig. 2(b).

② Local heat treatment after machining. The red area in Fig. 1(a) is the heat-treated area. The area was quenched by high-frequency induction at 1080 ± 20 °C, tempered at 500 °C, and air-cooled after holding for 4 h. The purpose of the local heat treatment was to improve the resistance of local areas to water erosion.

Fig. 2 Composition and mechanical properties of 1Cr12Ni3Mo2VN. (a) Contents of main elements of 1Cr12Ni3Mo2VN material [41]. (b) Mechanical property testing of blade matrix [41].

Fig. 2

The operating conditions of the LSRBs were as follows: ① The steam turbine was checked after one year of operation. The results showed that there were no abnormalities in the LSRB. ② After six years of operation, the steam turbine was rechecked. The results showed that cracks were found in some LSRBs. ③ After another year of operation, cracks were also found on four LSRBs that had not failed originally. Four cracks have been found on the inlet sides of the LSRBs 125–165 mm from the shroud, as shown in the blue dotted box in Fig. 1(a). In addition, many water erosion pits were found in the yellow dotted line box in Fig. 1(a), and the microstructure of the pits is shown in the lower left corner of Fig. 1(a). The crack lengths were 150 μm, 1.5 mm, 5.2 mm, and 13.5 mm, respectively, and the crack initiation was located at the tip of the pit on the inlet side, as shown in Fig. 1(b)–(e).

2.2 Methods

The study of cracks in the LSRB of LPST and the corrosion fatigue properties of materials included two parts: the first part was the failure analysis of the LSRB, while the second part was the rotary bending corrosion fatigue test, as shown in Fig. 3.Fig. 3 Study methods of crack initiation and propagation peculiarities.

Fig. 3

2.2.1 Failure analysis of the LSRB

The failure analysis of the LSRB included analyses of the microstructure, microhardness, and tensile strength of the inlet side; crack analysis of four lengths; and FEA analysis of centrifugal tensile stress and alternating bending stress, as shown in Fig. 3. The crack analysis of four lengths comprised four parts: ① microstructure analysis of the crack region, ② grain orientation analysis of the crack tip region via electron backscatter diffraction (EBSD), ③ fracture morphology analysis, and ④ energy-dispersive X-ray spectroscopy (EDS) analysis, as shown in Fig. 3.

Metallographic samples with four cracks of different lengths, EBSD samples, and fracture samples were prepared. First, the cracks with lengths of 100 μm (100-μm crack), 1.5 mm (1.5-mm crack), and 13.5 mm (13.5-mm crack) were selected to prepare the metallographic samples. Confocal laser scanning microscopy (CLSM) (VK-X3000K) and scanning electron microscopy (SEM) (Crossbeam 550) were used to observe the microstructure of the crack region. After microstructural analysis, the EBSD sample was prepared with a 1.5-mm crack, and EBSD (EDAX Hikari Plus FEI NOVA NANOSEM 450) was used to detect the grain orientation. After EBSD analysis, four cracks were prepared into fracture samples using the three-point bending method, and the fracture morphologies were analyzed. Because the 100-μm, 1.5-mm, and 13.5-mm cracks were corroded and contaminated when preparing metallographic samples, EDS analysis was only performed for the crack with a length of 1.5 mm (5.2-mm crack). SEM (Crossbeam 550) was used to photograph the fracture morphology and EDS (Oxford Xplore 30) analysis was performed in local areas.

2.2.2 Rotary bending corrosion fatigue test

According to ASTM F1801-97(2014), 10 rotary bending corrosion fatigue samples were prepared from the blade's bottom, as shown in Fig. 4(a). These samples were divided into two groups of five. One group (G-1) was subjected to high-frequency induction quenching, and the quenching depth was 1.5 mm. Conversely, the other group (G-2) was not subjected to this treatment. High-frequency induction surface treatment was carried out using a 20-kHz induction furnace (WKG-20KW) with a 7-turn coil; the total heating power was 20 kW and the total heating time was 4 s. After induction heating, water quenching was performed. The samples were then tempered at 500 °C for holding for 4 h before cooling with the furnace. The two groups of samples were installed on the support beam rotary bending corrosion fatigue tester (QBWP-6000J), and the corrosion fatigue test was carried out under the same conditions, as shown in Fig. 4(b). The number of loading cycles was recorded, and the microhardness values, microstructures, and fracture morphologies of the two groups of samples were analyzed.Fig. 4 (a) Rotary bending corrosion fatigue sample. (b) Support beam rotary bending corrosion fatigue tester.

Fig. 4

Because the ambient temperature of the steam turbine blades gradually decreases across all levels, humidity increases correspondingly. The LSRB operates in a weakly corrosive wet steam environment containing Cl− at a temperature of 35.2–68.2 °C [41]. In order to accelerate failure, the sample was sprayed with 3.5 % NaCl solution at 52 °C. To prevent interference of the vibration of the rotary bending corrosion fatigue tester during high-speed rotation, the speed was set to 3000 r/min (50 Hz). Each revolution of the sample subjects the material to one cycle of tensile stress followed by one cycle of compressive stress. According to (1), (2), (3), the loading force F was calculated to be 34.1 N. To accelerate the failure process, the actual loading force was set to 35 N.(1) σmax=MW

(2) W=πd332

(3) F=ML

where M is the maximum bending moment, W is the section modulus, L is the moment arm (L = 280 mm), d is the cross-section diameter of the circular sample (refer to Fig. 4(a), d = 6 mm), and σmax is the maximum bending stress (σmax was set to 450 Mpa, and σmax <0.5Rp0.2).

3 Results

3.1 Microstructure of the inlet side

A sample near the crack on the blade's inlet side was removed through wire cutting to analyze the material's microstructure and microhardness. After mechanical and electrolytic polishing and cleaning, the sample was observed under EBSD, as shown in Fig. 5. The results showed that the inlet side included three different microstructure regions, labeled I, II, and III, as shown by the white dotted lines in Fig. 5(b). Due to the direct action of high-frequency induction quenching, the microstructure of region I was fine martensite, as shown in Fig. 5(d). Region II was adjacent to the high-frequency induction quenching region, and its microstructure was medium-sized lath martensite, as shown in Fig. 5(e). Region III was far from the high-frequency induction quenching region, and its microstructure was coarse martensite, as shown in Fig. 5(f).Fig. 5 Microstructure and mechanical properties of the inlet side of the LSRB. (a) Structure of the LSRB. (b) Inlet side structure of the LSRB; the inlet side was divided into three different microstructure regions, labeled I, II, and III, as shown by the dotted white lines. (c) Microhardness of the inlet side. Series 1, 2, and 3 are the microhardness values corresponding to the three dotted lines in (b). (d)–(f) Electron backscatter diffraction (EBSD) images of the local areas of (b). The microstructures of (d) were fine martensite, the microstructures of (e) were medium-sized lath martensite, and the microstructures of (f) were coarse martensite. (g) Tensile strength test of the high-frequency induction quenching area. (h) Fracture morphology of the tensile specimen. (i) Dimple fracture morphology.

Fig. 5

3.2 Microhardness and tensile strength of the inlet side

Using a microhardness tester, the microhardness of the inlet side was measured in the direction of the three dotted lines in Fig. 5(b), and the spacing between two adjacent detection points was 2 mm. The measurement results are shown in Fig. 5(c). The microhardness of the fine martensite region (region I of Fig. 5(b)) was 450–480 HV0.1. The microhardness of the region dominated by medium-sized lath martensite (region II of Fig. 5(b)) and the coarse martensite region (region III of Fig. 5(b)) was 355–395 HV0.1. The relationship between the microhardness values was region I > region II ≈ region III, and region I was most susceptible to water erosion.

According to the ISO6892-1 standard, the high-frequency induction quenching area on the inlet side of the LSRB was sampled and processed into a plate tensile sample, as shown in Fig. 5(a). Tensile testing was carried out using a universal test machine (XBD5205), as shown in Fig. 5(g). The results showed that the tensile strength of the high-frequency induction quenching area was 1402 MPa, about 200 MPa higher than the tensile strength of the matrix area of 1200 MPa (Fig. 2(b)). The fracture morphology of the tensile sample was observed by SEM. The results showed that the fracture morphology was a dimple, and there was no intergranular fracture morphology, which proved that the intergranular fracture of the LSRB did not exist before service but occurred during service, as shown in Fig. 5(h) and (i).

3.3 Microstructure of crack region

The microstructures of the 100-μm and 1.5-mm crack regions were analyzed. Based on the microstructures of the three regions (I, II, and III) on the inlet side of Fig. 5(b), both crack regions exhibited fine martensite laths, both of which were consistent with the microstructure of the fine martensite region (I). The cracks did not extend into the medium-sized lath martensite (II), as shown in Fig. 6. There were three cracks with lengths of less than 60 μm around the 100-μm crack. Secondary cracks were present near the tips of some cracks, and the initiation positions of the cracks were all the tips of the pits, as shown in Fig. 6(a). The 1.5-mm crack was well preserved, and the overall path of crack propagation was straight, while the specific direction of crack propagation was tortuous, as shown in Fig. 6(b).Fig. 6 Microstructures of cracks with lengths of 100 μm and 1.5 mm. (a) Microstructure of the 100-μm crack region. (b) Microstructure of the 1.5-mm crack region.

Fig. 6

For the 1.5-mm crack, the crack initiation location was the pit tip, and the crack width gradually narrowed as the crack propagated, as shown in Fig. 7(a). From Fig. 7(a)–(c), it can be found that there were multiple secondary cracks in the initiation, middle, and tip regions of the crack. Some micro-particles were separated from the material by the main and secondary cracks, and these micro-particles may have been the prior austenite grains, as shown in Fig. 7(b). Fig. 7(d) and (e) show the SEM images of secondary crack tips in Fig. 7(a) and (c), respectively. The images show apparent prior austenite grain boundaries, with the propagation directions of the secondary crack's tips approximately parallel to the prior austenite grain boundaries. Therefore, the 1.5-mm crack could propagate along the prior austenite grain boundaries. The separated microparticles were the prior austenite grains. In Fig. 7(c), the main crack tip in the yellow dotted line frame could not be confirmed to spread along the prior austenite grain boundary, and further confirmation by EBSD was needed.Fig. 7 Microstructure of the 1.5-mm crack region. (a) Scanning electron microscopy (SEM) image of the entire crack. (b), (c) Confocal laser scanning microscopy (CLSM) images of the crack middle, and tip in (a), respectively. (d), (e) SEM images of the secondary crack and main crack tips in (a) and (c), respectively.

Fig. 7

Compared with the 1.5-mm crack, the 13.5-mm crack had a longer propagation path. Therefore, the initial propagation and later propagation of the 13.5-mm crack were analyzed. It was found that there was a secondary crack, which propagated along the prior austenite grain boundary during the initial propagation of the 13.5-mm crack, as shown in Fig. 8(a). There was a secondary crack that passed through the martensite lath in the later propagation of the 13.5-mm crack, as shown in Fig. 8(b). Therefore, the 13.5-mm crack could have two opposite fracture modes, which were early intergranular fracture and late transgranular fracture.Fig. 8 Microstructure of the secondary crack region of 13.5-mm crack. (a) Secondary crack propagated along the prior austenite grain boundary in the initial propagation of the 13.5-mm crack. (b) Secondary crack passed through the martensite lath in the later propagation of the 13.5-mm crack.

Fig. 8

3.4 Grain orientation of crack tip region

After grinding and electropolishing the metallographic sample with the 1.5-mm crack, the EBSD grain orientation of the main crack tip region in the yellow dashed box region in Fig. 7(c) was observed via SEM. Due to the grinding and electrolytic polishing, the surface material was removed, so the crack shape in the EBSD grain orientation diagram is not exactly the same as those in Fig. 7(a) and (c). However, the trend of crack propagation was basically the same, as shown in Fig. 9. The results showed that the sample had almost no residual austenite, and the grains had small lath shapes. The average equivalent grain diameters were mostly 3–11 mm, and the grains were fine lath martensite, as shown in Fig. 9(a). By comparing Fig. 5(d) and (e) with Fig. 9, it was found that the tip of the 1.5-mm crack did not extend to the medium-sized lath martensite (Region II). Except for local transgranular propagation, most of the 1.5-mm crack tip was fractured along the boundaries of the gains in different orientations, and the crack tip as a whole underwent intergranular propagation, as shown in Fig. 9(b).Fig. 9 Grain orientation in the tip region of the 1.5-mm crack region. (a) Image quality + boundaries + grain size. (b) Inverse pole figure + image quality + boundaries.

Fig. 9

3.5 Fracture morphologies

3.5.1 Fracture morphology of 100-μm crack

The 100-μm crack that was made into a metallographic sample was opened by the three-point bending method. After ultrasonic cleaning, the fracture morphology was observed via SEM (Crossbeam 550). The fracture included the pit zone and transient fracture zone, and there was a clear boundary between these zones, as shown in Fig. 10(a). Similar to the results in Fig. 6(a), there were also multiple micro-cracks at the tip of the pit, as shown by the black arrow in Fig. 10(b). There were no brittle fractures in the fracture region, but there were evident dimple features caused by tearing.Fig. 10 Morphology of the 100-μm crack fracture. (a) SEM image of the complete fracture; the upper right corner is the fracture diagram of the two sides opened by the three-point bending method. (b) Enlarged view of a local area in (a), with black arrows pointing to microcracks and white arrows pointing to dimples.

Fig. 10

3.5.2 Fracture morphology of 1.5-mm crack

The fracture of the 1.5-mm crack included a pit zone, propagation zone, and transient fracture zone, as shown in Fig. 11(a). There was a clear boundary between the pit zone and the propagation zone, as shown in Fig. 11(b). A rock-candy-shaped morphology was found in the propagation zone, the fracture surface was covered with many corrosion products, and there were secondary cracks and a small number of transgranular fracture characteristics, as shown in Fig. 11(c). The rock-candy-shaped fracture, had the appearance of many pieces of rock candy stacked on top of each other. The cracks were mainly intergranular fractures. The boundary between the propagation zone and the transient fracture zone was evident, and the transient fracture zone had apparent dimples, as shown in Fig. 11(d). The corrosion products in the propagation zone may have been formed by corrosion reactions during the preparation of the metallographic samples and EBSD samples or by the presence of corrosive media in the service environment. Therefore, the corrosion products in this fracture propagation zone could not be used as evidence of corrosion in the crack propagation process, and EDS analysis was not required.Fig. 11 Fracture morphology of the 1.5-mm crack. SEM images of the (a) complete fracture, where the upper right corner shows the fracture on both sides of the 1.5-mm crack, (b) pit zone and propagation zone, (c) propagation zone, and (d) propagation zone and transient fracture zone.

Fig. 11

3.5.3 Fracture morphology of 5.2-mm crack

The fracture of the 5.2-mm crack was covered with black corrosion products, as shown in the upper right corner of Fig. 12(a). The pit zone, propagation zone, and transient fracture zones can be clearly distinguished on the fracture, as shown in Fig. 12(a). Similar to the results in Fig. 11(b), there was also a clear boundary between the pit and propagation zones, as shown in Fig. 12(b). Compared with Fig. 11(c), the rock-candy-shaped morphology was also found in the propagation zone, and there were secondary cracks and a small number of transgranular fracture characteristics, as shown in Fig. 12(c). Compared with Fig. 11(d), there was also a sharp boundary between the propagation zone and the transient fracture zone, and the transient fracture zone was characterized by a significant dimple with no transition area near the boundary, as shown in Fig. 12(d).Fig. 12 Fracture morphology of the 5.2-mm crack. SEM image of the (a) complete fracture, where the upper right corner shows the fracture on both sides of the 5.2-mm crack, (b) pit zone and propagation zone, (c) propagation zone, and (d) propagation zone and transient fracture zone.

Fig. 12

3.5.4 Fracture morphology of 13.5-mm crack

Similar to the fractures of the 1.5- and 5.2-mm cracks, the fracture of the 13.5-mm crack also included a pit zone, propagation zone, and transient fracture zone. The difference was that the fracture propagation of the 13.5-mm crack was divided into initial and later propagation zones, as shown in Fig. 13(a). By comparing Fig. 5, Fig. 13, it was found that the crack had expanded from region I, dominated by fine martensite, to region II, dominated by medium-sized lath martensite. Similar to the results in Fig. 11, Fig. 12, there is also a clear boundary between the pit and propagation zones in Fig. 13(b). In the initial propagation zone, similar to the results in Fig. 11, Fig. 12, rock-candy-shaped intergranular fracture features and corrosion products were evident, as shown in Fig. 13(c). A small number of apparent transgranular fracture features were also present. The small planes and steps in the transgranular fracture had no noticeable tear ridge.Fig. 13 Morphology of the 13.5-mm crack fracture. (a) SEM image of complete fracture; the upper right corner is the fracture opened by the three-point bending method. (b)–(h) Enlarged images of different regions in (a): (b) pit zone and initial propagation zone, (c) initial propagation zone, (d) initial propagation and late propagation zones, (e)–(g) late propagation zone, and (h) late propagation zone and transient fracture zone.

Fig. 13

Two significantly different fracture characteristics were observed in the interface zone of the initial and late propagation. There was a relatively smooth plane in the initial propagation zone, while significant plastic deformation appeared in the later propagation, as shown in Fig. 13(d). A noticeable tear ridge, fatigue characteristics, and a few facets were found in the late propagation zone, as shown in Fig. 13(e)–(g). There was a clear boundary between the late propagation zone and the transient fracture zone, and many dimples were present in the transient fracture zone, as shown in Fig. 13(h). Therefore, the fracture morphology with a 13.5-mm crack had a larger late propagation zone with transgranular fracture characteristics compared with those of the fractures with 1.5- and 5.2-mm cracks, as shown in Fig. 11, Fig. 12, Fig. 13.

3.6 Energy-dispersive X-ray spectroscopy (EDS) analysis of local fracture area

After the fracture morphology analysis, EDS analysis was performed on the interface area between the propagation and transient fracture zones of the fracture of the 5.2-mm crack, as shown in Fig. 14(a). The surface scanning results showed that due to the enrichment of corrosion products in the propagation zone, the O content was significantly higher than that in the transient fracture zone, as shown in Fig. 14(b). Corrosive Cl and S were present in addition to O, as shown in Fig. 14(c). In contrast, no corrosive elements, such as Cl and S, were detected in the transient fracture zone.Fig. 14 EDS analysis of 5.2-mm crack fracture. (a) Interface zone between the propagation and transient fracture of the 5.2-mm crack. (b) O elemental surface scan result of (a). (c) Elemental content of the propagation zone.

Fig. 14

3.7 FEA analysis of centrifugal tensile stress and alternating bending stress

Due to the rotation driven by the steam load, the LSRB of the LPST is subjected to the combined action of centrifugal forces and alternating steam loads [42]. The frequency of the alternating bending stress is equal to the product of the number of stationary blades of the same stage and the rotor speed. The distributions of centrifugal tensile stress and steam bending stress of the last-stage stage rotor blade were obtained by finite element analysis (FEA). As shown in Fig. 15(a)–(g-1), the analysis process included reverse scanning, point data processing, reverse modeling, model optimization, meshing, and Abaqus simulation analysis. In the cyclic symmetry, the plane on which the black dashed line frame was located was the active plane, and the opposite side was the driven plane. The boundary condition was that the bottom of the model was completely fixed, and the load for the centrifugal tensile stress analysis was 314 rad/s, as shown in Fig. 15(f). Periodic pressure was selected in the load options to simulate alternating steam pressure in the area where it was applied [41], as shown in Fig. 15(g).Fig. 15 FEA analysis of centrifugal tensile stress and alternating bending stress: (a) LSRB of the low-pressure cylinder of a steam turbine (LPST) already in service. (b) Scan of the LSRB to obtain the point data model. (c) Reverse modeling using Geomagic Design X software. (d) Simplification of the model of the blade and rotor using SolidWorks software. (e) Grid partitioning using Hypermesh2021 software. (f) and (g) shows the boundary conditions, cyclic symmetry, and load types; (f-1) is the analysis result of the centrifugal tensile stress, while (g-1) is the result of the steam bending stress analysis.

Fig. 15

The FEA results showed that the tensile stress generated by the centrifugal force was large in the area 120–260 mm from the shroud, and reaches its maximum at the blade root, as shown in Figs. 15(f–1). In contrast, the maximum bending stress was generated by the alternating steam load on the inlet side, occurring at a distance of 140–200 mm from the shroud, as shown in Figs. 15(g–1). The four cracks were located at the inlet side 125–165 mm from the shroud, as shown in Fig. 15(a). Therefore, the actual crack location closely coincided with both the region of elevated centrifugal tensile stress and the region of maximum steam bending stress.

A grid independence analysis was conducted as follows: First, Hypermesh2021 software was used to divide the grids into four different sizes, as shown in Fig. 16(a)–(d). Then, FEA Analysis was completed using Abaqus 2019 software. The results showed that the maximum bending stress was localized in the specific area on the inlet side about 120–260 mm from the shroud, as shown in Fig. 17(a–1)–(d-2).Fig. 16 Grid partitioning using Hypermesh2021 software.

Fig. 16

Fig. 17 Results of the steam bending stress analysis: (a-1), (b-1), (c-1), and (d-1) use cyclic symmetric interactions, while (a-2), (b-2), (c-2), and (d-2) do not use cyclic symmetry interactions. The material properties, loading forces, and boundary conditions for all eight FEA analyses were consistent.

Fig. 17

3.8 Rotary bending corrosion fatigue test

3.8.1 Microhardness and microstructure

There was a hardness gradient in the G-1 material, as shown in Fig. 18. The surface microhardness of the G-1 material was high, while the core microhardness was low. In the range of 1.5 mm from the outer surface to the core, the microhardness of the material was 455–480 HV0.1, while the microhardness of the core material was 355–380 HV0.1. The G-2 material did not have a hardness gradient and had a microhardness of 360–380 HV0.1. There was a microstructural gradient in the G-1 material, as shown in Fig. 19. The surface region subjected to high-frequency induction quenching fine martensite, as shown in Fig. 19(a), while the core region was coarse martensite, as shown in Fig. 19(b). However, the G-2 material was coarse martensite, with no microstructural gradient, as shown in Fig. 19(c).Fig. 18 Microhardness values of two groups of samples.

Fig. 18

Fig. 19 Microstructures of the two groups of tests. (a) Microstructure of the G-1 surface region. (b) Microstructure of the G-1 core region. (c) Microstructure of the G-2 material.

Fig. 19

3.8.2 Loading cycle number and fracture morphology

It can be seen from the loading cycle number of the rotary bending corrosion fatigue test that the corrosion fatigue performance of G-1 was weaker than that of G-2, as shown in Fig. 20. The average loading cycle numbers of G-1 and G-2 were 376,159 (run time: 125 min and 23 s) and 559,397 (run time: 186 min and 28 s), respectively. Therefore, high-frequency induction quenching process can easily reduce the corrosion fatigue performance.Fig. 20 Loading cycle numbers of corrosion fatigue tests of the two groups of samples.

Fig. 20

According to the fracture morphologies, the fractures can be divided into a crack initiation zone, crack propagation zone, and transient fracture zone, as shown in Fig. 21, Fig. 22. The crack initiation locations were discontinuous machining defects, and there was distinct radiation from the crack initiation location to the material direction in the fracture, as shown in Fig. 21, Fig. 22. Machining defects can easily cause stress concentration and crack initiation. There were two types of fracture morphologies in the crack propagation zone of G-1. In the initial propagation zone, the fracture morphology was rock-candy-shaped intergranular fracture, and a few transgranular fractures were found, as shown in Fig. 21(b). However, there were quasi-cleavage and fatigue characteristics in the late propagation zone, as shown in Fig. 21(c) and (c-1). No rock-candy-shaped intergranular fracture was found in the G-2 propagation zone, and only quasi-cleavage and fatigue striation were found, as shown in Fig. 22(b), (c), and (c-1). Therefore, the high-frequency induction quenching process could easily cause intergranular fracture of the LSRB material in the rotary bending corrosion fatigue test.Fig. 21 Fracture morphology of G-1. (a) Complete morphology of the fracture. (b), (c) Local enlarged image of (a), and (c-1) local enlarged image of (c).

Fig. 21

Fig. 22 Fracture morphology of G-2. (a) Complete morphology of the fracture. (b), (c) Local enlarged image of (a), and (c-1) local enlarged image of (c).

Fig. 22

4 Discussion

4.1 Crack initiation

Due to the impact of water erosion, the pits on the inlet side gradually deepened [20,43], and the stress concentration coefficient at the tip of the pit increased. From the tip of the blade to the axis of rotation, the size of the pit gradually decreased. The largest pit was located at the tip of the blade, and no crack area was detected in the region. Based on the distribution of the centrifugal tensile stress, the tensile stress at the blade bottom was the largest, but no pit was present there. The centrifugal tensile stress was greater at distances of 120–260 mm from the shroud, and the crack region was in this region. Therefore, crack initiation was not only related to centrifugal tensile stress but also to the pit, as shown in Fig. 23.Fig. 23 Crack initiation is not only related to centrifugal tensile stress, but also to pits.

Fig. 23

At the same time, due to the corrosive elements O, Cl, and S in the wet steam, the Cr-poor zone at the prior austenite grain boundary in the fine martensite region was corroded, and the stress concentration sensitivity of the grain boundary in this region increased [41]. Finally, the prior austenite grain boundary near the tip of the pit induced crack initiation under the combined action of steam bending stress [41] and centrifugal tensile stress. The detection results of the microstructure and fracture morphology of the 100-μm crack both indicated that the crack originated at the tip of the pit, as shown in Fig. 6, Fig. 10. After crack initiation, the main and secondary cracks spread rapidly, and the prior austenite grains near the tip of the pit separated from the material, as shown in Fig. 7(b) and (c). Therefore, the crack initiation may have been caused by the combination of stress concentration at the tip of the pit, corrosion in the Cr-poor zone near the prior austenite grain boundary, centrifugal tensile stress, and steam bending stress.

4.2 Initial intergranular fracture

If the Cr-poor zone near the prior austenitic grain boundary of the material were corroded, it could easily cause stress corrosion cracking and lead to rapid crack propagation. After the blade material was machined, the tempering temperature of the heat treatment was 500 °C. The tempering interval of 460–540 °C of martensitic stainless steel was sensitized, the carbide was precipitated along the prior austenite grain boundaries, and Cr was separated from the solid solution at the grain boundaries [[44], [45], [46], [47], [48]]. Due to the slow diffusion rate of Cr, it could not be supplemented from the solid solution within the grain to the boundary, so it could only consume Cr near the grain boundary, resulting in a Cr-poor area near the grain boundary [49]. In the corrosive media, the Cr-poor zone near the prior austenite grain boundary was the anode, and the prior austenite grain boundary and other parts were the cathodes, thus forming a short-circuit battery effect. Because the Cr-poor zone was narrow, it dissolved quickly [50,51]. Under stress corrosion, the cracks propagated along the prior austenite grain boundaries, as shown in Fig. 24.Fig. 24 Crack propagation along the prior austenite grain boundary.

Fig. 24

The evidence of the initial crack propagation along the prior austenite grain boundary was as follows. ① The CLSM analysis showed that the 1.5-mm crack could propagate along the prior austenite grain boundaries, and main and secondary cracks separated the prior austenite from the material, as shown in Fig. 7(b). The secondary crack propagated along the prior austenite grain boundary in the initial propagation of the 13.5-mm crack, as shown in Fig. 8(a). ② The SEM analysis showed that the crack tips of the secondary cracks propagated along the prior austenite grain boundaries, as shown in Fig. 7(d) and (e). ③ The EBSD detection results showed that the tip of the 1.5-mm main crack mainly expanded along the boundary of fine martensite with different grain orientations, as shown in Fig. 9. ④ The entire propagation zone of the 1.5- and 5.2-mm crack fracture and the initial propagation zone of the 13.5-mm crack had typical rock candy shape characteristics, as shown in Fig. 11, Fig. 12, Fig. 13. The following conditions of stress corrosion have been verified in this study: ① The crack initiation locations coincided with the larger centrifugal tensile stress distribution region and the maximum steam bending stress distribution region [41], as shown in Fig. 15(f–1)–(g-1). ② Corrosive elements O, Cl, and S were detected in the fracture propagation zone, as shown in Fig. 14.

4.3 Initial transgranular fracture

In the fine martensite region on the inlet side (region I in Fig. 5(b)), in addition to the Cr-poor region near the prior austenite grain boundaries due to the enrichment of carbide, a small number of Cr-poor regions were also present at the martensite lath boundaries. Therefore, under stress corrosion, cracks could propagate through the prior austenite grain boundaries and along the martensite lath boundaries [52].

As shown in Fig. 25(a), crack segments 1, 2, 5, and 6 underwent intergranular propagation, and crack segments 3 and 4 underwent transgranular propagation. The propagation direction of crack segment 2 deviated significantly from prior austenite grain boundaries a and b, while the direction deviation from the martensite lath boundary was slight, which led to the continuation of the crack propagation along the martensite lath boundary and the formation of crack segment 3. The propagation direction of crack segment 3 slightly differed from the direction of the next martensite lath boundary, resulting in crack propagation along the direction of crack segment 4. The propagation direction of crack segment 4 deviated significantly from the direction of the martensite lath boundary in the next prior austenite grain, resulting in further propagation of the crack along the prior austenite grain boundary and forming crack segment 5. The small number of transgranular fracture features without prominent tearing characteristics in the initial propagation zone shown in Fig. 11, Fig. 12, Fig. 13 confirmed this conjecture.Fig. 25 Path of initial crack propagation. (a) Schematic diagram of crack propagation along the prior austenite grain boundaries and martensite lath boundaries. A few martensite laths may be torn during the process of intergranular propagation. (b) Partial enlargement of (a). (c) Partial enlargement of (b).

Fig. 25

Due to the disordered growth of martensite laths in the prior austenite, cracks could be blocked by the martensite laths during the propagation process, which tore the martensite laths and then continued to propagate along the boundaries of the laths, as shown in Fig. 25(b) and (c). This hypothesis was confirmed by the tearing of martensite laths in the enlarged area of the white box in Fig. 9(b). Therefore, after the crack initiation at the inlet side of the blade, under the action of stress corrosion, the crack mainly propagated along the prior austenite grain boundaries. A small number of cracks spread along the martensite lath boundaries, and a small number of martensite laths were torn during the process of intergranular propagation.

4.4 Late transgranular fracture

As the crack propagated from the fine martensite region (region I) to the region dominated by medium-sized lath martensite (region II), the crack propagation pattern gradually changed from intergranular fracture to transgranular fracture. This transition from intergranular to transgranular propagation may have been triggered by the material's microstructure and stress state [35,[53], [54], [55]]. The crack propagated to the region dominated by medium-sized lath martensite for the following reasons: ① the lath martensite and prior austenite grains were large, and it was difficult to continue propagation along the grain boundary [56], ② the LSRB was subjected to alternating steam bending stress in addition to centrifugal forces, and ③ there were corrosive elements O, Cl, and S in the operating environment, so the material experienced corrosion fatigue. Under corrosion fatigue, the material at the crack tip underwent plastic deformation and slipped due to reaching the fatigue limit. Due to the slip of the material, the stress concentration sensitivity of the crack tip material decreased, and the crack propagation rate dropped [54]. Due to alternating bending stress, the material at the crack tip was torn, and a periodic tear edge was formed. The cracks gradually changed from intergranular propagation to transgranular propagation.

In the process of crack propagation, a small number of cracks propagated along the grain boundary, resulting in the formation of a facet of the fracture and a tear edge at the end of the facet [57], which was confirmed by the discovery of facets in Fig. 13(e), (f), and (g). At the same time, the tear edges and fatigue characteristics observed via SEM of the 13.5-mm crack fracture confirmed that the material experienced fatigue and transgranular propagation, as shown in Fig. 13(e), (f), and (g). In addition, the 5.2-mm crack was not made into metallographic and EBSD samples, so the fracture was not corroded and contaminated. The discovery of corroding elements O, Cl, and S in the propagation zone confirmed that corrosion occurred during the crack propagation, as shown in Fig. 14. Therefore, due to corrosion fatigue, the cracks gradually transitioned from initial intergranular propagation to later transgranular propagation, as shown by the yellow arrow in Fig. 26. Moreover, when the crack extended to the martensite lath boundary, the crack propagated along the martensite lath boundary due to the action of stress corrosion, as shown by the red arrow in Fig. 26.Fig. 26 Crack propagation patterns in coarse martensite.

Fig. 26

4.5 Corrosion fatigue properties of blade materials

In the rotary bending corrosion fatigue test, the sample was subjected to periodic tensile and compressive stresses in the corrosive environment, as shown in Fig. 27(a). First, cracks initiated from machining defects on the surface of the material, such as discontinuous knife marks, as shown in Fig. 21, Fig. 22. Then, the cracks propagated under the combined action of the alternating stresses and corrosive media. After the crack initiation in the surface layer of the G-1 sample, the crack propagation path was from the high-frequency induction quenching area of the surface layer to the matrix material area of the core, as shown in Fig. 27(b) and (c). Finally, the sample broke.Fig. 27 Experimental principle of corrosion fatigue in rotary bending. (a) Sample is subjected to periodic loading forces in a corrosive environment. (b) Stress in the crack region. (c) After half a cycle, the stress in the crack region.

Fig. 27

By comparing the microhardness and microstructure of the inlet side, it was found that the microstructure of the G-1 material also had a gradient of the microhardness and microstructure, as shown in Fig. 5, Fig. 18, Fig. 19. Furthermore, by comparing the rotary bending corrosion fatigue test results with the fracture morphology of the failed blade, it was found that the fracture morphology of G-1 was consistent with that of the failed blade, as shown in Fig. 13, Fig. 21. The initial propagation of cracks was mainly intergranular fracture, and a small amount was transgranular fracture, as shown in Fig. 13, Fig. 21. The late propagation form exhibited transgranular fracture with fatigue characteristics, as shown in Fig. 13(e)–(g) and 21(c) and (c-1). It is worth noting that cracks appeared in the LSRB after 6–7 years of operation, and G-1 fractured after 47 min and 49 s of the rotary bending corrosion fatigue test. Therefore, the rotary bending corrosion fatigue test, which could effectively simulate the crack propagation process of the blade, is an important method to test the corrosion fatigue properties of blade materials.

According to the test results of the rotary bending corrosion fatigue, although the surface microhardness of G-1 was higher than that of G-2, the number of loading cycles of G-1 was smaller than that of G-2, as shown in Fig. 18, Fig. 20. This showed that the high-frequency induction quenching technology could improve the microhardness of the surface material of the blade, improve the water corrosion resistance of the surface material, and reduce the corrosion fatigue resistance of the blade material. Therefore, it is necessary to optimize the high-frequency induction quenching process to improve the corrosion fatigue resistance of the material surface while improving the water erosion resistance.

5 Conclusions

In this study, through the comparative analysis of the microstructure, EBSD results, fracture morphology, and EDS results of cracks of different lengths in the LSRB of the LPST, as well as rotary bending corrosion fatigue tests, the crack propagation characteristics on the inlet side 125–165 mm from the shroud were investigated. The conclusions were as follows:(1) The crack initiation occurred at the tip of the pit. The crack propagation was divided into the initial propagation stage, dominated by intergranular fracture, and the later propagation stage, dominated by transgranular fracture. The crack initiation may have been caused by the combination of stress concentration at the tip of the pit, corrosion in the Cr-poor zone near the prior austenite grain boundary, centrifugal tensile stress, and steam bending stress.

(2) The initial intergranular propagation of cracks was mainly caused by stress corrosion. Under the action of stress corrosion, the cracks mainly propagated along the prior austenite grain boundaries. A small number of cracks spread along the martensite lath boundaries, and a small number of martensite laths were torn during the process of intergranular propagation.

(3) The late transgranular propagation of cracks was mainly caused by corrosion fatigue. As the crack propagated from the fine martensite region (region I) to the region dominated by medium-sized lath martensite (region II), the crack propagation pattern gradually changed from intergranular fracture to transgranular fracture. Under the combined action of alternating bending stress and a weak corrosion environment, the material at the crack tip reached the fatigue limit, and the crack gradually changed from intergranular fracture to transgranular fatigue fracture.

(4) The rotary bending corrosion fatigue test could effectively simulate the crack propagation process of the blade. The high-frequency induction quenching technology could improve the microhardness of the surface material of the blade, improve the water corrosion resistance of the surface material, and reduce the corrosion fatigue resistance of the blade material.

Data availability statement

No data was used for the research described in the article.

CRediT authorship contribution statement

Jiashun Gao: Writing – original draft, Software, Methodology, Investigation, Formal analysis, Conceptualization. Zhuolin Tang: Writing – original draft, Data curation. Bicheng Guo: Supervision, Resources. Zhilong Xu: Writing – review & editing, Supervision, Resources, Funding acquisition, Conceptualization. Ming Liu: Validation, Software, Funding acquisition. Wenbing Sun: Writing – review & editing, Visualization. Zhenye Zhao: Visualization, Investigation.

Declaration of competing interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

Acknowledgments

The work described in this paper was supported by the Science and Technology Project of Fujian Province, in China [No. 2022HZ024009 , 2021HZ024006 ], 10.13039/501100003392 Fujian Provincial Natural Science Foundation , in China [No. 2020 J01693 ], Special Fund for Marine Economic Development of Fujian Province, in China [No. FJHJF-L-2021-9 ], the Education Research Project of the Fujian Provincial Department of Education, in China [No. JAT210806 ], and the Major Science and Technology Project of Xiamen, Fujian, in China [No. 3502Z20231011 ].
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