Peculiarities of the Creep Behavior of 15Kh2NMFAA Vessel Steel at High Temperatures

Round 1

Reviewer 1 Report

Comments and Suggestions for Authors

The paper describes the creep behavior of a low alloyed steels commonly used in nuclear plants. The focus is placed on the high temperature range, much higher than the service temperature.

I have some doubts regarding the setup of the paper, which presents an admirably large amount of experimental creep data, but their interpretation isn't entirely clear to me. The parametric Larson-Miller approach allows for the extrapolation of creep performance under conditions where direct testing is not easily performed for various reasons (usually very long times). In this case, it is used to describe experimental results regarding emergency operating conditions, for which creep tests on the order of 100 hours are sufficient. Therefore, in my opinion, the emphasis on the Larson-Miller modeling is misplaced. Furthermore, the importance of grouping experimental data on the basis of the temperature range could be questioned: it shouldn’t be surprising that a specific interpolation offers a more accurate description than a unique interpolation of data at different temperatures.

At the end of the paper, the relevant conclusion is that in the A1-A3 range the creep behavior is peculiar and differs from both “high” and “low” temperature creep. The evidence supporting this conclusion is mainly the Larson-Miller analysis.

In conclusion, I believe that this paper, as it is now, is hardly acceptable as a scientific paper. I usually don’t recommend “reject” – and especially in this case, because the experimental work is really impressive – but I think that it needs a profound revision.

 

In the following, some relatively minor suggestion/comments

Section 2 should give some more details about the experimental setup: what kind of test rig was used (dead weight testing machine, universal testing machines, ...)?, what was the gauge length of the sample? how was the temperature controlled? how was strain measured? Could you please also add some words about the metallographic etching?

 

Row 24: maybe “design” should be replaced with “designed”

Row 41: “at least of 30 years” should be “at least 30 years”, I guess

Row 58: could the authors explain better the statement about the “more than 20 times” difference? Was there over/underestimation? What was the performance parameter taken into account (minimum creep rate, time to rupture, time to specific strain, …)

Row 66: “Cylindrical specimens with a diameter of ∅8 mm for long-term strength tension tests were fabricated of the bottom part of unirradiated WWER pressure vessel”: if I understood correctly, the samples were machined off from an actual unserviced vessel, is it correct? In such case, I would suggest to rephrase “fabricated of the bottom part”

Row 69: “in an argon” should be either “in argon” or “in argon atmosphere”

Row 77: please add some words to suggest that table A1 can be found in the Appendix

Row 82: please consider the use of “minimum creep rate” instead of “minimum creep strain rate”, which is, in my experience, less frequent. Just a suggestion, anyway, the expression used by the authors is correct.

Row 86: there is an extra “it”

Row 90: the coefficient of determination perhaps should be indicated with a capital R

Figure 1, caption: “long term strength curves” could be replaced by “Creep curves”, which is more reader-friendly. I suggest to use this terminology throughout the paper. Moreover, some might argue that 10 hours are not so “long term”

Row 137: “Low-alloyed steels with ferrite-carbide microstructure have relatively high yield strength due to precipitation and dislocation hardening and therefore usually do not require significant strain hardening to resist creep well”. It is hard for me to catch the point the authors are making. I guess this is an explanation on why the samples exhibit virtually no primary stage, but it needs to be further elaborated in my opinion.

Figure 6: The caption (and the text) are somewhat vague. Maybe it’s just me, but I suggest to clearly state in the caption of Figure 6 that the same experimental data in Fig. 2 were modeled grouping ranges of temperatures and calculating temperature-specific coefficients. However, I would raise some concerns about calculating coefficient at 800 °C with only two tests, I don’t think that an optimization is possible.

Author Response

Dear Reviewer.

On behalf of the authors’ team, I thank you for the time you have devoted to a thorough and detailed analysis of our article. We also thank you for giving an honest and competent opinion on the research approach we have proposed and on the article as a whole. Below we offer responses to your comments and remarks.

 

Comment 1

I have some doubts regarding the setup of the paper, which presents an admirably large amount of experimental creep data, but their interpretation isn't entirely clear to me. The parametric Larson-Miller approach allows for the extrapolation of creep performance under conditions where direct testing is not easily performed for various reasons (usually very long times). In this case, it is used to describe experimental results regarding emergency operating conditions, for which creep tests on the order of 100 hours are sufficient. Therefore, in my opinion, the emphasis on the Larson-Miller modeling is misplaced. Furthermore, the importance of grouping experimental data on the basis of the temperature range could be questioned: it shouldn’t be surprising that a specific interpolation offers a more accurate description than a unique interpolation of data at different temperatures

The purpose of fragmentation the Larson-Miller relationship was not the desire to increase the accuracy of the linear approximation of experimental data, but to obtain the ability to more accurately estimate the ultimate long-term strength of the steel under study, taking into account the known structural changes during heating, i.e. this fragmentation has a physical justification. This is important, since the use of the generalized Larson-Miller dependence when predicting the time to failure gives a fairly large error, for example, for 750°C the failure time is overestimated by almost 20 times, which is unacceptable when modeling the residual life of the reactor vessel in emergency operating modes. The fragmentation of the parametric Larson-Miller dependence into intervals does not have scientific significance in itself, but when modeling an emergency situation it allows relying on more accurate data that take into account the microstructure changes in steel.

 

Comment 2

At the end of the paper, the relevant conclusion is that in the A1-A3 range the creep behavior is peculiar and differs from both “high” and “low” temperature creep. The evidence supporting this conclusion is mainly the Larson-Miller analysis.

Although we have mentioned the A1-A3 range several times in “discussion” and “conclusions” sections, it is not entirely clear which conclusion you mean. In this temperature range, the phase composition changes from ferrite to austenite, so the creep characteristics are not entirely clear. On the one hand, the value of the K coefficient in the Monkman-Grant equation in this temperature range (unfortunately, there are data only for the temperature of 800C, which falls within this range) is the highest among all values in the temperature range of 500-1200C (please, see Figure 4). On the other hand, the values of the creep stress exponent n become constant already at temperatures of 750C and higher, i.e. the A1-A3 range entirely falls within this interval of n stability (Figure 3).

We understand that this issue requires more in-depth research, including detailed creep tests in this temperature range and microstructure studies. Considering that the phase composition, and accordingly, the deformation behavior of the steel in the temperature range A1-A3 continuously change depending on the temperature and time of testing, we see this as the goal of a separate, rather labor-intensive study.

 

Comment 3

In conclusion, I believe that this paper, as it is now, is hardly acceptable as a scientific paper. I usually don’t recommend “reject” – and especially in this case, because the experimental work is really impressive – but I think that it needs a profound revision.

Thank you for your honest opinion. We understand it.

When we submitted this article for consideration, we understood that the analytical work we conducted, aimed at obtaining valuable practical knowledge about the behavior of the studied steel at high temperatures, is relevant and has high significance. And the results we obtained are necessary for the correct assessment of the behavior of steel in emergency situations.

Of course, we understand that the explanation of this type of steel behavior in each of the temperature ranges requires a specific explanation, especially from the point of view of the microstructure evolution. We presented some general data on these processes in the paper, including the analysis of the carbide phase. In the revised version of the article, we tried to strengthen this issue as much as possible.

At the same time, given the complex behavior of steel in each of the ranges, a detailed explanation requires a large amount of research. As we mentioned above, for example, the study of steel in the A1-A3 temperature range is definitely a topic for a separate study, since to obtain reliable results it is necessary to analyze test data with a temperature step of at least 10 degrees. And for each test temperature it is necessary to analyze the microstructure components and phase composition. Otherwise, this explanation will not make sense, since the general patterns of deformation of other similar steels are generally known. Therefore, we hope that the approach we propose based on the analysis of experimental data with the receipt of practically significant results and an explanation of the patterns of change in the steel behavior at high temperatures, is acceptable for presenting this information in the form of an article.

 

Comment 4

Section 2 should give some more details about the experimental setup: what kind of test rig was used (dead weight testing machine, universal testing machines, ...)?, what was the gauge length of the sample? how was the temperature controlled? how was strain measured? Could you please also add some words about the metallographic etching?

Thank you. The information about the testing machines was added to the paper.

 

Comment 5

Row 24: maybe “design” should be replaced with “designed”

Corrected. Thank you.

 

Comment 6

Row 41: “at least of 30 years” should be “at least 30 years”, I guess

Corrected. Thank you.

 

Comment 7

Row 58: could the authors explain better the statement about the “more than 20 times” difference? Was there over/underestimation? What was the performance parameter taken into account (minimum creep rate, time to rupture, time to specific strain, …)

The assessment of the failure time obtained on the basis of the Larson-Miller parametric relation generalized for the entire temperature range yields a large error, since it does not take into account the microstructure state of the steel at different temperatures. Thus, for an applied stress of 35 MPa at a temperature of 750°C, the estimated failure time is 3410 hours, whereas during testing, the specimens were failed in 20-21 hours.

The appropriate text has been added to the article.

 

Comment 8

Row 66: “Cylindrical specimens with a diameter of ∅8 mm for long-term strength tension tests were fabricated of the bottom part of unirradiated WWER pressure vessel”: if I understood correctly, the samples were machined off from an actual unserviced vessel, is it correct? In such case, I would suggest to rephrase “fabricated of the bottom part”

Yes, you understood it correctly. We rephrased the sentence in accordance with your proposal.

 

Comment 9

Row 69: “in an argon” should be either “in argon” or “in argon atmosphere”

Corrected. Thank you.

 

Comment 10

Row 77: please add some words to suggest that table A1 can be found in the Appendix

Corrected. Thank you.

 

Comment 11

Row 82: please consider the use of “minimum creep rate” instead of “minimum creep strain rate”, which is, in my experience, less frequent. Just a suggestion, anyway, the expression used by the authors is correct.

Thank you. We have corrected the term according to your recommendation.

 

Comment 12

Row 86: there is an extra “it”

Corrected. Thank you.

 

Comment 13

Row 90: the coefficient of determination perhaps should be indicated with a capital R

You are right, a capital R is better suited to denote the coefficient of determination, but in this paper the gas constant is already denoted as R (please, see the equation (3)). Therefore, I propose to leave the notation "r", or remove the notation of the coefficient of determination, since it actually does not appear anywhere else in the article.

 

Comment 14

Figure 1, caption: “long term strength curves” could be replaced by “Creep curves”, which is more reader-friendly. I suggest to use this terminology throughout the paper. Moreover, some might argue that 10 hours are not so “long term”

Thank you. Here we have also corrected these terms in the paper.

 

Comment 15

Row 137: “Low-alloyed steels with ferrite-carbide microstructure have relatively high yield strength due to precipitation and dislocation hardening and therefore usually do not require significant strain hardening to resist creep well”. It is hard for me to catch the point the authors are making. I guess this is an explanation on why the samples exhibit virtually no primary stage, but it needs to be further elaborated in my opinion.

Yes, you are right; we meant that the creep curves quickly pass to the second stage. In fact, this is the phrase we started our analysis of the experimental data with to show the absence of a clearly defined primary region in the creep curves (Figure 1). Perhaps we should simply delete the first paragraph of the "discussion" section?

 

Comment 16

Figure 6: The caption (and the text) are somewhat vague. Maybe it’s just me, but I suggest to clearly state in the caption of Figure 6 that the same experimental data in Fig. 2 were modeled grouping ranges of temperatures and calculating temperature-specific coefficients. However, I would raise some concerns about calculating coefficient at 800 °C with only two tests, I don’t think that an optimization is possible.

 

Thanks for the comment. The following text was added to the paper:

Figure 6 illustrates “σ - P_LM” dependence for the specified temperature ranges. When constructing Figure 6, data from the same creep tests were used as for constructing the generalized Larson-Miller dependence shown in Figure 2.

As for the value of the coefficient at 800°C, it is built on the results of four points, just two pairs of points were located quite close to each other. The designation of these points was changed for better visualization (please, see new Figure 6).

Reviewer 2 Report

Comments and Suggestions for Authors

the authors have expressed a significant knowledge and experience in creep modelling and application of coefficients 

microstructure characterisation and the discussion based on the effects of microstructure on creep can be improved

please see the pdf file for the detailed comments

Comments for author File: Comments.pdf

Author Response

Dear Reviewer!

Thank you for your careful review of our article.

Most of your comments are related to the fact that the article does not sufficiently cover the microstructure of the steel under study. We agree with almost all of your comments and have tried to answer each of them.

At the same time, we ask you to look at this article from another point of view. The main work carried out in the article is aimed at obtaining and analyzing real data on the deformation behavior of 15Kh2NMFAA steel under creep conditions at temperatures corresponding to abnormal operating conditions of the reactor vessel. These data are of great value for assessing the reliability of power equipment at nuclear power plants. We have conducted a detailed analysis of the experimental data and offer a number of results, indicated in the "Conclusions" section, useful for specialists in assessing the reliability and durability of power engineering equipment.

These results will certainly be even more valuable if we substantiate the deformation behavior of steel from the point of view of microstructure. We agree that it is the microstructure that can explain the behavior of steel and the resulting dependencies. At the same time, the volume of such studies required to obtain reliable results in each temperature range requires a huge amount of research. Therefore, in this paper we tried to give some explanation taking into account our own studies (as much as we can) and the general laws of microstructure evolution for this type of steels.

Our responses to your comments are presented below.

 

Comment 1

Paper fragment:

It was shown that the results of predicted rupture time for 700°C obtained using the Larson-Miller equation for temperatures between 500°C and 900°C without taking into consideration microstructure changes of the 15Kh2NMFAA steel differs from ones obtained by creep tensile test in more than 20 times [36]. It is obvious that the connection between the microstructure and the creep behavior has not been studied in enough detail for this steel yet.

Reviewer comment:

It's clear that the authors did not study the effect of microstructure previously, but it does not mean that nobody did it. A more detailed literature review on the effect of microstructure is required, which will support novelty of this paper. It may occur that this literature review will go beyond the studied steel composition, but still may be beneficial for the community.

Authors’ response:

A large number of papers are devoted to the study of the microstructure influence on steels creep. Herewith, most of the papers are naturally devoted to temperature ranges close to the operating temperature. Indeed, there are papers that study the influence of microstructure on the creep characteristics of reactor steels, for example, SA508 steel. However, the A1 temperature is much lower for SA508 steel, so for this steel, the transition to hot deformation coincides with the onset of phase transformations. The behavior of 15Kh2NMFAA steel, which contains a larger number of ferritizing elements and has significantly higher A1 and A3 temperatures, is different. The transition to hot deformation is observed even in the ferrite-carbide state, at a temperature of about 700C, when we observed a kink in the Larson-Miller parametric dependence.

We made some edits to the manuscript taking into account your comments.

 

Comment 2

Paper fragment:

Cylindrical specimens with a diameter of ∅8 mm for long-term strength tension tests were fabricated of the bottom part of unirradiated WWER pressure vessel. Long-term strength tests were carried out on tensile machines at temperatures of 500-1200°C.

Reviewer comment:

The machine type should be presented

Authors’ response:

Thank you. The information was added to the article.

 

Comment 3

Paper fragment:

At temperatures up to 650°C, tests were carried out in an air atmosphere, at temperatures of 700-900°C - in an argon, at temperatures of 900-1200°C - in a vacuum.

Reviewer comment:

If a special machine with a temperature chamber was used, this should be clearly presented, because the machine design will influence the accuracy of the results

Authors’ response:

Thank you. Some information about testing machine was added to the article.

 

Comment 4

Paper fragment: Figure 1.

Reviewer comment:

It's clear that a higher stress - faster failure, but it would be beneficial to use different line types for the same temperature / colour (dash maybe)

Authors’ response:

Thank you. Corrected.

 

Comment 5

Paper fragment:

A plot of stress σ vs the Larson-Miller parameter P_LM for all tested specimens is shown in Figure 2, taking a value of C = 20. A perceptible slope change of “σ – P_LM” relation in the range of 700-750°C occurs, which could not be related to the steel microstructure type change, because the starting phase transition temperature A₁ is noticeably higher (A₁ ≈ 760-770°C) for this steel [39].

Reviewer comment:

The phase transformation temperature may vary with stress - a literature reference is needed for this.

Authors’ response:

We completely agree with your statement. However, it should be taken into account that the heating of the samples was carried out without loading, and the tests started after the temperature had stabilized, which took several hours. That is, at the tests start moment, the phase composition of the steel was in any case close to equilibrium.

We also admit that during the tests, the ratio of ferrite and austenite could change under the influence of stresses, but this requires separate studies.

 

Comment 6

Paper fragment:

The most possible cause of this phenomenon consists in microstructure changes that are taking place in the region of A1 and A3 temperatures.

Reviewer comment:

Different carbide dissolution temperature for various carbide compositions and dislocation annihilation rate which also varies with temperature.

In general, it's not good to discuss properties before the initial microstructure was presented - was it studied? Initial microstructure should be studied and presented before the properties.

Authors’ response:

We completely agree with you. In this paper, the dynamics of carbide coagulation were not studied in detail; this may be the subject of a separate study. In addition, the purpose of the paper is to study the features of high-temperature creep, and the description of the steel microstructure is given only to explain the features of steel behavior in different temperature ranges. The original microstructure of the steel is represented by a ferrite-carbide mixture obtained after normalization and subsequent high tempering. The results of the microstructure study were given in the authors' previous paper [39]. That is why we decided not to include photographs of the microstructure in this article, so as not to clutter the article.

A brief microstructure description was included in the article.

 

Comment 7

Paper fragment:

It is obvious from the obtained graphs that the microstructure has a decisive impact on the steel behavior under creep conditions. Dispersed particles, primarily carbides, play an important role in resistance to deformation under creep conditions. Distribution of carbides inclusions was studied by scanning electron microscope (Figure 5).

Reviewer comment:

No. What is obvious is the variation of creep behaviour with temperature, but the data on microstructure variation with temperature is insufficient to discuss the effect of microstructure on creep.

Carbides were observed, however chemistry needs a proof (EDS for example) and the size distribution variation with temperature is needed to support the effect of carbides on creep.

Authors’ response:

You are right; this phrase can be considered not as a statement, but rather as an assumption.

Unfortunately, we do not have enough data on the change in size distribution depending on temperature to present a detailed picture. Therefore, we also relied on well-known literary references.

We also present you some of the SEM results we have (please, see the attached file). The SEM results at temperatures of 650, 800 and 850C show the dynamics of coagulation and then dissolution of carbides. We can add this information to the paper, but I am not sure about it. Because, in order to draw conclusions that can be confident, such studies need to be carried out on a large number of samples after tests with different temperatures and creep test times. Anyway, this is a large amount of research.

 

Comment 8

Paper fragment:

The size of vanadium-enriched carbides found in the specimen tested at 650C was smaller and the volume fraction of carbides was less in comparison with the specimens tested at 800C and 850C.

Reviewer comment:

A table with sizes and number densities for all carbide chemistries and for all tested temperatures should be presented

Authors’ response:

We agree that these data would be very interesting. However, in this paper we limited ourselves to a quantitative assessment, since the general patterns of carbide coagulation were studied before us and we consider this sufficient to explain the change in creep characteristics.

 

Comment 9

Paper fragment:

The distributed VC particles critical size to remain coherence is approximately 33-50 nm. [41]. Carbides enriched with vanadium, found in specimens after testing at temperatures of 700C and above, can be attributed to relatively large particles that are no longer crucial for deformation resistance under creep conditions.

Reviewer comment:

Coherency is a big story, it's better not to touch it here, because:

1 - there is no TEM, coherency disappears at smaller than 33 particles sizes; and

2 - there is no proof in this paper that coherency may influence the dislocation and grain boundary pinning efficiency

What particle sizes are critical ? What is the proof that whatever size is critical ?

Authors’ response:

We completely agree with you. We tried to get around this point and removed this phrase from the article.

 

Comment 10

Paper fragment:

This fact is indirectly supported by an abrupt change in value of the creep stress exponent n for the temperature range of 650-700°C (Figure 4). A high value of n > 8 is observed at test temperatures below 650°C, which is typical for dispersion-hardened alloys in the “cold” deformation temperature region, when dislocations are not able to redistribute and are forced to move through obstacles in the form of carbides. At 700C, value of the exponent n is 5.35, corresponding to an intermediate state between “warm” deformation (low temperature dislocation climb controlled by pipe diffusion dominates) and “hot” deformation (high temperature dislocation climb controlled by lattice diffusion dominates).

Reviewer comment:

Good, but reference are needed regarding the variation in the dislocation behaviour with temperature

Authors’ response:

Thank you. The reference was added.

 

Comment 11

Paper fragment:

An important result of this paper is that at temperature range of A₃-1200C the creep parameters remain stable and predictable due to the microstructure stability (Figures 3 and 4).

Reviewer comment:

Show the carbide dissolution temperatures in the studied steel

Authors’ response:

We believe that this sentence should be reformulated as "An important result of this paper is that at temperature range of A₃-1200°C the creep parameters remain stable and predictable", removing the part about "due to the microstructure stability", since we cannot unambiguously state the microstructure stability in this temperature range. The microstructure does change, carbides dissolve, and grain growth is observed. However, this does not affect several specified creep characteristics (n, B, K), as presented at figures 3 and 4.

 

Comment 12

Paper fragment:

As a result, we obtained a significant practical conclusion, which is that for this steel in the range of 850-1200°C we can estimate the time to rupture by the minimum creep strain rate, even without knowing the value of the actual temperature and its change.

Reviewer comment:

Even the microstructure is not needed, just test a specimen and measure the time to failure

Authors’ response:

No. Now we don't even need a specimen and we don't need a test.

Here we are talking about the fact that in real conditions, when emergency situations arise, we only need data from the displacement or strain sensors to estimate the rupture of this steel, even if a temperature fluctuation in the specified range will be occurred.

Author Response File: Author Response.pdf

Round 2

Reviewer 1 Report

Comments and Suggestions for Authors

I still think that Larson-Miller analysis for such emergency operating conditions is not an appropriate tools, since long term data extrapolation is meaningful if the operating conditions are meant to last for long times, which - hopefully - is not the case here. I'd like to point out to the authors that I don't really see the lack of deep microstructural investigations as a weakness of the paper: I know it is useful, I know that scientifically speaking it is highly relevant, but I also know that designers need to know numbers.

However, the explanation given in the reply and the corrections in the paper made me reconsider my opinion and I suggest to publish this work.

Just a final consideration about the accuracy in temperature measurements. Unfortunately right now I can't check the standards, but I suspect that the indicated accuracy is lower than the prescribed values. On the other hand, results are consistent and I guess that the actual temperature control was better than +/-5 to +/-12

Author Response

Dear reviewer, thank you for your opinion and comments.

I still think that Larson-Miller analysis for such emergency operating conditions is not an appropriate tools, since long term data extrapolation is meaningful if the operating conditions are meant to last for long times, which - hopefully - is not the case here. I'd like to point out to the authors that I don't really see the lack of deep microstructural investigations as a weakness of the paper: I know it is useful, I know that scientifically speaking it is highly relevant, but I also know that designers need to know numbers. However, the explanation given in the reply and the corrections in the paper made me reconsider my opinion and I suggest to publish this work.

Just a final consideration about the accuracy in temperature measurements. Unfortunately right now I can't check the standards, but I suspect that the indicated accuracy is lower than the prescribed values. On the other hand, results are consistent and I guess that the actual temperature control was better than +/-5 to +/-12.

As for the accuracy of temperature measurements, you are right. We took the data on temperature fluctuations of +/-5 to +/-12 from the documentation for the testing machine. I think the equipment manufacturers played it safe and indicated the measurement error values ​​with some reserve. The actual temperature fluctuations during testing were usually significantly smaller. According to the requirements of our regulatory documents, the permissible temperature fluctuations are +/-3 at temperatures up to 600C, +/-4 at temperatures of 600...900C, and +/-6 at temperatures of 900...1200C. I believe that in fact we fit into this range. At the same time, in order not to provide unverified information, we propose to exclude the line with the accuracy of temperature maintenance, so as not to raise questions from readers.

____

I would like to express my gratitude to you once again for your competent opinion on our work and the time you devoted to it. Thank you.

Reviewer 2 Report

Comments and Suggestions for Authors

good answer to review

Author Response

Dear reviewer! On behalf of the authors team, I would like to thank you for your understanding, for working on our article, and for your valuable comments on the study of the relationship between microstructure and the patterns of mechanical properties change at high temperatures.

Thank you!