当前位置:首页 >> 能源/化工 >>

Some aspects of metallurgical assessment of boiler tubes—Basic principles and case studies


Materials Science and Engineering A 432 (2006) 90–99

Some aspects of metallurgical assessment of boiler tubes—Basic principles and case studies
Satyabrata Chaudhuri ?
Nat

ional Metallurgical Laboratory, Jamshedpur 831007, India Received 1 March 2006; received in revised form 11 May 2006; accepted 12 June 2006

Abstract Microstructural changes in boiler tubes during prolong operation at high temperature and pressure decrease load bearing capacity limiting their useful lives. When the load bearing capacity falls below a critical level depending on operating parameters and tube geometry, failure occurs. In order to avoid such failures mainly from the view point of economy and safety, this paper describes some basic principles behind remaining life assessment of service exposed components and also a few case studies related to failure of a reheater tube of 1.25Cr–0.5Mo steel, a carbon steel tube and ?nal superheater tubes of 2.25Cr–1Mo steel and remaining creep life assessment of service exposed but unfailed platen superheater and reheater tubes of 2.25Cr–1Mo steel. Sticking of ?y ash particles causing reduction in effective tube wall thickness is responsible for failure of reheater tubes. Decarburised metal containing intergranular cracks at the inner surface of the carbon steel tube exhibiting a brittle window fracture is an indicative of hydrogen embrittlement responsible for this failure. In contrast, ?nal superheater tube showed that the failure took place due to short-term overheating. The in?uence of prolong service revealed that unfailed reheater tubes exhibit higher tensile properties than that of platen superheater tubes. In contrast both the tubes at 50 MPa meet the minimum creep rupture properties when compared with NRIM data. The remaining creep life of platen superheater tube as estimated at 50 MPa and 570 ? C (1058 ? F) is more than 10 years and that of reheater tube at 50 MPa and 580 ? C (1076 ? F) is 9 years. ? 2006 Elsevier B.V. All rights reserved.
Keywords: Metallurgical assessment; Platen superheater; Reheater; Final superheater; Creep rupture; Tensile; Microstructures; Remaining creep life

1. Introduction Boiler tubes are energy conversion components where heat energy is used to convert water into high pressure superheated steam, which is then delivered to a turbine for electric power generation in thermal power plants, or to run plant and machineries in a process or manufacturing industry. Heat energy is obtained from combustion gases produced by burning coal or oil in the furnace. The combustion gases evaporate water into steam in the waterwall tubes and thereafter pass over the superheater and reheater tubes. After turning down they encounter primary superheater and economizer tubes. The combustion gases are fed through air preheater and cleaning devices before exiting through the stack. The boiler tubes designed for a speci?c period of time operate in a complex situation involving high temperature, pressure and corrosive environment. Several natural ageing processes such as creep, corrosion, fatigue, etc. occurring during prolong operation are responsible for accumulating
?

microstructural damages in the tubes. The effective strength, i.e. load bearing capacity of the tubes due to microstructural damages decreases. The failure occurs when it falls below a critical level determined by component geometry and loading. Such failure is the major problem concerning the availability of boilers. Since the failure results in non-availability of electric power, loss of industrial production, etc. life assessment exercise performed at regular intervals is a means to ensure avoidance of such failures. Some important remaining life assessment methodologies are based on empirical models using creep strain measurement, combined time–temperature parameter, tube wall thinning, oxide scale thickness measurement, hardness measurement and microstructural assessment, etc. The creep database [1–8] of indigenously produced, service exposed and failed boiler tubing materials is used for their metallurgical assessment and remaining life prediction. 2. Creep resistant steels Creep resistant steels are extensively used for large-scale chemical, thermal and nuclear power and petroleum industries.

Tel.: +91 657 2271709x2016; fax: +91 657 2270527. E-mail addresses: sc1@nmlindia.org, schaudhuri06@yahoo.co.in.

0921-5093/$ – see front matter ? 2006 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2006.06.026

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99 Table 1 Nominal composition of creep-resistant steels (wt.%) Steel type C Cr Mo Other elements

91

Group 1: carbon steels Tubes SA-192 0.06–0.18 Pipe SA-106 0.35 Group 2: low alloy steels and 9Cr–Mo steels 1Cr–0.5Mo 0.11 2.25Cr–1Mo 0.12 0.5Cr–0.5Mo–0.25V 0.11 9Cr–1Mo 0.10 Group 3: 12Cr steels 12Cr–Mo–V 0.12 12Cr–Mo–V–Nb 0.11 Group 4: austenitic steels Type 304 0.05 Type 316 0.05 Esshete 1250 0.10

Mn: 0.27–0.63, Si: 0.25 Mn: 0.29–1.06, Si: 0.10 minimun 1.0 2.25 0.5 9.0 12.0 11.0 18.5 18.0 15.0 0.5 1.0 0.5 1.0 0.6 0.6 Mn: 0.5, Si: 0.25 Mn: 0.5, Si: 0.25 Mn: 0.5, Si: 0.25, V: 0.25 Mn: 0.5, Si: 0.60 V: 0.2, Ni: 0.80 V: 0.2, Ni: 0.8, Nb: 0.35 Mn: 1.3, Ni: 10 Mn: 1.4, Si: 0.4, Ni: 10 Mn: 6, Si: 0.5, Ni:10, V: 0.3, Nb: 1

2.5 1.0

Some of the important factors to be considered for selection of such steels are resistance to creep deformation and rupture, resistance to environmental attack, creep rupture strength and ductility of weld metal and heat affected zone, adequate ductility of base material to avoid sudden failure and also to allow the material to deform rather than fracture in the regions of high stress concentrations. Lack of rupture ductility has been reported as the cause of cracking in power station steam pipes made of 0.5% Mo steel. This cracking resulted from repeated heating and cooling to and from service temperature of about 565 ? C (1049 ? F) during twoshift operation. The problem was eliminated by replacement of 0.5% Mo steel with more ductile 1% Cr–Mo steel. The normal compositions of some creep resistant steels are given in Table 1. The steels have been classi?ed into four groups for use in the increasing order of operating temperature. They are carbon steels, low alloy Cr–Mo/9Cr–Mo steels, 12Cr-steels and austenitic steels [9]. The ASME boiler and Pressure Vessel Code, Paragraph A-150 of section I states the criteria for determining allowable stresses. The allowable stresses are not to be higher than the lowest of the following: ? One-fourth of the speci?ed minimum tensile strength at room temperature. ? One-fourth of the tensile strength at elevated temperature. ? Two-third of the speci?ed minimum yield strength at room temperature. ? Two-third of the yield strength at elevated temperature. ? Stress to produce 1% creep in 100,000 h. ? Two-third of the average stress or four-?fth of the minimum stress to produce creep rupture in 100,000 h, whichever is minimum. These criteria are employed to establish allowable stresses for a range of steels as a function of temperature A comparison of the allowable stresses at various temperatures for commonly used steels is shown in Fig. 1 [10,11]. For 2.25Cr–1Mo steel, it is the creep or rupture strength that determines the allowable stress at a temperature of beyond 482 ? C (900 ? F). Therefore, in the

Fig. 1. Allowable stress for several grades of steels as a function of temperature.

evaluation of creep behavior of Cr–Mo steels, estimation of longterm rupture strength has received considerable importance. 3. Life assessment methodology Life assessment methodology can broadly be classi?ed into three levels [12]. Level 1 methodology is generally employed when service life of the components is less than 80% of their design lives. In level 1, assessments are performed using plant records, design stress and temperatures, and minimum values of material properties from literature. When service life exceeds 80% of the design life, level 2 methodology is employed. It involves actual measurements of dimensions and temperatures, stress calculations and inspections coupled with the use of the minimum material properties from literature. However, when life extension begins after attaining design life, level 3 methodology is employed. It involves in-depth inspection, stress analysis,

92

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

plant monitoring and generation of actual material data from samples removed from the component. The details and accuracy of the results increase from level 1 to level 3 but at the same time the cost of life assessment increases. Depending on the extent of information available and the results obtained, the analysis may stop at any level or proceed to the next level as necessary. One of the crucial parameters in estimation of creep life is the operating temperature. Although steam temperatures are occasionally measured in a boiler, local metal temperatures are rarely measured. Due to load ?uctuations and steam-side oxide-scale growth during operation, it is also unlikely that a constant metal temperature is maintained during service. It is therefore, more convenient to estimate mean metal temperature in service by examination of such parameters as hardness, microstructure, and thickness of the steam-side oxide-scale for tubes. Because the changes in these parameters are functions of time and temperature, their current values may be used to estimate mean metal temperature for a given operating time. The estimated temperature can then be used in conjunction with standard creep rupture data to estimate the remaining life. Several methods for estimation of metal temperature have been reviewed elsewhere [13]. 3.1. Hardness-based approach The strength of low alloy steel changes with service exposure depends on time and temperature. Thus change in hardness during service (Fig. 2) may be used to estimate mean operating temperature for the component. This approach is particularly suitable when strength changes in service occur primarily as a result of carbide coarsening neglecting stress induced softening. The database on changes in hardness due to long-term service is employed to assess remaining life [13].

3.2. Microstructure-based approach Toft and Mardsen [14] demonstrated that there are basically six stages of spheroidisation of carbides in ferritic steels. Using Sherby–Dorn parameter, they established a reasonable correlation of microstructure with mean service temperature. Similar semi quantitative and qualitative approaches involving database on changes in microstructure as a function of service history have been widely used [15]. 3.3. Oxide scale thickness-based approach Extensive data from literature indicate that in relatively pure steam, the growth of oxide scales is a function of temperature and time of exposure. Several expressions have been proposed in the literature to describe oxide scale growth kinetics [16,17]. This approach is particularly suitable when effective operating stress increases and strength changes due to the growth of oxide scale in service neglecting stress induced oxidation. 4. Failure of reheater tube A failed reheater tube (Fig. 3) in a 500 MW boiler was of pendent type placed in the convective zone. The operating pressure was 25 kg/cm2 and steam outlet temperature was 535 ? C (995 ? F). The design temperature of ?ue gas in the zone was 700–720 ? C (1292–1328 ? F). The tube material was 1.25Cr–0.5Mo steel. The tube had suffered extensive damage on the outer surface in the form of pits. The dimension of the pits at some places was as big as 40 mm × 10 mm with a maximum depth of 2 mm. The pitted surface bore brownish color in contrast to the damage free surfaces of the tube. The latter bore usual black oxide. The failure had occurred after about 24,000 h of service life. The investigation carried out at the laboratory showed that the tube had the typical microstructure consisting of ferrite and bainite and the mechanical properties were also found to be normal [18]. No appreciable diametrical expansion was observed. Since the tube showed extensive pitted surfaces, X-ray microanalysis using SEM was carried out on the pitted surfaces to ascertain the presence of corrosive elements. The presence of highly corrosive elements like K, Ca, Si, S and Cl were observed. The presence of these elements suggested that the attack was

Fig. 2. Changes in hardness with Larson–Miller parameter.

Fig. 3. Extensive damage on outer surface of failed reheater tube [4].

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

93

caused by the fusion of ash particles. Such attack usually occurs due to: ? Volatilization and condensation of volatile ash constituents containing Na2 SO4 or CaSO4 . ? Temperature excursion beyond 650 ? C (1202 ? F) particularly during the starting period when no steam ?ows through the reheater tubes. Ash particles of low fusion point can fuse and stick on the tube surfaces at such temperatures. It may be noted that alkali metals along with S and Fe can form ash with fusion temperature as low as 620 ? C (1148 ? F). The fuel oil used for the support can also cause this problem if the oil contains corrosive elements like V and S. Tube wall thinning due to sticking of ?y ash particles is primarily responsible for the failure of the reheater tube. 5. Failure of carbon steel tube A failed carbon steel tube (Fig. 4a) was of the brittle window type without the ?sh mouth appearance. The operating temperature of the tube was 350 ? C (662 ? F) The outer diameter and thickness of the tube are 45 and 4.5 mm, respectively. Visual examination showed a substantial degree of corrosion on the water surface leaving a rough area in the vicinity of rupture. Microstructural examination of a cross section through tube wall revealed decarburisation and extensive discontinuous ?ssures (Fig. 4b). The presence of oxide layer (Fig. 4b) and concentration of chlorides at the interface between oxide and metal at the base of corrosion pits (Fig. 4c) are also evident from microstructural examination. The combination of water side corrosion (oxidation), decarburisation and ?ssures, as observed in this case, led to the conclusion that the tube had failed because of hydrogen damage involving the formation of methane by the reaction of dissolved hydrogen with carbon in steel.

Table 2 Material speci?cation and operating parameters of superheater tubes Material Outer diameter Thickness Steam temperature Steam pressure Steam ?ow rate 2.25Cr–1Mo steel as per ASTM A213 T22 44.5 mm 10 mm 560–580 ? C (1040–1076 ? F) 185 kg/cm2 1700 tonnes/h

Fig. 5. Virgin, undamaged and failed ?nal superheater tube of a 500 MW boiler.

6. Failure of ?nal superheater tubes Failure of ?nal superheater tubes of a 500 M boiler occurred during trial run following a service exposure of about 100 h [4]. Material speci?cation and operating parameters of the tubes are summarized in Table 2. The tube samples selected for this investigation (Fig. 5) are a piece of tube from the zone of failure, a piece of tube adjacent to the failed tube, termed as undamaged tube and a piece of virgin tube.

Fig. 4. (a) Brittle window type fracture of a carbon steel tube; (b) microstructure through tube wall near rupture revealing decarburised metal and extensive discontinuous ?ssures; (c) microstructure revealing chloride distribution at the interface between oxide and metal at corrosion pits [4].

94 Table 3 Chemical composition (in wt.%) Element Carbon Manganese Silicon Phosphorus Sulphur Chromium Molybdenum Failed tube 0.12 0.41 0.22 0.013 0.003 2.16 1.05 Virgin tube 0.1 0.42 0.22 0.012 0.003 2.13 1.00

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

Speci?cation ASTM A213 T22 0.06–0.15 0.30–0.60 0.50 maximum 0.25 maximum 0.025 maximum 1.90–2.50 0.87–1.13

Table 4 Hardness values of virgin, undamaged and failed tube Tubes Hardness, HV20 Inner Virgin Undamaged Failed 145 143 179 Middle 149 143 178 Outer 151 145 180 Fig. 6. Tensile properties of virgin, undamaged and failed ?nal superheater tube of a 500 MW boiler.

The chemical composition of the failed and virgin tube, reported in Table 3, indicates that the chemistry of both tubes meets the ASTM speci?cation. The outer diameter of the failed tube was found to be 49.2 mm against the original diameter of 44.5 mm. The most signi?cant point in this case is the gross circumferential expansion of the failed tube up to about 19%. Such an extensive expansion cannot be expected under normal operating condition within a short span of service exposure of about 100 h. It is evident from the measurement of hardness at the inner surface, mid section and outer surface of all tubes (Table 4) that the hardness of the failed tube is signi?cantly higher than that of virgin and undamaged tubes.

The tensile properties of all tubes (Fig. 6) revealed that irrespective of test temperature all the tubes meet the minimum speci?ed properties of 2.25Cr–1Mo steel. The tensile strength of the failed tube is, however, signi?cantly higher than that of the other tubes. The microstructures of virgin (Fig. 7a) and undamaged (Fig. 7b) tubes are almost similar, consisting of ferrite and tempered bainite. In contrast the microstructure of the failed tube (Fig. 7c) showed the presence of freshly formed bainitic areas. Higher hardness and higher tensile strength as exhibited by the failed tube based on the comparison with other virgin and undamaged tubes con?rms the presence of freshly formed bainite in the failed tube. Besides, the oxide scale thickness on the inner surface of the failed tube (Fig. 8a) was found to be several folds more than that of the undamaged tube (Fig. 8b). In the failed tube the oxide scale thickness was 0.25 mm.

Fig. 7. Microstructures of: (a) virgin tube; (b) undamaged tubes consisting of ferrite and tempered bainite and (c) failed tube showing freshly formed bainitic regions.

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

95

Fig. 8. (a) 0.25 mm thick oxide scale at inner surface of failed tube and (b) insigni?cant oxide scale at inner surface of undamaged tube.

All the above observations can be reconciled in a situation only if the temperature had exceeded the lower critical temperature, which is reported to be about 800 ? C (1472 ? F) for 2.25Cr–1Mo steel. In order to predict the extent of temperature excursion beyond the critical temperature some detailed analysis was made based on oxide thickness as obtained on the inner surface of the failed tube. Published kinetic data on oxide scale growth of 2.25Cr–1Mo steel have been used for this purpose [11]. The predicted time–temperature pro?les in the temperature range of 500–850 ? C (932–1562 ? F) for a range of service exposure to develop 0.25 mm thick oxide at the inner surface of the failed tube are shown in Fig. 9. Since the presence of freshly formed bainite is indicative of temperature excursion beyond the lower critical temperature, the most probable pro?le that the tube experienced is the one having an exposure of 2 h, the maximum temperature being 830 ? C (1526 ? F). Corresponding to each time–temperature pro?le (Fig. 9), accumulation of strain and the diametrical expansion of the tube have been calculated and presented in Fig. 10. The creep strain predicted from the time–temperature pro?le for 2 h exposure comes to about 1%, which is indeed quite lower than the observed value of 19%. It is mainly because the existing material database in the temperature range of 500–600 ? C (932–1112 ? F) has been used for extrapolation. Clearly there is a need to collect creep data at temperatures beyond 600 ? C (1112 ? F).

In order to ensure whether it is possible to achieve a creep strain of about 19% within a short time at 830 ? C (1526 ? F), a short-term creep test has been carried out in the laboratory. Based on this it has been established that a creep strain of about 16% is achievable in less than 2 h at 830 ? C (1526 ? F) and a stress level of 30 MPa, which represents the hoop stress corresponding to the maximum operating pressure of 185 kg/cm2 for the tube in question. This, therefore, conclusively proves that the failure of the tube took place due to short-term overheating to a temperature of about 830 ? C (1526 ? F). Partial chocking of the tube by some foreign material could be responsible for such overheating. The other tubes, however, did not suffer any heavy temperature excursion beyond 650 ? C (1202 ? F). The presence of freshly formed bainitic structure and higher hardness in the damaged failed tube is responsible for higher strength. This is mainly because the ferrite and freshly formed bainitic structure in the damaged tube is stronger than ferrite and partially degenerated tempered bainite as observed in virgin and undamaged tubes. Although the high temperature mechanical properties of service exposed components are often found to be better than the minimum speci?ed level, the component dimensions often change as a result of prolonged service. The most prominent amongst these is the loss of tube wall thickness. The growth of oxide scale at tube surfaces is primarily responsible for this. Other external processes such as coal ash corrosion, ?ame impingement and ?y ash erosion also contribute to such damage development in service. A need is thus felt to look into the life

Fig. 9. Predicted time–temperature pro?les to develop 0.25 mm thick oxide layer at inner surface of failed tube.

Fig. 10. Predicted creep strain corresponding to time–temperature pro?le shown in Fig. 9.

96 Table 5 Material speci?cation for boiler tubes Tube sample PLSH coil (outlet) RH coil (outlet) Identi?cation L-R-CKT-25 L-R-CKT-1

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

Nominal dimension, o.d. × Th (mm) 51 × 8.8 54 × 3.6

Measured dimension, o.d. × Th (mm) 51.48 × 8.91 54.28 × 5.03

Grade of steel SA 213 T22 SA 213 T22

Table 6 Operating parameters of service exposed tubes Tube sample PLSH tube (outlet) RH tube (outlet) Designed steam temperature, ? C (? F) 544 ? C (1011 ? F) 580 ? C (1076 ? F) Operating pressure (kg/cm2 ) 158.2 33.5 Length of service (h) 59585 59585 Zone Platen superheater Reheater

prediction problems considering the in?uence of wall thinning due to the existence of corrosion and erosion processes during operation of the plant. Keeping this in view the methodology recently developed in the laboratory for creep life estimation of boiler tubes under wall thinning condition highlighting some typical results has been discussed elsewhere [4]. 7. Creep life assessment of service exposed platen superheater and reheater tubes Service exposed platen superheater (PLSH) and reheater (RH) tubes of a thermal power plant were identi?ed for remaining creep life assessment study [20]. The grade of steel, dimensions and identi?cation numbers of these tube samples are given in Table 5. The operating parameters of the service exposed tubes as obtained from the plant engineers are given in Table 6. The experimental work undertaken for assessing remaining life of service exposed boiler tubes includes tensile tests, hardness measurement, microstructural examination and creep rupture tests. The test data so generated have been analyzed and compared with National Research Institute for Metals (NRIM) data for 2.25Cr–1Mo steel to examine the in?uence of service exposure on mechanical properties and remaining life of boiler tubes. Longitudinal section of service exposed platen superheater and reheater tube samples Fig. 11(a and b) showed grayish oxide layer on the inner surfaces. The measured outer diameter and thickness of the service exposed tubes are given in Table 7. 7.1. Tensile tests and hardness measurement Standard tensile specimens were made from the longitudinal direction of the service exposed boiler tubes to carry out tensile tests in air using Instron 8562 servo electric machine at a conTable 7 Measured dimensions of service exposed tubes Tube nos. L-R-CKT-25 L-R-CKT-1 Outer diameter (mm) 51.48 54.28 Thickness (mm) 8.91 5.03

stant displacement rate of 0.008 mm/s. The range of temperature selected for tensile tests are from 25 to 650 ? C (77–1202 ? F) for both the platen superheater and reheater tubes. Hardness measurement at 25 ? C (77 ? F) was carried out on specimens selected from each type of boiler tube. The results of tensile tests viz. YS/0.2% PS, UTS and % elongation as a function of temperature as well as hardness measurement for PLSH and RH tubes are shown in Tables 8a and 8b. The temperature dependence of 0.2% PS/YS data and UTS data of both the tubes in the range of 25–650 ? C (77–1202 ? F) are shown graphically in Figs. 12 and 13, respectively. For the purpose of comparison, the minimum NRIM data are also shown in these ?gures. 7.2. Microstructural examination Specimens for microstructural examination using optical microscope were made following standard procedure from the transverse direction of the service exposed tubes. Microstructures of service exposed platen superheater and reheater tubes are shown in Fig. 14 and Fig. 15, respectively. The microstructures, in general, consist of degenerated bainitic regions in terms of spheroidisation of carbides. Such microstructural features are expected from the tubes, which have undergone a prolonged service exposure. The extent of degeneration of bainitic regions is almost similar in PLSH and RH tubes

Fig. 11. (a and b) Service exposed platen superheater and reheater tube samples.

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99 Table 8a Tensile properties and hardness measurement of service exposed PLSH tubes Tube no. L-R-CKT-25 L-R-CKT-25 L-R-CKT-25 Temperature, ? C (? F) 600 ? C (1112 ? F) 650 ? C (1202 ? F) 25 ? C (77 ? F) YS/0.2% PS (MPa) 248 143 117 UTS (MPa) 453 180 139 % Elongation 19 33 37

97

Hardness HV30 137

Table 8b Tensile properties and hardness measurement of service exposed RH tubes Tube no. L-R-CKT-1 L-R-CKT-1 L-R-CKT-1 Temperature, ? C (? F) 600 ? C (1112? F) 650 ? C (1202 ? F) 25 ? C (77 ? F) YS/0.2% PS (MPa) 259 166 129 UTS (MPa) 501 227 167 % Elongation 14 19 21 Hardness HV30 143

Fig. 14. Microstructure of service exposed platen superheater tubes. Fig. 12. Temperature dependence of 0.2% PS/YS of PLSH and RH tubes.

(Figs. 14 and 15). The presence of bainitic microstructures is also con?rmed from the fact that 2.25Cr–1Mo steel in normalized and tempered condition exhibits a wide range of microstructures consisting of ferrite and tempered bainite when normalized at 920 ? C (1688 ? F) to fully tempered bainitic microstruc-

tures when normalized at 990 ? C (1814 ? F) followed by forced air cooling, the tempering temperature remaining the same at 730 ? C (1346 ? F) [8]. The microstructural features in PLSH and RH tubes did not reveal presence of any signi?cant damages such as graphitisation, cavities, microcracks, etc. The thin oxide scale is present at the inner surfaces of both the tubes [20].

Fig. 13. Temperature dependence of UTS of PLSH and RH tubes.

Fig. 15. Microstructure of service exposed reheater tubes.

98

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99

Table 9a Creep rupture properties of service exposed PLSH tubes Tube no. (type) L-R-CKT-25 L-R-CKT-25 L-R-CKT-25 Temperature, ? C (? F) 650 ? C (1202 ? F) 600 ? C (1112 ? F) 600 ? C (1112 ? F) Stress (MPa) 50 100 80 Rupture time (h) 1073 144 851 % Elongation 29 31 28

Table 9b Creep rupture properties of service exposed RH tubes Tube no. (type) L-R-CKT-1 L-R-CKT-1 L-R-CKT-1 Temperature, ? C (? F) 650 ? C (1202 ? F) 600 ? C (1112 ? F) 600 ? C (1112 ? F) Stress (MPa) 50 100 80 Rupture time (h) 1022 1057 2904 % Elongation 13 21 Test interrupted

7.3. Creep rupture tests Standard test specimens were made from the longitudinal direction of the service exposed boiler tubes to carry out creep rupture tests in air using Mayes creep testing machines. The temperature levels selected for these tests are 600 and 650 ? C (1112 and 1202 ? F). The stress levels for these tests are selected to obtain rupture within a reasonable span of time. The results of creep rupture tests of the specimens made from PLSH and RH tubes are shown in Tables 9a and 9b. Larson–Miller parameter (LMP) as a function of stress were calculated using the formula LMP = T(20 + log tr ) where T is the temperature in (K) and tr is the rupture time in (h). The dependence of LMP values on stress is shown graphically in Fig. 16. For the purpose of comparison minimum NRIM data are also shown in Fig. 16. 7.4. Results and discussion Visual examination did not reveal presence of any signi?cant oxidation/corrosion damages on the outer surfaces of both tubes. The inner surfaces, however, showed the presence of grayish oxide layer Fig. 11(a and b). The reduction in tube wall thickness in reheater and platen superheater tubes as obtained from measurement and compared with the speci?ed thickness was not observed. The results of tensile tests of service exposed PLSH and RH tubes reveal the following features based on comparison with NRIM data for the same grade 2.25Cr–1Mo steel tube [19]. In general, deterioration in 0.2% PS (Fig. 12) and UTS (Fig. 13) of both PLSH and RH tubes, when compared with minimum NRIM data, was observed. The RH tubes exhibit higher 0.2% PS and UTS than that of PLSH tubes irrespective of test temperature. The extent of degradation is more pronounced in case of PLSH tube. The 0.2% PS at 600 ? C (1112 ? F) and UTS at 25 ? C (77 ? F) of RH tube meet the minimum NRIM data. The remaining lives of service exposed tubes were estimated using Larson–Miller parameter (LMP). The results of accelerated creep rupture tests of service exposed PLSH and RH tubing steel as obtained in the laboratory are utilized for this purpose. In absence of virgin PLSH and RH tubes, the accelerated creep rupture data of service exposed tubes are compared with minimum NRIM data. In contrast to deterioration in 0.2% PS and UTS of both PLSH and RH tubes, the applied stress dependence of LMP for these tubes as shown in Fig. 16 clearly indicate that the creep properties of both tubes at 50 MPa meet the minimum properties when compared with NRIM data. At a stress level of more than 50 MPa, the creep rupture data of PLSH tube fall marginally below the creep rupture data of RH tube. The deviation in terms of LMP is longer at higher applied stress of more than 50 MPa and gradually increases with increasing applied stress. Although the operating hoops stress as estimated from the nominal tube dimension and operating pressure is 37 MPa for PLSH tube and 16 MPa for RH tube, the remaining life has been estimated from experimentally obtained creep rupture data at the lowest stress level of 50 MPa and reported in Table 10. 7.5. Conclusion Based on accelerated creep rupture tests and analysis of data, it can be said that the service exposed platen superheater

Fig. 16. Stress vs. LMP plot of PLSH and RH tubes.

S. Chaudhuri / Materials Science and Engineering A 432 (2006) 90–99 Table 10 Remaining life of PLSH and RH tube Tube nos. L-R-CKT-25 L-R-CKT-1 Type of tube Platen superheater tube Reheater tube Estimated life at 50 MPa (years) 10 years at 570 ? C (1058 ? F) 9 years at 580 ? C (1076 ? F)

99

References
[1] S. Chaudhuri, R.N. Ghosh, ISIJ Int. 38 (1998) 881–887. [2] S. Chaudhuri, in: R. Singh, S.K. Sinha, S. Chaudhuri (Eds.), Power Plant Metallurgy, NML, Jamshedpur, India, 1997, pp. 56–73. [3] S. Chaudhuri, N. Roy, R.N. Ghosh, Acta Metall. Mater 41 (1993) 273–278. [4] S. Chaudhuri, R. Singh, in: S.R. Singh, et al. (Eds.), Failure Analysis, NML, Jamshedpur, India, 1997, pp. 107–120. [5] S. Chaudhuri, in: I. Chattoraj, et al. (Eds.), Boiler Corrosion, Proceedings of the National Workshop (NWBC-95), NML, Jamshedpur, India, 1995, pp. M1–M22. [6] R. Singh, S. Chaudhuri, et al., NML Report No. 11620034, December 1990. [7] R. Singh, S. Chaudhuri, et al, NML Report No. 11620035, December 1990. [8] S. Chaudhuri, Ph.D. Thesis, IIT, Kharagpur, 1994. [9] S. Chaudhuri, Proceedings of a National Workshop on Component Integrity Evaluation Program and Life Assessment, September 1996, p. 55. [10] ASME Boiler & Pressure Vessel Code, Section 1, ASME, New York. [11] R. Viswanathan, Damage Mechanisms and Life Assessment of High Temperature Components, ASM International, 1989, pp. 185–229. [12] R. Viswanathan, R.B. Dooley, in: International Conference on Creep, ASTM, Tokyo, April 14–18, JSME, J. Mech. E., ASME (1986) 349–359. [13] R. Viswanathan, J.R. Foulds, D.A. Roberts, Proceedings of the International Conference on Life Extension and Assessment, Hague, June 1988. [14] L.H. Toft, R.A. Mardsen, Conference on Structural Processes in Creep, JISI/JIM, London, 1963, p. 275. [15] R. Coade, Report No.SO/85/87, State Electricity Commission of Victoria, Australia, February 1885. [16] I.M. Rehn, W. Apblett, Report CS 1811m Electric Power Research Institute, Palo Alto, CA, 1981. [17] S.R. Paterson, T.W. Rettig, Project RP 2253-5, Final Report, Electric Power Research Institute, Palo Alktio, CA, 1987. [18] S. Banerjee, R. Singh, K. Prasad, NML Report, October 1991. [19] NRIM Creep Data Sheet No. 3A, 1976. [20] S. Chaudhuri, Final Report, NML, Jamshedpur, June 2003.

and reheater tubes are in a good state of health for its continued service provided no localized damages in the form of tube wall thinning, circumferential expansion, excessive oxide scale formation, microstructural degradation, etc. are present besides adherence to the speci?ed operating parameters in service. The PLSH tube may be allowed to remain in service for a period of 10 years under similar operating condition provided tube wall temperature does not exceed 570 ? C (1058 ? F). The RH tube may be allowed to remain in service for a period of 9 years under similar operating condition provided tube wall temperature does not exceed 580 ? C (1076 ? F). A more reliable estimate of remaining life is obtained by considering simultaneously NDT measurement at site and their analysis, actual operating history, destructive tests and their analysis in the laboratory. Acknowledgements Changes in microstructures and mechanical properties of service exposed boiler tubes presented in this paper is the outcome of applied sponsored research for component integrity evaluation program. The author wishes to thank Director, National Metallurgical Laboratory, Jamshedpur, India for his kind permission to publish this paper.


相关文章:
更多相关标签: