, Ostrava, Czech Republic HOW TO IMPROVE CREEP RUPTURE STRENGTH OF MICROALLOYED STEELS

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1 HOW TO IMPROVE CREEP RUPTURE STRENGTH OF MICROALLOYED STEELS Václav Foldyna a), Zdeněk Kuboň b), Tomáš Schellong a), Zdeněk Kübel a) a) JINPO PLUS a.s., Research and Development, Křišťanova 1113/2, Ostrava Přívoz, Czech Republic b) VÍTKOVICE a.s., Research and Development, Pohraniční 31, Ostrava Vítkovice, Czech Republic Abstract The microalloyed steels are delivered in the normalized condition or in the as rolled condition after normalizing forming. Precipitation strengthening of these is affected by means of small VCN or NbCN particles. After normalizing forming the mean particle diameter is 5 to 12 mm and the mean interparticle spacing IPS is smaller than about 100 nm. After following normalizing the mean IPS increases and therefore proof stress and creep rupture strength (CRS) decreases. Proof stress decreases after normalizing in vanadium bearing steel of about 3% and in Nb bearing steel of about 14%. CRS decreases after normalizing more significantly, in vanadium bearing steel of about 30% and in Nb bearing steel of about 40%. Only slight drop of proof stress and more significant drop of CRS were observed in both investigated steels. Taking into account, that the proof stress is indirect proportional to IPS, while time to rupture is indirect proportional to IPS cubed. With increasing creep rate decreases time to rupture and CRS. 1. Introduction Microalloyed steels with specified elevated temperatures have been developed for piping of power generating and chemical plant equipment where service temperature do not exceed 500 C. These steels are delivered in the normalized condition or in the as rolled condition after normalizing forming. It is supposed that the properties of steels in both mentioned conditions are the same or very similar. The aim of this paper is to show the way how to increase creep resistance of microalloyed steels. The effect of microstructure on the creep properties will be discussed too. 2. EXPERIMENTAL MATERIAL Steel grade is microalloyed with vanadium from 0.05 to 0.09 mass.% or niobium from 0.02 to 0.06 mass.%, carbon and manganese contents are in the range 0.14 to 0.20 mass.% and 0.6 to 1.0 mass.%, respectively. The creep rupture strength was investigated on three heats. Two of them were microalloyed with niobium (heats A and C), heat B was alloyed with vanadium. The chemical composition of these heats is shown in Table

2 Table 1. Chemical composition of the studied heats, (mass.%) Heat C Mn Si P S Ni Cu V Nb Al Note A V B V C n.a. n.a NH Notes: n.a. not analysed V produced in VÍTKOVICE, a.s., Ostrava NH- produced in NOVÁ HUŤ, a. s., Ostrava All heats were processed into seamless tubes of the wall thickness 20 mm and the diameter of 324 mm (heats A and B) and 219 mm (heat C), respectively. The rolling procedure was in all cases temperature-controlled [1, 2, 3]. Table 2 lists the room temperature mechanical properties and shows that normalizing after temperature-controlled rolling slightly impaired proof stress as well as tensile strength of the steels while improving the elongation. Table 2. Mechanical properties at 20 C of the investigated heats R p0.2 R m A 5 Heat MPa % Note A , 3 B , , 3 C , , , , 5 Notes: 1 controlled rolling 2 after normalizing 3 tube ø 324 x 20 mm 4 tube ø 219 x 20 mm 5 pipe ø 31.8 x 2.6 mm Normalizing after controlled rolling significantly decreases proof stress at room temperature in niobium bearing steel, at about 14% (heat C), while in vanadium bearing steel the drop in R p0.2 is only about 3% (heat B). How the proof stress at room temperature depends on the wall thickness and the effect of normalizing is shown in Fig CREEP RUPTURE STRENGTH ESTIMATION For creep rupture tests have been used material from tubes with the wall thickness 20 mm in the as rolled state (i.e. after controlled rolling) and after normalizing at 930 C with cooling in still air. All three heats were creep tested. Creep test to rupture were performed at temperatures 450, 475, 500 and 525 C and in the stress range from 40 to 160MPa, when 3 creep tests have attained more than hours and the longest test duration - 2 -

3 at present time is hours. All creep tests were performed in Research Institute of VÍTKOVICE. Much effort has been devoted to find the correct method for reliable creep rupture strength estimation [4]. Larson Miller parametric equation is most frequently used for evolution and prediction of long term creep rupture properties: P LM = T(C LM + log t r ) (1) where P LM is Larson Miller parameter and t r is time to rupture. The constant C LM can be either elected (ranging usually from 20 to 40) or calculated. Use of calculated C LM constant can be recommended. The stress dependence of P LM can be also described by a polynomial in the form: P LM =a 0 + a 1. log(σ)+a 2. log 2 (σ)+a 3. log 3 (σ)+ + a k. log k (σ) (2) where a 0, a 1, a 2,,a k are regression constants and k is elected up to 5. The performed creep tests were evaluated by means of P LM equation (2), using constant k=1, 2 or 3. The resultant creep rupture strength values are presented in Table 3. There is assumed that higher correlation index I proves higher reliability off the performed estimation. The example of logarithmic plots of time to rupture as a function of stress is shown in Figs. 2, 3, 4. Comparison of creep rupture strength (CRS) in 10 5 hours at 450 and 500 C is shown in Figure 5. Creep rupture strength at 450 C of vanadium bearing heat B is in as rolled condition distinguished higher than that of niobium bearing heats A and C. The creep rupture strength in the as rolled condition of all mentioned heats are at 500 C nearly the same (78 ± 3 MPa). Vanadium bearing steel B attained higher creep rupture strength than niobium bearing steel C also in normalized condition at 500 C (Fig. 5). Significant different temperature dependence of CRS in 10 5 hours was observed in vanadium and niobium bearing steels in as rolled condition as well as after normalizing (Fig. 6). Table 3. Creep rupture strength of the investigated heats Creep rupture strength in 10 5 h, MPa Correlation Heat 450 C 475 C 500 C 525 C k Index I Note A CR B CR N C Cr N Notes: CR after controlled rolling N normalized at 930 C and air cooled Normalizing after controlled rolling decreases proof stress at room temperature as well as creep rupture strength and this effect is more pronounced in niobium bearing steel (up to about 40% - heat C) than in vanadium bearing steel (up to about 30% - heat B) - 3 -

4 see Table 3. The creep rupture strength in 10 5 h of vanadium bearing steel (heat B) is superior at temperatures below 500 C. 4. Discussion How the proof stress and creep rupture strength of some carbon, microalloyed and low alloy steel depends on temperature is shown in Fig. 6. Creep rupture strength data attained for heats A1, B1 and C1 are presented in the condition after controlled rolling. Creep rupture strength for all mentioned steel grades is in agreement with corresponding standards [7, 8]. Comparison of CRS steel grade 20MnNb6 by means of data stated in corresponding standards [7, 8] and CRS of heats B and C attained after normalizing is shown in Fig. 7. It is clear, that CRS values stated in valid standards [7, 8] have been attained after normalizing. Creep rupture strength of examined heats in as rolled condition is much higher (Fig. 5). The low alloy steel 16Mo3 is alloyed with molybdenum from 0.25 to 0.35 mass.%. Using data stated in [7, 8] can be concluded that proof stress or CRS of both mentioned microalloyed steels is higher then that of alloy steel 16Mo3 at temperatures below 410 C (Fig. 6). Using data of heats A1 and C1 in as rolled condition can be expected that niobium microallyed steels are more convenient at temperatures below 440 C. In the case of vanadium microallyed steels in the as rolled condition heat B1; proof stress or CRS is higher than in the case of low alloy steel 16Mo3 at temperature below 470 C. The precipitation strengthening of investigated steels is affected by small particles of vanadium or niobium carbides. The mean particle diameter in the initial state ranges between 5 and 12 nm and the volume fraction between 0.05 and 0.25%, respectively [1]. It was found that the interparticle spacing in the initial state of examined steels is roughly the same as in low alloy steels of 0.5%Cr-0.25%V and 0.5%Cr-0.5%Mo-0.25%V [2], although the volume fraction of precipitates in the investigated steels is smaller. This is due to very small mean diameter of carbide particles in microalloyed steels. This conclusion is confirmed by the creep strength of examined steel after controlled rolling, which at temperatures up to about 525 C is nearly equalled those of low alloy 16Mo3 steel. The examined steels after normalizing exhibited larger MX particles and hence, given a constant volume fraction a correspondingly greater interparticle spacing is attained. This is, of course, reflected in a decline of its creep strength. The observed difference in decreasing creep rupture strength after normalizing between vanadium bearing and niobium bearing steels can be explained by the different solution temperature of vanadium carbide and niobium carbide. During normalizing vanadium carbides partly dissolve and re-precipitate again as a very fine particles during cooling. On the other hand, niobium carbides dissolve at much higher temperature and therefore during normalizing originally small particles only coarsen. As a result, the interparticle spacing of secondary carbides in vanadium bearing steel after normalizing is lower than that in niobium bearing steel. It is well known that, the creep rate in precipitation strengthened steels and alloys is proportional to IPS cubed [10]. As creep rate is indirect proportional to time to rupture, with increasing IPS increases creep rate, decreases time to rupture and creep rupture strength. When assessing the creep properties of microalloyed steels we must take into account not only the interparticle spacing in the as received condition, but also the changes of IPS during creep exposure. Coarsening of particles leads to increasing of mean diameter, decreasing number of particles and increasing IPS

5 Given identical coarsening rate of particles, IPS after creep exposure increases with decreasing mean diameter and volume fraction of particles ( see Figs. 8, 9 ) [9]. Given identical mean diameter of particles, IPS decreases with increasing volume fraction of particles, even in the as received condition ( Fig. 9 ). On the other hand, ratio l 5 10 /l 0 does not depend on the volume fraction of particles ( Fig. 8 ). As microalloyed steels are distinguished for small diameter and small volume fraction of particles, respectively, it is no point in increasing working temperature over 500 C. Especially in the case of vanadium bearing heat B the creep rupture strength decreases very quickly with increasing working temperature. 5. Conclusion With respect to attained experimental results and performed discussion, conclusion may be summarized as follows: Creep rupture strength (CRS) of microalloyed steels is higher in the as rolled condition (after normalizing forming) as after normalizing CRS of vanadium bearing steel is higher than that of niobium bearing steel in the as rolled condition as well as after normalizing. The allowable stress (which depends on the proof stress at elevated temperatures or CRS), of vanadium ( niobium ) bearing steel in as rolled condition is at temperatures up to about 470 C ( 430 C ) higher than that of low alloy molybdenum steel e.g. 16Mo3. CRS data introduced in the present standards corresponds to CRS attained after normalizing. CRS of microalloyed steels in as rolled condition is at about 500 C is quite comparable with CRS of molybdenum bearing low alloy steel e.g. 16Mo

6 REFERENCES [1] PURMENSKÝ, J., JAKOBOVÁ, A., FOLDYNA, V., BEMBENEK, Z., Microstructure and creep properties of carbon boiler steel microalloyed with niobium or vanadium, Proc. Symposium on High-Temperature Metallic Materials, Znojmo, Czech Republic, 1986, p. 97 [2] PURMENSKÝ, J, FOLDYNA, V., Microalloyed steel for chemical industry and power plant equipment, Seminar on New Application in Steel in view of the Challenge from Substitute Materials, R. 37, United Nations Economic Commission for Europe, Luxembourg, May 1988 [3] MELECKÝ, J., MAZANCOVÁ, E., OTTA, J., Hutník, 1988 (38), No. 8-9, p. 301 in Czech [4] FOLDYNA, V., JAKOBOVÁ, A., KUBOŇ, Z., Assessment of creep resistance of 9-12%Cr steels with respect to strengthening and degradation processes, Proc. of the International Symposium on Material Aging and Component Life Extension, Milan, Italy, October 1995, p. 15 [5] FOLDYNA, V., PURMENSKÝ, J., Czechoslovak Journal for Physics, 1989, vol. B39, p [6] FOLDYNA, V., PURMENSKÝ, J., KUBOŇ, Z., Use of microalloyed steels for construction of power plants, their reconstruction and modernisation, Proceedings Boiler and Boiler Equipment 2000, Brno, March 2000, p. 199, in Czech [7] Draft of Czechoslovak norm ČSN , November 1983 [8] European Standard, Draft pr. EN , December 1998 [9] Foldyna, V. et al.: Proc. 2 nd Inter. Symposium on Metallurgy and Materials Science, RISO, Denmark, 1981, p. 271 [10] Sherby, O.D., Klundt, R., Miller, A.K.: Metall. Trans. 8 A 1977, p

7 after controlled rolling after normalizing 442 Proof stress, MPa A1 B1 C2 C3 Heat Fig. 1: The effect of wall thickness and normalizing annealing on the proof stress at room temperature (1 pipe ø324x20 mm; 2 - pipe ø324x20 mm, 3 - tube ø31.8x2.6 mm) 1E+6 1E+5 1E+4 Time, h 1E C 475 C 500 C 525 C 1E Stress, MPa Fig. 2: Stress-temperature dependence of time to rupture of the heat A after controlled rolling, Larson-Miller equation with k=2-7 -

8 1E+6 1E+5 1E+4 Time, h 1E C 475 C 500 C 525 C 1E Stress, MPa Fig. 3: Stress-temperature dependence of time to rupture of the heat B after normalizing, Larson-Miller equation with k=3 1E+6 1E+5 Time, h 1E+4 1E+3 1E C 475 C 500 C 525 C Stress, MPa Fig. 4: Stress-temperature dependence of time to rupture of the heat C after controlled rolling, Larson-Miller equation with k=3-8 -

9 CRS at h, MPa A1 B1 B2 C1 C2 Heat C 500 C 44,7 Fig. 5: Comparison of creep rupture strength in as rolled condition ( A1, B1, C1 ) and after normalizing ( B2, C2 ) St 35.8 St 45.8 HEAT A1, C1 Heat B2, C1 Heat C1 Heat B1 16Mo3 20MnNb , C2 160 P235GH P265GH 16Mo3 140 STRESS [ MPa ] St 45.8 St 35.8 P265GH P235GH 80 20MnNb , C2 60 Heat A TEMPERATURE [ C ] Fig. 6: How the proof stress and creep rupture strength of carbon steels (St 35.8, St 45.8) microalloyed steels (P235GH, P265Gh, 20MnNb6, heats A1, B1, C1, C2) and - 9 -

10 low alloy steel (16Mo3) depends on the temperature. Heats A1, B1, C1 after normalizing forming. Heat B2, C2 after normalizing aging h [M Pa 5 Cr ee p ru pt ur e str en gt C 500 C ,7 0 20MnNb Heat B2 Heat C2 Fig. 7: Compression of creep rupture strength of steels 20MnNb6 and according corresponding standards with CRS of heats B2 and C2 attained after normalizing

11 d 0 [ nm] Fig. 8 : How the ratio of interparticle spacing after creep exposure in hours l 10 5 and the interparticle spacing in as received conditions l 0 ( l 10 5 ; l 0 ) depends on temperature and mean diameter of particles in the as received d 0. 1,4,7 before creep exposure 2,5,8 after creep exposure 550 C/10 5 h 369 after creep exposure 600 C/10 5 h 1,2,3 0,12C - 0,5Cr - 0,5Mo - 0,3V 4,5,6 0,12C - 9Cr - 1Mo 7,8,9 0,20C - 11Cr - 1Mo - 0,3V l [nm ] d 0[ nm] Fig. 9 : How the interparticle spacing l depends on the mean diameter of particles received conditions d 0, volume fraction f and temperature