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Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior
Accepted Manuscript Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior Wei Zhang, Xiaowei Wang, Xiang Li, Jianming Gong, Magd Abdel Wahab PII: DOI: Reference:
S0142-1123(18)30001-X https://doi.org/10.1016/j.ijfatigue.2018.01.001 JIJF 4530
To appear in:
International Journal of Fatigue
Received Date: Revised Date: Accepted Date:
13 September 2017 31 December 2017 2 January 2018
Please cite this article as: Zhang, W., Wang, X., Li, X., Gong, J., Abdel Wahab, M., Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior, International Journal of Fatigue (2018), doi: https://doi.org/10.1016/j.ijfatigue.2018.01.001
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Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior Wei Zhang 1,2, Xiaowei Wang 1,2,3,, Xiang Li 1,2, Jianming Gong 1,2, Magd Abdel Wahab 4,5 1
School of Mechanical and Power Engineering, Nanjing Tech University, Nanjing 211816, China
2
Jiangsu Key Lab of Design and Manufacture of Extreme Pressure Equipment, Nanjing 211816, China
3
Faculty of Engineering and Architecture, Ghent University, Zwijnaarde B-9052, Belgium
4
Institute of Research and Development, Duy Tan University, 03 Quang Trung, Da Nang, Vietnam
5
Soete Laboratory, Ghent University, Ghent 9000, Belgium
ABSTRACT This work is devoted to the study of the influence of prior low cycle fatigue (LCF) on microstructure evolution and subsequent creep behavior of P92 steel. Different strain amplitudes prior LCF tests ranging from 0.25% to 0.6% were performed at 650 °C with constant strain rate of 1.0×10-3 s-1 and thereafter subsequent creep tests were conducted at 130 MPa in the same temperature. Optical microscope (OM), scanning electron microscope (SEM) and transmission electron microscope (TEM) were utilized to clarify prior LCF damages of various strain amplitude on subsequent creep behaviors. Results reveal that the increase of lifetime fraction and the increase of strain amplitude lead to the nonlinear degradation of subsequent creep properties. However, the degradation behavior shows a tendency to be saturated after 70% lifetime fraction or after 0.4% strain amplitude, respectively. The degradation is lifetime fraction and strain amplitude of prior LCF dependent. Further analysis of microstructure indicates that the degradation behavior is attributed to the different evolution behavior of martensite laths driven by cyclic deformation during prior
Corresponding author. Tel.: +86 25 58139361; Fax: +86 25 58139361. E-mail address: [email protected] (Xiaowei Wang).
LCF. Additionally, fractographs of failed creep specimens were analyzed in detail to correlate the different subsequent creep behaviors. Key words: Prior low cycle fatigue; Creep behavior; Microstructure
1. Introduction ASME grade P92 steel (with the addition of W and B and a reduction of Mo in P91 steel) is widely used for ultra-supercritical (USC) power plants [1]. Excellent creep strength, lower thermal expansion coefficient, higher thermal conductivity, good weld ability and better corrosion resistance in water steam systems are important features for its suitability in USC industry. It is also considered as favored candidate materials for sodium or lead cooled reactors as well as gas-cooled reactors in innovative Generation IV plants because of its excellent irradiation resistance to void swelling [2,3]. Normally, during service operation of USC plants, as a result of frequent start-ups, shut-downs and compensating of renewable energy fluctuating output, the components often undergo repeated thermal stresses which result in strain controlled low cycle fatigue (LCF) damage of the used materials [4]. Moreover, long-term of constant loading at high temperature during continuous operation will result in unavoidable creep damage as well [5]. Consequently, components usually suffered from prior LCF with subsequent creep damage. Therefore, an awareness on the rate of degradation in creep properties due to progress of prior LCF is essential for the readiness against service failures and for remnant life assessment, and the influence of prior LCF on subsequent creep behavior should be clarified to ensure the engineering structures safety [6,7]. The conventional way of carrying out creep-fatigue interaction did throw some light on this problem by introducing a dwell time in strain controlled LCF tests where creep damage is superimposed on fatigue damage in each cycle [8-11]. Our previous work has also investigated the creep-fatigue damage under the cyclic loading with long dwell time [12]. However, the effect of extensive cyclic softening or microstructural
evolution encountered during fatigue stage on creep behavior is still unknown [13]. Therefore a sequential pattern of test with prior LCF followed by creep loading is a better choice. Up to now, there have been limited investigations [13-15] on the creep behavior with prior LCF damage. In general, the prior LCF increases the minimum creep rate and decreases creep lifetime. Furthermore, the damage level of prior LCF is concerned with the temperature, cycle number and strain rate. The increase of temperature, cycle number and strain rate causes a significant decrease in creep strength. This very significant effect affects the usual creep-fatigue damage calculation used in design codes [16]. However, to the best of the authors’ knowledge, the effect of different strain amplitudes prior LCF on subsequent creep behavior has not been clarified so far. Actually, the fatigue damage has been reported to be related to strain amplitude [17]. In our previous study [18], we concluded that the increase of strain amplitude from 0.2% to 0.8% at 650 °C led to the 90% decrease of fatigue lifetime which indicates that the creep behavior after prior LCF indeed has a direct relation with the strain amplitude. Therefore, influence of different strain amplitudes prior LCF on subsequent creep behavior should be elucidated for the readiness against service failures of USC plants. The objective of this paper is to characterize the influence of different strain amplitudes prior LCF on subsequent creep behavior of P92 steel. A comparative evaluation of creep properties of specimens interrupted after same lifetime fraction at different strain amplitudes is carried out to clarify the effect of strain amplitude of prior LCF. Additionally, fractographic feature observation of the creep failed specimens and microstructure characterization of the prior LCF and creep failed specimens are also performed. On the basis of correlation between microstructure evolution and creep rupture behavior, the damage mechanism of different strain amplitudes prior LCF on creep behavior is revealed.
2. Experimental procedures
P92 steel pipe with an outer diameter of 105 mm and a wall thickness of 24 mm was used in this study. The pipe was supplied in the normalized and tempered condition. The chemical compositions of this steel are listed in Table 1. Cylindrical specimens with 25 mm gauge length and 8 mm gauge diameter were machined from the pipe along the axial direction, as shown in Fig. 1. Specimens used in prior LCF tests and in the subsequent creep tests have all the same geometrical dimensions. In order to investigate the effect of different strain amplitudes prior LCF on subsequent creep behavior, experiments were conducted in three steps [19]. Firstly, reference samples were fatigue tested until final fracture at 650 °C under fully reversed triangular waveform (R = -1) in total-strain control mode. Strain amplitudes of 0.25%, 0.4% and 0.6% with constant strain rate of 1.0×10-3 s-1 were adopted. Temperature along the gauge length was controlled within ±2 °C during tests. After fatigue tests, the numbers of failure cycles (N f) of different strain amplitudes were determined [20], as shown in Fig. 2(a). Next, the cycle number (N) was normalized by the failure cycles (Nf) and prior LCF tests on new specimens were conducted at 650 °C to reach predetermined lifetime fractions, as shown in Fig. 2(b). Finally, these fresh prior LCF exposure specimens were creep tested to rupture at 650 °C under 130 MPa constant loading. The 130 MPa constant loading was adopted by considering the creep damage mechanisms [21,22]. Microstructure investigations were carried out by using a Zeiss Axio Imager A1m optical microscope (OM), a JSM-6360LV scanning electron microscope (SEM) and a JEOL JEM-2010 transmission electron microscope (TEM). TEM foils for the investigation of microstructure evolution behavior were prepared from transverse section of the tested specimens. Martensite lath width was measured from the TEM micrographs by the linear intercept method through counting all clearly visible lath boundaries.
3. Experimental results 3.1. Initial microstructure
The microstructure of the as-received P92 steel is shown in Figs. 3(a) and (b) by OM and TEM, respectively. The as-received microstructure is a typical tempered martensite structure. It consists of prior austenite grains, martensite laths, blocks and packets, as shown in Fig. 3(a). Observation of TEM micrograph indicates that some precipitates distribute along martensite lath boundaries. In addition, fine subgrain structure and high density dislocations within martensite laths can also be seen. Mean value of martensite lath width in the as-received condition is about 0.375 μm. It has been reported that the good fatigue and creep properties of P92 steel are mainly attributed to the complex microstructures [23]. This tempered microstructures, in particular subgrain and dislocation structures will undergo low energy configurations to achieve steady state equilibrium structures under prior LCF loading.
3.2. Creep properties after various prior LCF damage at different strain amplitudes
Figure 4 presents creep curves after various prior LCF exposures. The creep curve of the as-received specimen (without prior LCF damage) is also presented for comparison. It can be observed that at different strain amplitudes of 0.25% (Fig. 4(a)), 0.4% (Fig. 4(b)) and 0.6% (Fig. 4(c)), the introduction of prior LCF reduces the creep lifetime. However, no evident change of creep curve shape can be found. The creep deformation still consists of three stages: transient primary stage, obvious secondary stage and tertiary stage. In which the tertiary stage dominates the main deformation. In addition, it is noteworthy that obvious difference can be seen between different strain amplitudes even though the lifetime fraction of prior LCF is identical. To further investigate the effect of strain amplitude of prior LCF on subsequent creep deformation, creep curves after the same lifetime fraction, 20% (Fig. 5(a)) and 50% (Fig. 5(b)), were compared. It is evident that the creep lifetime decreases a lot with the increase of strain amplitude. However, the degradation is saturated irrespective of prior LCF damage when strain amplitude exceeds 0.4%. Moreover, the creep
rupture strain can also be depicted well by the creep strain at the end of the creep curves. As it can be seen, the creep rupture strain was reduced by prior LCF exposure, which indicates the degradation of creep ductility. Additionally, increasing strain amplitude of prior LCF results in a gradually decreased creep rupture strain and the creep rupture strain also shows a tendency to be saturated with further increasing strain amplitude after 0.4%. This is in accordance with the result of creep lifetime that has been achieved before. Creep lifetime versus lifetime fraction and creep lifetime versus strain amplitude are shown in Figs. 6(a) and (b), respectively. It’s evident that prior LCF definitely degrades the subsequent creep strength nonlinearly, regardless of the strain amplitude, as shown in Fig. 6(a). The extent of degradation is strain amplitude and lifetime fraction of prior LCF dependent. For example, comparison between 0.25% and 0.4% strain amplitude reveals that the prior LCF damage in 0.4% strain amplitude is severer. At 20% lifetime fraction of prior LCF, the creep lifetime of 0.25% strain amplitude is about 1085 h, which is higher than the creep lifetime of 0.4% strain amplitude, 865 h. This tendency keeps unchanged until the lifetime fraction comes to 70%. After that, further increase in lifetime fraction doesn’t decrease the creep lifetime anymore even in higher strain amplitudes. This observation is consistent with the limited work reported in [13,24] where the prior LCF damage tends to be saturated after a particular lifetime fraction. Consequently, the damage caused by 70% lifetime fraction is the maximum damage that can be caused by prior LCF. What’s more, 503 h thus can be identified as the shortest creep lifetime of P92 steel at 130 MPa and 650 °C, regardless of the strain amplitude and the lifetime fraction of prior LCF. Of course, this conclusion is valid without obvious outer surface cracks on the specimens after prior LCF. In this investigation, observation of specimen’s outer surface has been performed and will be discussed in Section 3.3. Further observation of Fig. 6(a) shows another interesting phenomenon. The higher strain amplitude of 0.6% at same lifetime fraction doesn’t affect the subsequent creep lifetime compared to 0.4% strain amplitude. This gives an indication that
prior LCF damage is only related to the cycle numbers when strain amplitude is higher than 0.4%. In order to elucidate the effect of strain amplitude in detail, Fig. 6(b) depicts the relationship between strain amplitude and creep lifetime at 20% and 50% prior LCF. It can be achieved that evident reduction of creep lifetime can be induced by low strain amplitude. However, the reduction tends to be saturated at higher strain amplitude. This is consistent with the observations in Fig. 6(a). The microstructure evolution during prior LCF will give an explanation for this behavior and this will be investigated in Section 3.4. Additionally, to demonstrate the effect of prior LCF on different creep stages, Fig. 7 depicts the creep rate curves with different strain amplitudes at 20% and 50% lifetime fraction. It can be seen that 20% and 50% lifetime fraction of prior LCF play a similar role in affecting the creep rate. In both Figs. 7(a) and (b), an obvious creep rate difference can be noticed in the steady and tertiary stages. With increasing strain amplitude of prior LCF, creep rate increases. However, the effect of prior LCF in primary stage seems less. It is also noteworthy that there is no obvious difference in creep rate between 0.4% and 0.6% strain amplitudes, which is consistent with Fig. 5. Since the minimum creep rate is an important parameter describing creep properties, Fig. 8 presents the evolution of the minimum creep rate with the strain amplitude. It reveals that increasing the strain amplitude of prior LCF leads to the increased minimum creep rate. This phenomenon can be found for both 20% and 50% lifetime fractions. Meanwhile, further increase in strain amplitude hardly causes evident increase of the minimum creep rate, which shows a tendency to be saturated even strain amplitude increases from 0.4% to 0.6%. The minimum creep rate evolution is also in agreement with the creep lifetime evolution shown in Fig. 6(b).
3.3. Fracture surface observation
As mentioned before, in order to check whether prior LCF affects the micro-cracks initiation or not during subsequent creep tests, Fig. 9 presents the longitudinal outer surface of the specimens near the creep
fracture location. In pure creep specimen (Fig. 9(a)), the outer surface is quite smooth without any cracks except the oxidation layer due to long-term high temperature exposure. With increasing lifetime fraction of prior LCF (Figs. 9(a)-(c)), the outer surface appears a slight rough tendency. Wang et al. [25] have reported that surface roughness increased with the increase of fatigue cycles. This is consistent with our present observation. Furthermore, comparison of outer surface at different strain amplitudes shows that the prior LCF exposure does not cause obvious surface cracks on 0.25% and 0.4% strain amplitude specimens, as shown in Figs. 9(b) and (c). Whereas, 0.6% strain amplitude specimens reveal apparent surface cracks. It can be easily deduced that the surface cracks can be attributed to the surface damage due to prior LCF. However, our previous observations reveal that prior LCF damage at 0.4% and 0.6% strain amplitudes should be identical due to the same degradation of subsequent creep lifetime. Consequently, the conflict between surface observation and damage investigation seems to appear. Actually, Mariappan et al. [26] have pointed that the change in the surface roughness was due to the microstructure evolution. Thus, the surface cracks in 0.6% strain amplitude can also be attributed to the microstructure evolution during prior LCF. This indicates that the information gathered from the outer surface may not be reliable to assess the prior LCF damage directly. Thereafter, investigation on creep fracture surface morphologies and microstructure evolutions should be further performed to evaluate the prior LCF damage and this would be discussed in the following sections. The fractograph of the failed sample tested under pure creep loading is shown in Fig. 10. Typical features of rough and dimpled fractography can be observed. Lee et al. [27] have reported that transgranular cracking led to the ductile failure. Fractographs of creep failed specimens with different strain amplitudes at 20% and 50% lifetime fractions are shown in Fig. 11 and Fig. 12, respectively. It can be observed that prior LCF doesn’t bring a lot of changes. Typical features of creep fracture are still vividly represented on the
fractographs. However, detailed observations of dimple size and density indicate that prior LCF still produce some effects. As the strain amplitude increases, the dimple becomes more homogeneous and denser, meanwhile, mean sizes of dimple diameter and depth decrease with increasing strain amplitude, as shown in Figs. 11(a)-(c) and Figs. 12(a)-(c). According to the report of Moorthy et al. [28], the change of dimple morphologies observed in Fig. 11 and Fig. 12 is a sign of the degradation of creep ductility, which is also in agreement with the results shown in Fig. 5. To further investigate the changing characteristics of dimples, Fig. 13 shows the variation of dimple diameters at different strain amplitudes. Dimple sizes were measured from a lot of SEM micrographs which are similar to Figs. 10-12 by statistical method and its distribution character have been confirmed to be able to reflect creep strength [29,30]. It can be seen from Fig. 13 that number of large dimples decreases with increasing strain amplitude, meanwhile, number of small dimples increases. This proves the degradation of creep ductility. Additionally, the mean dimple size also presents a tendency to be saturated after 0.4% strain amplitude which may give an explanation for the saturated phenomena as observed in Section 3.2. As observed before, since no visible cracks which may affect the subsequent creep fracture behavior appear during prior LCF, the different creep failure behavior and the difference in creep curves can then be attributed to the microstructure evolution. Mariappan et al. [26] have also reported that microstructure was the main reason that resulted in the difference in properties. Therefore, the microstructure evolution during prior LCF and its effect on the creep deformation should be clarified.
3.4. Microstructure evolution
TEM images of P92 steel after various prior LCF exposures are shown in Fig. 14. All of these images reveal that evident recovery of martensite laths occurs during prior LCF after comparing with the image of the as-received condition. Additionally, growth of subgrains inside the martensite laths can be detected as
well. The growth of subgrains is a result of the elimination of low-angle boundaries which contain subgrain and lath boundaries [31]. Meanwhile, there are some dislocation networks distributing in martensite laths, which is a sign of significant decline of dislocation density in the laths [28]. The growth of subgrains and the appearance of dislocation networks suggest that the unstable martensite laths in the original microstructure were gradually replaced by the grown subgrains along with the decreased dislocations during prior LCF exposure. One microstructural mechanism proposed to explain the P92 steel’s cyclic softening effect is based on the growth of subrains and decrease of dislocations during cyclic deformation [32]. Moreover, some isolated precipitates inside the laths can be detected, which is a strong indication of the elimination of low angle boundaries. Since most of the precipitates mainly distribute along boundaries in as-received P92 steel, as observed in Fig. 3(b), these isolated precipitates indicate the pre-existing boundaries. Normally, coarsening of these precipitates should emerge during high temperature exposure [33,34]. However, there is no appreciable coarsening of precipitates occurring during prior LCF, as the precipitate coarsening is strongly time-dependent. Such observation is in agreement with many others reported in the literature, concerning various tempered martensitic-ferritic steels subjected to cyclic deformation [35,36]. Further comparison of the TEM images between 20% and 50% lifetime fractions of prior LCF at 0.25% strain amplitude (Figs. 14(a) and (b)) and that at 0.4% strain amplitude (Figs. 14(c) and (d)) give an indication that laths and subgrains sizes increase with the increase of lifetime fraction, while dislocation density decreases. On the other hand, effect of strain amplitude on the microstructure evolution can be clarified by comparing Figs. 14(a) with (c) or by comparing Figs. 14(b) with (d). It shows that higher strain amplitude results in more recovery of martensite laths, higher reduction of dislocation density and more obvious growth of subgrains. Meanwhile, higher strain amplitude drives the lath subgrains into equiaxial shape, which has lower energy, as shown in Figs. 14(c) and (d). Additionally, as compared to the microstructure at 0.4% strain
amplitude, the microstructure at 0.25% strain amplitude is relatively heterogeneous. The heterogeneous phenomenon has been revealed to be related to local plastic deformation [37]. To investigate the influence of prior LCF on microstructure evolution during subsequent creep process, TEM micrograph of failed P92 steel under pure creep loading is presented for comparison, as shown in Fig. 15. The microstructure still consists of martensite laths, subgrains and dislocations. However, in comparison with the as-received image (Fig. 3(b)), eliminations of both low-angle and high-angle lath boundaries could be noticed, as well as the coarsening of precipitates. These precipitates have been confirmed to be M 23C6 and MX by energy dispersive spectrometer (EDS) analysis, as shown in Figs. 15(b) and (c). This means the diffusion process which causes precipitates to coarsen occurs during creep process. In addition, the TEM micrograph of P92 steel after pure creep loading also shows the growth of subgrains and decrease of dislocation density due to the elimination of lath boundaries. The growth of subgrains and decrease of dislocation density further resulted in the martensite laths recovery [38]. Therefore, the dislocation process causing martensite laths recovery in combination with the diffusion process causing precipitates to coarsen are the major process during creep loading which have been proved [39]. Figure 16 presents the TEM images of creep failed samples with various lifetime fractions of 0.25% and 0.4% strain amplitude prior LCF. It is noteworthy that during subsequent creep process, the recovery of martensite laths, the agglomeration of precipitates, the growth of subgrains and the recovery of excess dislocations take place on the basis of the microstructure recovery during prior LCF process. The martensite laths recover to wider laths by the absorption of dislocations and the recombination of two boundaries. Meanwhile, precipitates increasingly coarsen can be also observed. Most of these precipitates distribute along the martensite lath boundaries and some distribute within the laths. These precipitates in the laths suggest a trace of previous lath boundaries as well [40]. In addition to this, the martensite laths evolution
after subsequent creep exposure shows a similar tendency with that after prior LCF exposure. The martensite lath width increases with the increase of lifetime fraction and strain amplitude of prior LCF. Moreover, the martensite laths of failed creep samples with prior LCF are larger than that of pure creep loading (Fig. 15), which indicates that the prior LCF causes additional damage to the subsequent creep.
4. Discussion 4.1. Damage mechanism of prior LCF on subsequent creep behavior
The results shown in the above sections reveal that prior LCF causes a significant degradation on subsequent creep properties of P92 steel. The decline of creep lifetime, the decrease of creep rupture strain and the increase of the minimum creep rate all represent the degradation. However, the degradation of creep properties shows a tendency to be saturated at large lifetime fraction and higher strain amplitude of prior LCF. Firstly, it should be noticed that the fatigue strength of P92 steel depends on the stability of microstructures: dislocations, subgrains, martensite laths and precipitates, as shown in Fig. 3(b). These fine microstructure obstacles hinder the dislocation motion and consequently decrease the viscoplatic strain rate during cyclic deformation. However, a cyclic softening effect occurs during LCF exposure because of microstructure evolution. The microstructure observations have shown that the martensite lath recovery, which is the result of subgrains growth and dislocations recovery, takes place during LCF, as shown in Fig. 14. The subgrains growth and the decrease of the dislocation density could explain the cyclic softening. The initial peak stress may be higher due to the reinforcement of the initial fine martensite lath structures’ size. However, as a result of subgrains growth and dislocations recovery during cycling, the strengthening decreases. With the strengthening of (sub)boundaries and dislocations declining, the hindering effect of obstacles to dislocation motion was degraded, therefore, the stress amplitude decreases during LCF tests, as
observed in Fig. 2. Additionally, some isolated precipitates inside the laths with no appreciable coarsening have been observed in the present work. This means the cyclic softening effect is in little relation to precipitates. Such observation is in agreement with some others reported in the literature [41]. Considering the fact that the creep strength of P92 steel also depends on the obstacles hindering the dislocation motion [42], a similar analysis may be carried out for the degeneration of subsequent creep strength. In material containing fine martensite lath structures, the martensite laths and corresponding fine (sub)boundaries play the main role in hindering dislocation motion during creep deformation [43]. Precipitates also play a role in creep resistance by stabilizing the dislocation network and (sub)boundaries, which retard the dislocation motion [44]. During creep loading, the dislocation process causing martensite laths recovery and the diffusion process causing precipitates to agglomerate take place, as shown in Figs. 15 and 16. The dislocation process absorbed extra dislocations and caused martensite laths recovery dominates the main microstructure evolution during creep deformation. Such behavior is observed among 9-12% Cr martensite lath steels [39]. As explained above, the martensite laths recovery (growth of subrains and the decrease of the dislocation density) without any appreciable coarsening of precipitates contributes to the cyclic softening occurrence during prior LCF process. This can be used to give an explanation for the degeneration of subsequent creep strength. In fact, Chaudhuri [45] and Sawada [46] have pointed that the initial microstructure of materials affects the creep primary stage. During prior LCF exposure, as a result of subgrains growth and cellular dislocation network formation driven by cyclic deformation, the martensite laths recovery takes place and reduces the effective barriers of dislocation motion in subsequent creep load [47]. Therefore, the early transition of secondary creep stage occurs, and thus the minimum creep rate increases, as shown in Fig. 7 and Fig. 8. Additionally, Zhang et al. [48] reported that the movement of slip bands driven by cyclic deformation was another important reason caused cyclic softening. During this
process, the slip bands may become carriers to convey the defects, such as vacancies and dislocations, from the grain interiors to the lath boundaries. These boundary defects generated during prior LCF hence facilitate the formation of cavitation (Fig. 11, Fig. 12) and then accelerate the transition of the tertiary stage and decrease the creep rupture strain (Fig. 4, Fig. 5). Moreover, Mariappan [26] has also pointed out that the slip bands lead to the rough surface. Consequently, surfaces after subsequent creep loading show a slight rough trend, as shown in Fig. 9. These phenomena all explain that the recovery of martensite lath structure degrades creep resistance (Figs. 4-8). However, after softening during prior LCF, the remaining strenghtening due to the (sub)boundaries and dislocations is still available. These are not further affected by cycling at particular lifetime fraction and strain amplitude. Therefore, creep strength still remains quite high, and this phenomenon will be discussed in the next section.
4.2. Effect of strain amplitude of prior LCF
In the results section, it has been shown that with increasing the lifetime fraction and strain amplitude of prior LCF, the degradation of creep strength increases and eventually tends to be saturated. The effect of lifetime fraction of prior LCF on creep behavior has been discussed a lot [13,15]. Therefore, the effect of strain amplitude of prior LCF was focused in this part. The creep results have proved that increase of the strain amplitude leads to the decreased creep strength, as shown in Figs. 5-8, especially when strain amplitude increase from 0.25% to 0.4%, the decline is quite obvious. This means that the damage caused by 0.25% strain amplitude prior LCF is much lower than that at 0.4% strain amplitude. As explained above, the growth of subgrains and decrease of dislocation density driven by cyclic deformation should be responsible for the creep strength decline. This may induce a deduction that these obstacles evolution trends at lower strain amplitude differ strongly from higher strain amplitude. It should be noted that although cyclic softening effects at 0.25% and 0.4% strain amplitude show a similar trend, the
extent of cyclic softening is obviously different, as shown in Fig. 2. The cyclic softening at lower strain amplitude is apparently slower than higher strain amplitude. This means the cyclic softening is a function of strain amplitude. Similarly, growth of subgrains and decrease of dislocation density driven by cyclic deformation are also dependent on strain amplitude. At 0.25% strain amplitude, the recovery of martensite laths is heterogeneous, as shown in Fig. 14(a) and (b). Even though most of martensite laths have recovered to low energy microstructure (equiaxed subgrains and cellular dislocation networks), some martensite laths keep unchanged. This is due to that, during prior LCF exposure, subgrains recovery is related to local plastic deformation under each cycle [49]. The local plastic deformation occurred in each cycle at low strain amplitude (0.25%) is quite small. Only favorably oriented grains will be subjected to plastic deformation, while others remain unchanged and thus a rather heterogeneous distribution of martensite laths at 0.25% strain amplitude was observed. On the contrary, because of bigger local plastic deformation at higher strain amplitude, microstructure recovery at 0.4% strain amplitude is therefore much more homogeneous, as shown in Fig. 14(c) and (d). Almost all the martensite laths recover and the subgrains recover to equiaxed subgrains. These evolutions indicate that the martensite laths recovery, concerning dislocation density and subgrain structures, intensifies even further with increase in applied strain amplitude [50]. Consequently, it is expected that strengthening due to (sub)boundaries and dislocations decrease at 0.25% strain amplitude prior LCF is lower than 0.4% strain amplitude. The remaining strengthening due to the (sub)boundaries and dislocations available at 0.25% strain amplitude is therefore more than 0.4% strain amplitude. Moreover, defects transported by slip bands at 0.25% strain amplitude are also obviously fewer than that at 0.4% strain amplitude. Therefore, the cyclic softening differs from one strain amplitude to the other (Fig. 2). Similarly, during subsequent creep exposure, (sub)boundaries’ resistance to creep dislocation motion weakened by 0.25% strain amplitude prior LCF is much lower. And defects accelerating cavitation formation
transported by slip bands at 0.25% strain amplitude are also fewer. These evolutions ultimately incite later attainment of rupture state for 0.25% strain amplitude prior LCF exposure specimens during subsequent creep exposure, as shown in Fig. 5. Fig. 17 further shows the plot of average martensite lath size versus strain amplitude of prior LCF. The martensite lath width was measured from the many TEM micrographs, which are similar to Fig. 3(b) and Figs. 14-16 through the linear intercept method. It is noteworthy that at 0.4% strain amplitude, after prior LCF exposure, the martensite lath structure recovery is very fast. Therefore, the strengthening due to the (sub)boundaries and dislocations is significantly reduced and more obvious cyclic softening effect occurs, as well as subsequent creep lifetimes are significantly degraded. In addition, Fig. 17 also shows that with increasing strain amplitude, the extent of martensite laths recovery during subsequent creep exposure decreases. This proves that the damage caused by high strain amplitude is larger. In addition to this, at 20% lifetime fraction of prior LCF, with the increase of strain amplitude, the martensite lath width exhibits linear behavior while the 50% presents nonlinear behavior. This also means that with increasing lifetime fraction of prior LCF, the prior LCF damage tends to be saturated regardless of strain amplitude which is consistent with the creep lifetime observation in Fig. 6. However, there is no obvious difference between creep properties at 0.4% and 0.6% strain amplitude in experimental results, as shown in Figs. 5-8. The saturated behavior of microstructure evolution at higher strain amplitude could explain the phenomenon. When strain amplitude increases above 0.4%, the transformation of unstable martensite lath structure shows an obvious tendency to be saturated. This indicates that during subsequent creep process, the remaining strengthening due to the (sub)boundaries and dislocations is still available but saturated. These are not further affected by increasing strain amplitude. Therefore, the creep strength and dimple morphologies present the similar saturated behavior. As discussed above, the conclusion that the creep behavior of P92
steel at elevated temperature is influenced significantly by the strain amplitude of prior LCF in a particular range can be achieved.
5. Conclusions In the present investigation, the effect of different strain amplitudes prior LCF on creep behavior was investigated. Microstructure evolutions during prior LCF exposure followed by creep were analyzed to clarify the damage mechanism. The experimental analysis and discussions lead to the following major conclusions: (1) The subsequent creep properties are degraded nonlinearly by the prior LCF. The degradation increases with increase in lifetime fraction and strain amplitude of prior LCF. The prior LCF damage eventually got saturated at 70% lifetime faction, irrespective of strain amplitude. In addition, when strain amplitude exceeds 0.4%, no additional prior LCF damage is induced by the higher strain amplitude. This marks the maximum damage of prior LCF on creep at that lifetime fraction or at that strain amplitude. (2) P92 steel demonstrates a rough and dimpled characteristic under pure creep and prior LCF with subsequent creep loading conditions. Increases of lifetime fraction and strain amplitude of prior LCF contribute to the homogeneous and denser dimples indicating the degradation of creep ductility. (3) Transformation of unstable martensite lath structure to dislocation cell and subgrain structure driven by cyclic deformation during prior LCF process mainly contributes to the degradation of creep properties. Martensite laths evolution behavior is different for different lifetime fractions and different strain amplitudes. The martensite lath structures affect the creep lifetime, the minimum creep rate and the creep rupture strain.
Acknowledgements The authors gratefully acknowledge the financial support of the China Postdoctoral Science Foundation (No. 2016M600405) and innovation program for graduate students in Jiangsu Province of China (No. KYCX17_0935). We also thank Mr. Yi-Nong Lv from School of Material Science and Engineering, Nanjing Tech University, Nanjing, for his help with the TEM specimens’ preparation and observation.
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Fig. 1 Specimen geometry for fatigue followed by creep tests (unit: mm).
Fig. 2. (a) Determination of failure lifetime (Nf) at different strain amplitudes, (b) normalized cycle number at different strain amplitudes
Fig. 3. OM micrograph (a) and TEM micrograph (b) of as-received P92 steel
Fig. 4. Creep curves of P92 steel with various lifetime fraction of prior LCF at strain amplitudes of 0.25% (a), 0.4% (b) and 0.6% (c)
Fig. 5. Creep curves of P92 steel after 20% (a) and 50% (b) lifetime fraction of prior LCF at various strain amplitudes
Fig. 6. (a) Creep lifetime versus lifetime fraction of prior LCF, (b) creep lifetime versus strain amplitude of prior LCF
Fig. 7. Creep rate curves of P92 steel with different strain amplitudes at 20% (a) and 50% (b) lifetime fraction prior LCF.
Fig. 8. The minimum creep rate versus strain amplitude of prior LCF
Fig. 9. Optical micrographs of the outer surface near the fracture location with different lifetime fraction of prior LCF: (a) 0%; (b) 20%; (c) 50%
Fig. 10. Fracture surface of P92 steel subjected to pure creep at 650 °C
Fig. 11. Fracture surface of P92 steel subjected to 20% prior LCF with different strain amplitudes followed by creep loading: (a) 0.25%; (b) 0.4%; (c) 0.6%
Fig. 12. Fracture surface of P92 steel subjected to 50% prior LCF with different strain amplitudes followed by creep loading: (a) 0.25%; (b) 0.4%; (c) 0.6%
Fig. 13. Variation of dimple diameters with different strain amplitudes at 20% (a) and 50% (b) lifetime fraction prior LCF
Fig. 14. TEM micrographs of P92 steel after different prior LCF (a) 20%, 0.25%, (b) 50%, 0.25%, (c) 20%, 0.4%, (d) 50%, 0.4%
Fig. 15. (a) TEM image of P92 steel after pure creep loading, (b) EDS analysis of M 23C6 (point A), (c) EDS analysis of MX (point B)
Fig. 16. TEM micrographs of P92 steel after subsequent creep exposure (a) 20%, 0.25%, (b) 50%, 0.25%, (c) 20%, 0.4%, (d) 50%, 0.4%
Fig. 17. Relationship of martensite lath size versus strain amplitude of prior LCF obtained from measurement of TEM images
Table 1 Chemical composition of the as received P92 steel (wt%). Element
C
Mn
Si
P
S
Cr
Mo
V
N
Ni
Al
Nb
W
B
Amount
0.106
0.361
0.235
0.017
0.0081
9.18
0.368
0.182
0.061
0.108
0.0059
0.078
1.85
0.0022
Highlights (1) Creep properties change with lifetime fraction and strain amplitude of prior LCF. (2) Degradation of subsequent creep strength due to microstructure evolution. (3) Prior LCF damage comes to the maximum after 70% lifetime fraction or after 0.4% strain amplitude, respectively.
S0142-1123(18)30001-X https://doi.org/10.1016/j.ijfatigue.2018.01.001 JIJF 4530
To appear in:
International Journal of Fatigue
Received Date: Revised Date: Accepted Date:
13 September 2017 31 December 2017 2 January 2018
Please cite this article as: Zhang, W., Wang, X., Li, X., Gong, J., Abdel Wahab, M., Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior, International Journal of Fatigue (2018), doi: https://doi.org/10.1016/j.ijfatigue.2018.01.001
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Influence of prior low cycle fatigue on microstructure evolution and subsequent creep behavior Wei Zhang 1,2, Xiaowei Wang 1,2,3,, Xiang Li 1,2, Jianming Gong 1,2, Magd Abdel Wahab 4,5 1
School of Mechanical and Power Engineering, Nanjing Tech University, Nanjing 211816, China
2
Jiangsu Key Lab of Design and Manufacture of Extreme Pressure Equipment, Nanjing 211816, China
3
Faculty of Engineering and Architecture, Ghent University, Zwijnaarde B-9052, Belgium
4
Institute of Research and Development, Duy Tan University, 03 Quang Trung, Da Nang, Vietnam
5
Soete Laboratory, Ghent University, Ghent 9000, Belgium
ABSTRACT This work is devoted to the study of the influence of prior low cycle fatigue (LCF) on microstructure evolution and subsequent creep behavior of P92 steel. Different strain amplitudes prior LCF tests ranging from 0.25% to 0.6% were performed at 650 °C with constant strain rate of 1.0×10-3 s-1 and thereafter subsequent creep tests were conducted at 130 MPa in the same temperature. Optical microscope (OM), scanning electron microscope (SEM) and transmission electron microscope (TEM) were utilized to clarify prior LCF damages of various strain amplitude on subsequent creep behaviors. Results reveal that the increase of lifetime fraction and the increase of strain amplitude lead to the nonlinear degradation of subsequent creep properties. However, the degradation behavior shows a tendency to be saturated after 70% lifetime fraction or after 0.4% strain amplitude, respectively. The degradation is lifetime fraction and strain amplitude of prior LCF dependent. Further analysis of microstructure indicates that the degradation behavior is attributed to the different evolution behavior of martensite laths driven by cyclic deformation during prior
Corresponding author. Tel.: +86 25 58139361; Fax: +86 25 58139361. E-mail address: [email protected] (Xiaowei Wang).
LCF. Additionally, fractographs of failed creep specimens were analyzed in detail to correlate the different subsequent creep behaviors. Key words: Prior low cycle fatigue; Creep behavior; Microstructure
1. Introduction ASME grade P92 steel (with the addition of W and B and a reduction of Mo in P91 steel) is widely used for ultra-supercritical (USC) power plants [1]. Excellent creep strength, lower thermal expansion coefficient, higher thermal conductivity, good weld ability and better corrosion resistance in water steam systems are important features for its suitability in USC industry. It is also considered as favored candidate materials for sodium or lead cooled reactors as well as gas-cooled reactors in innovative Generation IV plants because of its excellent irradiation resistance to void swelling [2,3]. Normally, during service operation of USC plants, as a result of frequent start-ups, shut-downs and compensating of renewable energy fluctuating output, the components often undergo repeated thermal stresses which result in strain controlled low cycle fatigue (LCF) damage of the used materials [4]. Moreover, long-term of constant loading at high temperature during continuous operation will result in unavoidable creep damage as well [5]. Consequently, components usually suffered from prior LCF with subsequent creep damage. Therefore, an awareness on the rate of degradation in creep properties due to progress of prior LCF is essential for the readiness against service failures and for remnant life assessment, and the influence of prior LCF on subsequent creep behavior should be clarified to ensure the engineering structures safety [6,7]. The conventional way of carrying out creep-fatigue interaction did throw some light on this problem by introducing a dwell time in strain controlled LCF tests where creep damage is superimposed on fatigue damage in each cycle [8-11]. Our previous work has also investigated the creep-fatigue damage under the cyclic loading with long dwell time [12]. However, the effect of extensive cyclic softening or microstructural
evolution encountered during fatigue stage on creep behavior is still unknown [13]. Therefore a sequential pattern of test with prior LCF followed by creep loading is a better choice. Up to now, there have been limited investigations [13-15] on the creep behavior with prior LCF damage. In general, the prior LCF increases the minimum creep rate and decreases creep lifetime. Furthermore, the damage level of prior LCF is concerned with the temperature, cycle number and strain rate. The increase of temperature, cycle number and strain rate causes a significant decrease in creep strength. This very significant effect affects the usual creep-fatigue damage calculation used in design codes [16]. However, to the best of the authors’ knowledge, the effect of different strain amplitudes prior LCF on subsequent creep behavior has not been clarified so far. Actually, the fatigue damage has been reported to be related to strain amplitude [17]. In our previous study [18], we concluded that the increase of strain amplitude from 0.2% to 0.8% at 650 °C led to the 90% decrease of fatigue lifetime which indicates that the creep behavior after prior LCF indeed has a direct relation with the strain amplitude. Therefore, influence of different strain amplitudes prior LCF on subsequent creep behavior should be elucidated for the readiness against service failures of USC plants. The objective of this paper is to characterize the influence of different strain amplitudes prior LCF on subsequent creep behavior of P92 steel. A comparative evaluation of creep properties of specimens interrupted after same lifetime fraction at different strain amplitudes is carried out to clarify the effect of strain amplitude of prior LCF. Additionally, fractographic feature observation of the creep failed specimens and microstructure characterization of the prior LCF and creep failed specimens are also performed. On the basis of correlation between microstructure evolution and creep rupture behavior, the damage mechanism of different strain amplitudes prior LCF on creep behavior is revealed.
2. Experimental procedures
P92 steel pipe with an outer diameter of 105 mm and a wall thickness of 24 mm was used in this study. The pipe was supplied in the normalized and tempered condition. The chemical compositions of this steel are listed in Table 1. Cylindrical specimens with 25 mm gauge length and 8 mm gauge diameter were machined from the pipe along the axial direction, as shown in Fig. 1. Specimens used in prior LCF tests and in the subsequent creep tests have all the same geometrical dimensions. In order to investigate the effect of different strain amplitudes prior LCF on subsequent creep behavior, experiments were conducted in three steps [19]. Firstly, reference samples were fatigue tested until final fracture at 650 °C under fully reversed triangular waveform (R = -1) in total-strain control mode. Strain amplitudes of 0.25%, 0.4% and 0.6% with constant strain rate of 1.0×10-3 s-1 were adopted. Temperature along the gauge length was controlled within ±2 °C during tests. After fatigue tests, the numbers of failure cycles (N f) of different strain amplitudes were determined [20], as shown in Fig. 2(a). Next, the cycle number (N) was normalized by the failure cycles (Nf) and prior LCF tests on new specimens were conducted at 650 °C to reach predetermined lifetime fractions, as shown in Fig. 2(b). Finally, these fresh prior LCF exposure specimens were creep tested to rupture at 650 °C under 130 MPa constant loading. The 130 MPa constant loading was adopted by considering the creep damage mechanisms [21,22]. Microstructure investigations were carried out by using a Zeiss Axio Imager A1m optical microscope (OM), a JSM-6360LV scanning electron microscope (SEM) and a JEOL JEM-2010 transmission electron microscope (TEM). TEM foils for the investigation of microstructure evolution behavior were prepared from transverse section of the tested specimens. Martensite lath width was measured from the TEM micrographs by the linear intercept method through counting all clearly visible lath boundaries.
3. Experimental results 3.1. Initial microstructure
The microstructure of the as-received P92 steel is shown in Figs. 3(a) and (b) by OM and TEM, respectively. The as-received microstructure is a typical tempered martensite structure. It consists of prior austenite grains, martensite laths, blocks and packets, as shown in Fig. 3(a). Observation of TEM micrograph indicates that some precipitates distribute along martensite lath boundaries. In addition, fine subgrain structure and high density dislocations within martensite laths can also be seen. Mean value of martensite lath width in the as-received condition is about 0.375 μm. It has been reported that the good fatigue and creep properties of P92 steel are mainly attributed to the complex microstructures [23]. This tempered microstructures, in particular subgrain and dislocation structures will undergo low energy configurations to achieve steady state equilibrium structures under prior LCF loading.
3.2. Creep properties after various prior LCF damage at different strain amplitudes
Figure 4 presents creep curves after various prior LCF exposures. The creep curve of the as-received specimen (without prior LCF damage) is also presented for comparison. It can be observed that at different strain amplitudes of 0.25% (Fig. 4(a)), 0.4% (Fig. 4(b)) and 0.6% (Fig. 4(c)), the introduction of prior LCF reduces the creep lifetime. However, no evident change of creep curve shape can be found. The creep deformation still consists of three stages: transient primary stage, obvious secondary stage and tertiary stage. In which the tertiary stage dominates the main deformation. In addition, it is noteworthy that obvious difference can be seen between different strain amplitudes even though the lifetime fraction of prior LCF is identical. To further investigate the effect of strain amplitude of prior LCF on subsequent creep deformation, creep curves after the same lifetime fraction, 20% (Fig. 5(a)) and 50% (Fig. 5(b)), were compared. It is evident that the creep lifetime decreases a lot with the increase of strain amplitude. However, the degradation is saturated irrespective of prior LCF damage when strain amplitude exceeds 0.4%. Moreover, the creep
rupture strain can also be depicted well by the creep strain at the end of the creep curves. As it can be seen, the creep rupture strain was reduced by prior LCF exposure, which indicates the degradation of creep ductility. Additionally, increasing strain amplitude of prior LCF results in a gradually decreased creep rupture strain and the creep rupture strain also shows a tendency to be saturated with further increasing strain amplitude after 0.4%. This is in accordance with the result of creep lifetime that has been achieved before. Creep lifetime versus lifetime fraction and creep lifetime versus strain amplitude are shown in Figs. 6(a) and (b), respectively. It’s evident that prior LCF definitely degrades the subsequent creep strength nonlinearly, regardless of the strain amplitude, as shown in Fig. 6(a). The extent of degradation is strain amplitude and lifetime fraction of prior LCF dependent. For example, comparison between 0.25% and 0.4% strain amplitude reveals that the prior LCF damage in 0.4% strain amplitude is severer. At 20% lifetime fraction of prior LCF, the creep lifetime of 0.25% strain amplitude is about 1085 h, which is higher than the creep lifetime of 0.4% strain amplitude, 865 h. This tendency keeps unchanged until the lifetime fraction comes to 70%. After that, further increase in lifetime fraction doesn’t decrease the creep lifetime anymore even in higher strain amplitudes. This observation is consistent with the limited work reported in [13,24] where the prior LCF damage tends to be saturated after a particular lifetime fraction. Consequently, the damage caused by 70% lifetime fraction is the maximum damage that can be caused by prior LCF. What’s more, 503 h thus can be identified as the shortest creep lifetime of P92 steel at 130 MPa and 650 °C, regardless of the strain amplitude and the lifetime fraction of prior LCF. Of course, this conclusion is valid without obvious outer surface cracks on the specimens after prior LCF. In this investigation, observation of specimen’s outer surface has been performed and will be discussed in Section 3.3. Further observation of Fig. 6(a) shows another interesting phenomenon. The higher strain amplitude of 0.6% at same lifetime fraction doesn’t affect the subsequent creep lifetime compared to 0.4% strain amplitude. This gives an indication that
prior LCF damage is only related to the cycle numbers when strain amplitude is higher than 0.4%. In order to elucidate the effect of strain amplitude in detail, Fig. 6(b) depicts the relationship between strain amplitude and creep lifetime at 20% and 50% prior LCF. It can be achieved that evident reduction of creep lifetime can be induced by low strain amplitude. However, the reduction tends to be saturated at higher strain amplitude. This is consistent with the observations in Fig. 6(a). The microstructure evolution during prior LCF will give an explanation for this behavior and this will be investigated in Section 3.4. Additionally, to demonstrate the effect of prior LCF on different creep stages, Fig. 7 depicts the creep rate curves with different strain amplitudes at 20% and 50% lifetime fraction. It can be seen that 20% and 50% lifetime fraction of prior LCF play a similar role in affecting the creep rate. In both Figs. 7(a) and (b), an obvious creep rate difference can be noticed in the steady and tertiary stages. With increasing strain amplitude of prior LCF, creep rate increases. However, the effect of prior LCF in primary stage seems less. It is also noteworthy that there is no obvious difference in creep rate between 0.4% and 0.6% strain amplitudes, which is consistent with Fig. 5. Since the minimum creep rate is an important parameter describing creep properties, Fig. 8 presents the evolution of the minimum creep rate with the strain amplitude. It reveals that increasing the strain amplitude of prior LCF leads to the increased minimum creep rate. This phenomenon can be found for both 20% and 50% lifetime fractions. Meanwhile, further increase in strain amplitude hardly causes evident increase of the minimum creep rate, which shows a tendency to be saturated even strain amplitude increases from 0.4% to 0.6%. The minimum creep rate evolution is also in agreement with the creep lifetime evolution shown in Fig. 6(b).
3.3. Fracture surface observation
As mentioned before, in order to check whether prior LCF affects the micro-cracks initiation or not during subsequent creep tests, Fig. 9 presents the longitudinal outer surface of the specimens near the creep
fracture location. In pure creep specimen (Fig. 9(a)), the outer surface is quite smooth without any cracks except the oxidation layer due to long-term high temperature exposure. With increasing lifetime fraction of prior LCF (Figs. 9(a)-(c)), the outer surface appears a slight rough tendency. Wang et al. [25] have reported that surface roughness increased with the increase of fatigue cycles. This is consistent with our present observation. Furthermore, comparison of outer surface at different strain amplitudes shows that the prior LCF exposure does not cause obvious surface cracks on 0.25% and 0.4% strain amplitude specimens, as shown in Figs. 9(b) and (c). Whereas, 0.6% strain amplitude specimens reveal apparent surface cracks. It can be easily deduced that the surface cracks can be attributed to the surface damage due to prior LCF. However, our previous observations reveal that prior LCF damage at 0.4% and 0.6% strain amplitudes should be identical due to the same degradation of subsequent creep lifetime. Consequently, the conflict between surface observation and damage investigation seems to appear. Actually, Mariappan et al. [26] have pointed that the change in the surface roughness was due to the microstructure evolution. Thus, the surface cracks in 0.6% strain amplitude can also be attributed to the microstructure evolution during prior LCF. This indicates that the information gathered from the outer surface may not be reliable to assess the prior LCF damage directly. Thereafter, investigation on creep fracture surface morphologies and microstructure evolutions should be further performed to evaluate the prior LCF damage and this would be discussed in the following sections. The fractograph of the failed sample tested under pure creep loading is shown in Fig. 10. Typical features of rough and dimpled fractography can be observed. Lee et al. [27] have reported that transgranular cracking led to the ductile failure. Fractographs of creep failed specimens with different strain amplitudes at 20% and 50% lifetime fractions are shown in Fig. 11 and Fig. 12, respectively. It can be observed that prior LCF doesn’t bring a lot of changes. Typical features of creep fracture are still vividly represented on the
fractographs. However, detailed observations of dimple size and density indicate that prior LCF still produce some effects. As the strain amplitude increases, the dimple becomes more homogeneous and denser, meanwhile, mean sizes of dimple diameter and depth decrease with increasing strain amplitude, as shown in Figs. 11(a)-(c) and Figs. 12(a)-(c). According to the report of Moorthy et al. [28], the change of dimple morphologies observed in Fig. 11 and Fig. 12 is a sign of the degradation of creep ductility, which is also in agreement with the results shown in Fig. 5. To further investigate the changing characteristics of dimples, Fig. 13 shows the variation of dimple diameters at different strain amplitudes. Dimple sizes were measured from a lot of SEM micrographs which are similar to Figs. 10-12 by statistical method and its distribution character have been confirmed to be able to reflect creep strength [29,30]. It can be seen from Fig. 13 that number of large dimples decreases with increasing strain amplitude, meanwhile, number of small dimples increases. This proves the degradation of creep ductility. Additionally, the mean dimple size also presents a tendency to be saturated after 0.4% strain amplitude which may give an explanation for the saturated phenomena as observed in Section 3.2. As observed before, since no visible cracks which may affect the subsequent creep fracture behavior appear during prior LCF, the different creep failure behavior and the difference in creep curves can then be attributed to the microstructure evolution. Mariappan et al. [26] have also reported that microstructure was the main reason that resulted in the difference in properties. Therefore, the microstructure evolution during prior LCF and its effect on the creep deformation should be clarified.
3.4. Microstructure evolution
TEM images of P92 steel after various prior LCF exposures are shown in Fig. 14. All of these images reveal that evident recovery of martensite laths occurs during prior LCF after comparing with the image of the as-received condition. Additionally, growth of subgrains inside the martensite laths can be detected as
well. The growth of subgrains is a result of the elimination of low-angle boundaries which contain subgrain and lath boundaries [31]. Meanwhile, there are some dislocation networks distributing in martensite laths, which is a sign of significant decline of dislocation density in the laths [28]. The growth of subgrains and the appearance of dislocation networks suggest that the unstable martensite laths in the original microstructure were gradually replaced by the grown subgrains along with the decreased dislocations during prior LCF exposure. One microstructural mechanism proposed to explain the P92 steel’s cyclic softening effect is based on the growth of subrains and decrease of dislocations during cyclic deformation [32]. Moreover, some isolated precipitates inside the laths can be detected, which is a strong indication of the elimination of low angle boundaries. Since most of the precipitates mainly distribute along boundaries in as-received P92 steel, as observed in Fig. 3(b), these isolated precipitates indicate the pre-existing boundaries. Normally, coarsening of these precipitates should emerge during high temperature exposure [33,34]. However, there is no appreciable coarsening of precipitates occurring during prior LCF, as the precipitate coarsening is strongly time-dependent. Such observation is in agreement with many others reported in the literature, concerning various tempered martensitic-ferritic steels subjected to cyclic deformation [35,36]. Further comparison of the TEM images between 20% and 50% lifetime fractions of prior LCF at 0.25% strain amplitude (Figs. 14(a) and (b)) and that at 0.4% strain amplitude (Figs. 14(c) and (d)) give an indication that laths and subgrains sizes increase with the increase of lifetime fraction, while dislocation density decreases. On the other hand, effect of strain amplitude on the microstructure evolution can be clarified by comparing Figs. 14(a) with (c) or by comparing Figs. 14(b) with (d). It shows that higher strain amplitude results in more recovery of martensite laths, higher reduction of dislocation density and more obvious growth of subgrains. Meanwhile, higher strain amplitude drives the lath subgrains into equiaxial shape, which has lower energy, as shown in Figs. 14(c) and (d). Additionally, as compared to the microstructure at 0.4% strain
amplitude, the microstructure at 0.25% strain amplitude is relatively heterogeneous. The heterogeneous phenomenon has been revealed to be related to local plastic deformation [37]. To investigate the influence of prior LCF on microstructure evolution during subsequent creep process, TEM micrograph of failed P92 steel under pure creep loading is presented for comparison, as shown in Fig. 15. The microstructure still consists of martensite laths, subgrains and dislocations. However, in comparison with the as-received image (Fig. 3(b)), eliminations of both low-angle and high-angle lath boundaries could be noticed, as well as the coarsening of precipitates. These precipitates have been confirmed to be M 23C6 and MX by energy dispersive spectrometer (EDS) analysis, as shown in Figs. 15(b) and (c). This means the diffusion process which causes precipitates to coarsen occurs during creep process. In addition, the TEM micrograph of P92 steel after pure creep loading also shows the growth of subgrains and decrease of dislocation density due to the elimination of lath boundaries. The growth of subgrains and decrease of dislocation density further resulted in the martensite laths recovery [38]. Therefore, the dislocation process causing martensite laths recovery in combination with the diffusion process causing precipitates to coarsen are the major process during creep loading which have been proved [39]. Figure 16 presents the TEM images of creep failed samples with various lifetime fractions of 0.25% and 0.4% strain amplitude prior LCF. It is noteworthy that during subsequent creep process, the recovery of martensite laths, the agglomeration of precipitates, the growth of subgrains and the recovery of excess dislocations take place on the basis of the microstructure recovery during prior LCF process. The martensite laths recover to wider laths by the absorption of dislocations and the recombination of two boundaries. Meanwhile, precipitates increasingly coarsen can be also observed. Most of these precipitates distribute along the martensite lath boundaries and some distribute within the laths. These precipitates in the laths suggest a trace of previous lath boundaries as well [40]. In addition to this, the martensite laths evolution
after subsequent creep exposure shows a similar tendency with that after prior LCF exposure. The martensite lath width increases with the increase of lifetime fraction and strain amplitude of prior LCF. Moreover, the martensite laths of failed creep samples with prior LCF are larger than that of pure creep loading (Fig. 15), which indicates that the prior LCF causes additional damage to the subsequent creep.
4. Discussion 4.1. Damage mechanism of prior LCF on subsequent creep behavior
The results shown in the above sections reveal that prior LCF causes a significant degradation on subsequent creep properties of P92 steel. The decline of creep lifetime, the decrease of creep rupture strain and the increase of the minimum creep rate all represent the degradation. However, the degradation of creep properties shows a tendency to be saturated at large lifetime fraction and higher strain amplitude of prior LCF. Firstly, it should be noticed that the fatigue strength of P92 steel depends on the stability of microstructures: dislocations, subgrains, martensite laths and precipitates, as shown in Fig. 3(b). These fine microstructure obstacles hinder the dislocation motion and consequently decrease the viscoplatic strain rate during cyclic deformation. However, a cyclic softening effect occurs during LCF exposure because of microstructure evolution. The microstructure observations have shown that the martensite lath recovery, which is the result of subgrains growth and dislocations recovery, takes place during LCF, as shown in Fig. 14. The subgrains growth and the decrease of the dislocation density could explain the cyclic softening. The initial peak stress may be higher due to the reinforcement of the initial fine martensite lath structures’ size. However, as a result of subgrains growth and dislocations recovery during cycling, the strengthening decreases. With the strengthening of (sub)boundaries and dislocations declining, the hindering effect of obstacles to dislocation motion was degraded, therefore, the stress amplitude decreases during LCF tests, as
observed in Fig. 2. Additionally, some isolated precipitates inside the laths with no appreciable coarsening have been observed in the present work. This means the cyclic softening effect is in little relation to precipitates. Such observation is in agreement with some others reported in the literature [41]. Considering the fact that the creep strength of P92 steel also depends on the obstacles hindering the dislocation motion [42], a similar analysis may be carried out for the degeneration of subsequent creep strength. In material containing fine martensite lath structures, the martensite laths and corresponding fine (sub)boundaries play the main role in hindering dislocation motion during creep deformation [43]. Precipitates also play a role in creep resistance by stabilizing the dislocation network and (sub)boundaries, which retard the dislocation motion [44]. During creep loading, the dislocation process causing martensite laths recovery and the diffusion process causing precipitates to agglomerate take place, as shown in Figs. 15 and 16. The dislocation process absorbed extra dislocations and caused martensite laths recovery dominates the main microstructure evolution during creep deformation. Such behavior is observed among 9-12% Cr martensite lath steels [39]. As explained above, the martensite laths recovery (growth of subrains and the decrease of the dislocation density) without any appreciable coarsening of precipitates contributes to the cyclic softening occurrence during prior LCF process. This can be used to give an explanation for the degeneration of subsequent creep strength. In fact, Chaudhuri [45] and Sawada [46] have pointed that the initial microstructure of materials affects the creep primary stage. During prior LCF exposure, as a result of subgrains growth and cellular dislocation network formation driven by cyclic deformation, the martensite laths recovery takes place and reduces the effective barriers of dislocation motion in subsequent creep load [47]. Therefore, the early transition of secondary creep stage occurs, and thus the minimum creep rate increases, as shown in Fig. 7 and Fig. 8. Additionally, Zhang et al. [48] reported that the movement of slip bands driven by cyclic deformation was another important reason caused cyclic softening. During this
process, the slip bands may become carriers to convey the defects, such as vacancies and dislocations, from the grain interiors to the lath boundaries. These boundary defects generated during prior LCF hence facilitate the formation of cavitation (Fig. 11, Fig. 12) and then accelerate the transition of the tertiary stage and decrease the creep rupture strain (Fig. 4, Fig. 5). Moreover, Mariappan [26] has also pointed out that the slip bands lead to the rough surface. Consequently, surfaces after subsequent creep loading show a slight rough trend, as shown in Fig. 9. These phenomena all explain that the recovery of martensite lath structure degrades creep resistance (Figs. 4-8). However, after softening during prior LCF, the remaining strenghtening due to the (sub)boundaries and dislocations is still available. These are not further affected by cycling at particular lifetime fraction and strain amplitude. Therefore, creep strength still remains quite high, and this phenomenon will be discussed in the next section.
4.2. Effect of strain amplitude of prior LCF
In the results section, it has been shown that with increasing the lifetime fraction and strain amplitude of prior LCF, the degradation of creep strength increases and eventually tends to be saturated. The effect of lifetime fraction of prior LCF on creep behavior has been discussed a lot [13,15]. Therefore, the effect of strain amplitude of prior LCF was focused in this part. The creep results have proved that increase of the strain amplitude leads to the decreased creep strength, as shown in Figs. 5-8, especially when strain amplitude increase from 0.25% to 0.4%, the decline is quite obvious. This means that the damage caused by 0.25% strain amplitude prior LCF is much lower than that at 0.4% strain amplitude. As explained above, the growth of subgrains and decrease of dislocation density driven by cyclic deformation should be responsible for the creep strength decline. This may induce a deduction that these obstacles evolution trends at lower strain amplitude differ strongly from higher strain amplitude. It should be noted that although cyclic softening effects at 0.25% and 0.4% strain amplitude show a similar trend, the
extent of cyclic softening is obviously different, as shown in Fig. 2. The cyclic softening at lower strain amplitude is apparently slower than higher strain amplitude. This means the cyclic softening is a function of strain amplitude. Similarly, growth of subgrains and decrease of dislocation density driven by cyclic deformation are also dependent on strain amplitude. At 0.25% strain amplitude, the recovery of martensite laths is heterogeneous, as shown in Fig. 14(a) and (b). Even though most of martensite laths have recovered to low energy microstructure (equiaxed subgrains and cellular dislocation networks), some martensite laths keep unchanged. This is due to that, during prior LCF exposure, subgrains recovery is related to local plastic deformation under each cycle [49]. The local plastic deformation occurred in each cycle at low strain amplitude (0.25%) is quite small. Only favorably oriented grains will be subjected to plastic deformation, while others remain unchanged and thus a rather heterogeneous distribution of martensite laths at 0.25% strain amplitude was observed. On the contrary, because of bigger local plastic deformation at higher strain amplitude, microstructure recovery at 0.4% strain amplitude is therefore much more homogeneous, as shown in Fig. 14(c) and (d). Almost all the martensite laths recover and the subgrains recover to equiaxed subgrains. These evolutions indicate that the martensite laths recovery, concerning dislocation density and subgrain structures, intensifies even further with increase in applied strain amplitude [50]. Consequently, it is expected that strengthening due to (sub)boundaries and dislocations decrease at 0.25% strain amplitude prior LCF is lower than 0.4% strain amplitude. The remaining strengthening due to the (sub)boundaries and dislocations available at 0.25% strain amplitude is therefore more than 0.4% strain amplitude. Moreover, defects transported by slip bands at 0.25% strain amplitude are also obviously fewer than that at 0.4% strain amplitude. Therefore, the cyclic softening differs from one strain amplitude to the other (Fig. 2). Similarly, during subsequent creep exposure, (sub)boundaries’ resistance to creep dislocation motion weakened by 0.25% strain amplitude prior LCF is much lower. And defects accelerating cavitation formation
transported by slip bands at 0.25% strain amplitude are also fewer. These evolutions ultimately incite later attainment of rupture state for 0.25% strain amplitude prior LCF exposure specimens during subsequent creep exposure, as shown in Fig. 5. Fig. 17 further shows the plot of average martensite lath size versus strain amplitude of prior LCF. The martensite lath width was measured from the many TEM micrographs, which are similar to Fig. 3(b) and Figs. 14-16 through the linear intercept method. It is noteworthy that at 0.4% strain amplitude, after prior LCF exposure, the martensite lath structure recovery is very fast. Therefore, the strengthening due to the (sub)boundaries and dislocations is significantly reduced and more obvious cyclic softening effect occurs, as well as subsequent creep lifetimes are significantly degraded. In addition, Fig. 17 also shows that with increasing strain amplitude, the extent of martensite laths recovery during subsequent creep exposure decreases. This proves that the damage caused by high strain amplitude is larger. In addition to this, at 20% lifetime fraction of prior LCF, with the increase of strain amplitude, the martensite lath width exhibits linear behavior while the 50% presents nonlinear behavior. This also means that with increasing lifetime fraction of prior LCF, the prior LCF damage tends to be saturated regardless of strain amplitude which is consistent with the creep lifetime observation in Fig. 6. However, there is no obvious difference between creep properties at 0.4% and 0.6% strain amplitude in experimental results, as shown in Figs. 5-8. The saturated behavior of microstructure evolution at higher strain amplitude could explain the phenomenon. When strain amplitude increases above 0.4%, the transformation of unstable martensite lath structure shows an obvious tendency to be saturated. This indicates that during subsequent creep process, the remaining strengthening due to the (sub)boundaries and dislocations is still available but saturated. These are not further affected by increasing strain amplitude. Therefore, the creep strength and dimple morphologies present the similar saturated behavior. As discussed above, the conclusion that the creep behavior of P92
steel at elevated temperature is influenced significantly by the strain amplitude of prior LCF in a particular range can be achieved.
5. Conclusions In the present investigation, the effect of different strain amplitudes prior LCF on creep behavior was investigated. Microstructure evolutions during prior LCF exposure followed by creep were analyzed to clarify the damage mechanism. The experimental analysis and discussions lead to the following major conclusions: (1) The subsequent creep properties are degraded nonlinearly by the prior LCF. The degradation increases with increase in lifetime fraction and strain amplitude of prior LCF. The prior LCF damage eventually got saturated at 70% lifetime faction, irrespective of strain amplitude. In addition, when strain amplitude exceeds 0.4%, no additional prior LCF damage is induced by the higher strain amplitude. This marks the maximum damage of prior LCF on creep at that lifetime fraction or at that strain amplitude. (2) P92 steel demonstrates a rough and dimpled characteristic under pure creep and prior LCF with subsequent creep loading conditions. Increases of lifetime fraction and strain amplitude of prior LCF contribute to the homogeneous and denser dimples indicating the degradation of creep ductility. (3) Transformation of unstable martensite lath structure to dislocation cell and subgrain structure driven by cyclic deformation during prior LCF process mainly contributes to the degradation of creep properties. Martensite laths evolution behavior is different for different lifetime fractions and different strain amplitudes. The martensite lath structures affect the creep lifetime, the minimum creep rate and the creep rupture strain.
Acknowledgements The authors gratefully acknowledge the financial support of the China Postdoctoral Science Foundation (No. 2016M600405) and innovation program for graduate students in Jiangsu Province of China (No. KYCX17_0935). We also thank Mr. Yi-Nong Lv from School of Material Science and Engineering, Nanjing Tech University, Nanjing, for his help with the TEM specimens’ preparation and observation.
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Fig. 1 Specimen geometry for fatigue followed by creep tests (unit: mm).
Fig. 2. (a) Determination of failure lifetime (Nf) at different strain amplitudes, (b) normalized cycle number at different strain amplitudes
Fig. 3. OM micrograph (a) and TEM micrograph (b) of as-received P92 steel
Fig. 4. Creep curves of P92 steel with various lifetime fraction of prior LCF at strain amplitudes of 0.25% (a), 0.4% (b) and 0.6% (c)
Fig. 5. Creep curves of P92 steel after 20% (a) and 50% (b) lifetime fraction of prior LCF at various strain amplitudes
Fig. 6. (a) Creep lifetime versus lifetime fraction of prior LCF, (b) creep lifetime versus strain amplitude of prior LCF
Fig. 7. Creep rate curves of P92 steel with different strain amplitudes at 20% (a) and 50% (b) lifetime fraction prior LCF.
Fig. 8. The minimum creep rate versus strain amplitude of prior LCF
Fig. 9. Optical micrographs of the outer surface near the fracture location with different lifetime fraction of prior LCF: (a) 0%; (b) 20%; (c) 50%
Fig. 10. Fracture surface of P92 steel subjected to pure creep at 650 °C
Fig. 11. Fracture surface of P92 steel subjected to 20% prior LCF with different strain amplitudes followed by creep loading: (a) 0.25%; (b) 0.4%; (c) 0.6%
Fig. 12. Fracture surface of P92 steel subjected to 50% prior LCF with different strain amplitudes followed by creep loading: (a) 0.25%; (b) 0.4%; (c) 0.6%
Fig. 13. Variation of dimple diameters with different strain amplitudes at 20% (a) and 50% (b) lifetime fraction prior LCF
Fig. 14. TEM micrographs of P92 steel after different prior LCF (a) 20%, 0.25%, (b) 50%, 0.25%, (c) 20%, 0.4%, (d) 50%, 0.4%
Fig. 15. (a) TEM image of P92 steel after pure creep loading, (b) EDS analysis of M 23C6 (point A), (c) EDS analysis of MX (point B)
Fig. 16. TEM micrographs of P92 steel after subsequent creep exposure (a) 20%, 0.25%, (b) 50%, 0.25%, (c) 20%, 0.4%, (d) 50%, 0.4%
Fig. 17. Relationship of martensite lath size versus strain amplitude of prior LCF obtained from measurement of TEM images
Table 1 Chemical composition of the as received P92 steel (wt%). Element
C
Mn
Si
P
S
Cr
Mo
V
N
Ni
Al
Nb
W
B
Amount
0.106
0.361
0.235
0.017
0.0081
9.18
0.368
0.182
0.061
0.108
0.0059
0.078
1.85
0.0022
Highlights (1) Creep properties change with lifetime fraction and strain amplitude of prior LCF. (2) Degradation of subsequent creep strength due to microstructure evolution. (3) Prior LCF damage comes to the maximum after 70% lifetime fraction or after 0.4% strain amplitude, respectively.