Effect of Overheating Temperature on Thermal Cycling Creep Properties of K465 Superalloy

Abstract

Turbine blades in aircraft engines may encounter overheating and suffer serious creep property degradation. In this study, the thermal cycling creep experiments were conducted on K465 superalloy under (900 °C/30 min–1100 °C/3 min)/50 MPa, (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa. The investigated thermal cycling creep properties were dramatically degraded, and increasing the overheating temperatures significantly decreased the thermal cycling creep life. The secondary γ′ precipitates obviously dissolved and the area fraction decreased to around 35.2% under (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa, which was almost half that after the standard solution treatment. The decline of the thermal cycling creep properties was mainly due to the significant dissolution of γ′ precipitates. The creep holes/cracks were mainly distributed at the M₆C carbides and γ/γ′ eutectics interfaces, M₆C carbides and γ′ film interfaces in the grain boundaries, and resulted in the final intergranular fracture.

1. Introduction

Ni-based superalloys are generally selected as the materials of aircraft turbine blades [1,2]. During actual operation, the engine encounters emergencies such as inlet distortion [3] or one engine inoperative (OEI) conditions [4], which may result in a sharp temperature jump. This can pose a serious threat to the safe service of turbine blades, and the thermal cycling creep is always used to modify this specific service condition by airworthiness authorities such as European EASA [5].

Studies on thermal cycling creep have attracted attention since the 2010s due to the work of Cormier et al. on single-crystal Ni-based superalloys [5,6,7,8,9]. Recently, Guo et al. [10], Yuan et al. [11] and An et al. [12] began to investigate the thermal cycling creep behavior of equiaxed cast and directionally solidified superalloys, and Antonov et al. [13] briefly summarized the thermal cycling creep properties of conventionally cast, directionally solidified and single crystal superalloys. Overheating can result in the obvious dissolution of strengthening γ’ precipitates, and thermal stress due to the temperature variation during the thermal cycling creep promotes the initiation of creep holes. The thermal cycling creep properties are even worse than those under overheating isothermal creep properties. However, there is still a lack of research on polycrystalline alloys, and the evolution law of thermal cycling creep properties of polycrystalline superalloys is not clear enough.

K465 superalloy is widely used to manufacture the aero-engine turbine blades, while the overheating service caused by the sudden temperature jump also poses a serious threat to its safe service [14]. Previous studies have shown that the thermal cycling creep property under (900 °C/30 min–1100 °C/3 min)/70 MPa was even worse when compared to the isothermal creep property under 1100 °C/70 MPa and 900 °C/70 MPa [10], while the effect of overheating temperature variation on the thermal cycling creep performance of the K465 alloy is still unclear. Exploring the above point is much more meaningful, considering that K465 alloy will encounter overheating at different temperatures, such as 1100 °C and 1150 °C.

In this study, we conducted three thermal cycling creep experiments on K465 superalloy to reveal the overheating effect on thermal cycling creep properties. The microstructure characterization and creep cracks were investigated. This study aims to briefly introduce the degradation law of the thermal cycling creep of conventionally cast superalloys.

2. Materials and Methods

All the as-received solid bars were manufactured using investment casting technology and experienced the standard solution treatment, which composed of solution treatment at 1210 °C for 4 h, followed by air cooling. The main chemical composition of K465 superalloy is Ni, 9.90 W, 9.50 Co, 8.66 Cr, 5.10 Al, 2.52 Ti, 1.44 Mo, 0.95 Nb, 0.16 C (in wt%). The microstructure morphologies have already been reported in Ref. [10]. The thermal cycling creep tests were conducted using our own creep testing rig, which has been systematically described in our previous article [10]. The normal allowable maximum service temperature of K465 alloy is 1000 °C and 1050 °C when used as turbine blades and vanes, respectively.

Figure 1a depicts the procedure of thermal cycling creep; creep temperature cycles were repeated at a temperature under 50 MPa. Each cycle consisted of a 30 min duration at lower temperature T₁, a 4 min heating from T₁ to T₂, a 3 min overheating at T₂ and a 5 min cooling from T₂ to T₁. Figure 1b depicts the size of the experimental specimen, and the gauge length and diameter was set as 40 mm and 5 mm, respectively. The creep tests under (900 °C/30 min–1100 °C/3 min)/50 MPa were interrupted after 191.8 h and the creep tests under (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa were carried out until the samples fractured. It should be noted that the deformation is expressed by the total macro displacement of the sample.

Table 1. The creep testing conditions of the thermal cycling creep of K465 superalloy.

Specimen Number	T1/°C	T2/°C	Displacement/mm	Creep Time/h	Cycles
1#	900	1100	2.52	191.8	268
2#	900	1150	5.22	56.5	69
3#	1000	1150	6.47	20.4	26

shows the measured creep property with various overheating temperatures.

Samples were prepared for microstructural observation after grinding and polishing. The detail etching method has been described in detail in Ref. [15]. Microstructural observations were conducted through a ZEISS AXIO image A2m optical microscope (OM) and ZEISS SUPRE 55 scanning electron microscope (SEM). The quantitative statistics of the area fractions of γ′ precipitates were based on Chinese standard GB/T 15749-2008 through Photoshop software. At least three images were used to get a statistical result.

3. Results and Discussion

3.1. Thermal Cycling Creep Properties

Figure 2a presents the experimental creep deformation displacement vs. the time of the thermal cycling creep under (900 °C/30 min–1100 °C/3 min)/50 MPa (1# specimen). The overall displacement of the 1# specimen increased with extending the creep time, and the creep test was interrupted at the 269th cycle. A partially enlarged image of the creep curve is shown in Figure 2b. A displacement peak appears with the temperature jump, and is consistent with the reports by Cormier [16], Guo [10] and An [12] et al. Generally, the peak strain is caused by the changes in creep temperature, and consists of elastic strain, plastic strain and thermal strain. Figure 2c shows the experimental creep deformation vs. the time of the thermal cycling creep under (900 °C/30 min–1150 °C/3 min)/50 MPa (2# specimen) and (1000 °C/30 min–1150 °C/3 min)/50 MPa (3# specimen); the alloy fractured after 69 and 26 cycles, respectively. Under the above two conditions, the final fracture of the alloy occurred in the overheating stage of 1150 °C/3 min. After removing the elastic and thermal strains, the plastic strain was obtained [12]. Figure 2d depicts the plastic creep displacement vs. time curves under (900 °C/30 min–1100 °C/3 min)/50 MPa, (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa. Obviously, increasing the creep temperatures T₁ and T₂ during the thermal cycle creep will significantly reduce the creep properties of the alloy.

3.2. Microstructure Characterizations

To reveal the overheating effect on the thermal cycling creep properties of K465 alloy, the main reinforcing phase γ′ phase and creep voids were systematically characterized. The morphologies of the γ’ precipitates in creep specimens 1# and 2# are shown in Figure 3a,b. For 1# specimen, the secondary γ′ precipitates were obviously dissolved and directionally coarsened. It is worth noting that there were multiple ultra-fine γ′ particles in the γ channel, and they were precipitated after the creep testing during the cooling process and are not concerned in the present work. When T₂ temperature rose to 1150 °C, the fracture microstructure showed that γ’ precipitates dissolved more obviously, as shown in Figure 3b. The specific thermal exposure experiments of K465 alloy were carried out to analyze and evaluate the degradation of γ′ phase. Figure 3c–e are SEM-SE images showing the morphologies of γ’ precipitates in the dendrite core regions of K465 superalloy after being heat treated at 900 °C/1000 h 1100 °C/3 min and 1150 °C/3 min, respectively. The area fraction of γ’ precipitates are listed in

Table 2. Area fraction of γ′ precipitates in the dendrite core regions of K465 superalloy after the standard solution treatment, during thermal cycling creep under different conditions and after heat treatment at 900–1150 °C.

Condition	Specimen Number	γ′ Area Fraction/%	Interrupted Temperature/°C
Standard solution treatment	-	61.7 ± 1.4	-
Thermal cycling creep	1#	46.5 ± 3.2	1100
	2#	36.1 ± 1.8	1150
	3#	35.2 ± 2.4	1150
900 °C/1000 h	-	67.8 ± 1.2	900
1100 °C/3 min	-	50.5 ± 3.3	1100
1150 °C/3 min	-	34.1 ± 3.1	1150

.

In normal conditions (below 1000 °C), γ′ precipitates basically maintained cubic morphology, as shown in Figure 3c. The area fraction of γ′ phase kept 67.8 ± 1.2% after aging at 900 °C for 1000 h, even 6% larger than that after the standard solution treatment. This is mainly because the standard heat treatment temperature was semi solid solution treatment, and the forming elements of γ′ phase are still supersaturated after air cooling. Therefore, when exposed to a relatively low heat treatment temperature of 900 °C, γ′ phase precipitated completely, resulting in an increase of its area fraction [17]. The solvus temperature of γ′ phase in the dendrite core region is 1210 °C [18]. After short-time overheating at 1100 °C and 1150 °C, significant dissolution of γ′ phase occurred and the higher the temperature was, the more obvious this phenomenon was (Figure 3d,e). This is mainly due to the increase of the saturation of γ′ phase enriched elements with the overheating temperature increasing, resulting in the decrease of area fraction of the γ′ precipitates [18]. Especially after overheating treatment at 1150 °C, the area fraction of the γ′ phase decreased to 34.1 ± 3.1%, which was only 1/2 of that after the standard heat treatment.

The temperature change will lead to the dissolution and re-reprecipitation of γ′ precipitates continuously during the thermal cycling creep [10,12]. We previously identified that when crept under (900 °C/30 min–1100 °C/3 min)/70 MPa, γ’ precipitates dissolved during the (T₁–T₂)/4 min and T₂/3 min stages, and partially recovered during the (T₂–T₁)/5 min and T₁/30 min stages [10]. The results of this study show that for the 1# specimen, the area fraction of the γ’ precipitates decreased to 46.5 ± 3.2% and was even lower than that after overheating at 1100 °C/3 min, which may be due to the acceleration of stress on γ’ dissolution during the creep process. For the 2# and 3# specimens, the area fractions of γ′ precipitates were both close to that after overheating at 1150 °C/3 min, and the microstructural degradation of γ′ precipitates was serious. However, it should be noted that the temperature T₁ of 2# specimen was 900 °C, which was lower than the 1000 °C of 3# specimen. Therefore, in other stages of the thermal cycling creep, including T₁/30 min stage, (T₁–T₂)/4 min stage, (T₂–T₁)/5 min stage, the area fraction of γ′ precipitates in 2# specimen was higher than that in 3# specimen. More detailed characterizations of γ′ phase microstructure are in progress, but it is certain that the significant γ′ phase degradation is mainly responsible for the degradation of the thermal cycling creep of K465 superalloy. Whether T₁ or T₂, the higher the overheating temperature, the more serious the damage of thermal cycling creep performance (Table 1 and Figure 2).

At the same time, the weak regions in the process of thermal cycling creep, namely the creep hole/crack initiation regions, were systematically studied. Under the three creep conditions in this study, the distribution characteristics of creep holes/cracks were basically the same. Figure 4a,b are SEM-SE images showing the creep cracks distributed at the interfaces between M₆C carbides and γ/γ′ eutectics when crept under (900 °C/30 min–1100 °C/3 min)/50 MPa and (900 °C/30 min–1150 °C/3 min)/50 MPa. Compared with the standard solution treated microstructure [19], the coupling effect of creep stress and cyclic thermal stress promoted the formation of M₆C carbides, resulting in the initiation of more creep holes/cracks. In addition, the interfaces of M₆C carbides and γ′ film also easily initiated creep holes/cracks and lead to the final intergranular fracture (Figure 4c,d). Overall, the interfaces of M₆C carbides and γ/γ′ eutectics in the interdendritic regions and that of M₆C carbides and γ′ film in the grain boundaries easily initiated creep holes during the thermal cycling creep, and the creep cracks in the grain boundaries finally resulted in the intergranular fracture.

4. Conclusions

The thermal cycling creep was conducted under (900 °C/30 min–1100 °C/3 min)/50 MPa, (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa. Combined with the systematic microstructural characterization, the overheating effect on thermal cycling creep properties of K465 superalloy was revealed. The main conclusions are summarized as follows:

(1)

Creep life of thermal cycling creep was longer than 268 h under (900 °C/30 min–1100 °C/3 min)/50 MPa (268 cycles), while it decreased to 56.5 h and 20.4 h under (900 °C/30 min–1150 °C/3 min)/50 MPa (69 cycles) and (1000 °C/30 min–1150 °C/3 min)/50 MPa (26 cycles), respectively. Increasing the creep temperatures T₁ and T₂ may significantly decrease the thermal cycling creep property of K465 superalloy.

(2)

The secondary γ’ precipitates dissolved significantly, and the area fraction decreased to 46.5 ± 3.2%, 36.1 ± 1.8%, 35.2 ± 2.4% under (900 °C/30 min–1100 °C/3 min)/50 MPa, (900 °C/30 min–1150 °C/3 min)/50 MPa and (1000 °C/30 min–1150 °C/3 min)/50 MPa, respectively. The obvious degradation of γ′ precipitates plays the most important role in the degradation of the thermal cycling creep properties.

(3)

The creep holes/cracks were mainly distributed at the interfaces of M₆C carbides and γ/γ′ eutectics, M₆C carbides and γ′ film in the grain boundaries. Intergranular fracture finally occurred under the investigated thermal cycling creep conditions.