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Fitness for Service of Degraded Grade 91 Pipe

Conference Paper · July 2012

DOI: 10.1115/PVP2012-78589

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Fitness for Service of Low Hardness Grade 91 Pipe

TP202 November 2012

Prepared By

Marvin J. Cohn, P.E.

Intertek APTECH Sunnyvale, CA USA

Steve R. Paterson

PIKA Solutions Los Gatos, CA USA

Organizer - European Technology Development

Cyclic Operation of Power Plants, November 29, 2012 Charlotte, NC

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Fitness for Service of Low Hardness Grade 91 Pipe Marvin J. Cohn, P.E

Intertek APTECH Sunnyvale, CA USA

Steve R. Paterson PIKA Solutions

Los Gatos, CA USA

1. Abstract The use of creep strength enhanced ferritic alloys, such as Grade 91 materials, in fossil power plants has become popular for high temperature piping applications. Since Grade 91 components have higher stress allowables than Grade 22 components, a piping designer can specify a thinner component wall thickness for the same applied stress, resulting in lower through-wall thermal stresses during transient events and lower material and piping support costs. Grade 91 has been used in the power industry over the past two decades. In some instances, there have been some incidents of non-optimal weldment microstructure and low hardness measurements. In this case study, Brinell hardness tests of an ASME A182 Grade F91 (F91) wye block, including upstream and downstream F91 spools, revealed several readings of soft material, as low as 168HB. In this paper, a life consumption evaluation was performed for low hardness Grade 91 weldments based on a lower bound stress rupture curve for weldment anomaly creep rupture tests. Life fraction analyses were performed, considering the redistributed maximum principal stresses based on simulation of piping thermal displacements obtained from the hot and cold walkdowns. This study also considered the recent history of the specific piping system operating pressures and temperatures. In particular, the life consumption evaluation considered standby (low power) and normal operations, where the unit is at about 60% pressure nearly 50% of the operating time. It was determined that the Grade F91-to-F91 low hardness weldments had less than 12% life consumption and the remaining lives were at least 20 years.

Keywords: Fitness for Service, life consumption, Grade 91, low hardness, high energy piping Disclaimer: The views expressed in this paper are strictly those of the authors and do not necessarily reflect those of their affiliated corporations.

2. Introduction This study is an evaluation of low hardness Grade 91 piping material in a main steam (MS) piping system.

The piping system began operation prior to 1980 and has accumulated more than 200,000 operating hours. The initial piping layout design was Grade 11 material. The piping system isometric is illustrated in Figure 1.

In 1994, the original ASTM A182 Grade F11 wye block was replaced with a significantly thinner ASTM A182 Grade F91 wye block, including ASTM A182 Grade F91 spools welded to the upstream and downstream wye block connections. In addition, an ASTM A369 Grade FP22 transition spool was added to the upstream leg and an ASTM A369 Grade FP 22 transition spool was added to each of the downstream legs of the wye block. An illustration of the wye block and adjacent welds is provided in Figure 2.

In May 1998, indications were revealed at Weld 15 (the west leg Grade F91 wye block to F91 pup transition spool). A comprehensive root cause investigation that included monitoring the local pipe temperatures during

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shutdown and startup transients revealed that the cracking was driven by a combination of 1) quenching of the wye block from fan cooling of the boiler when necessary for maintenance (the rapid cooling of the steam in the superheater tubes would condense, migrate, and quench the MS wye block), 2) increased sustained load stresses due to malfunctioning piping supports, 3) poor control of the weld post weld heat treating temperatures leading to a soft base metal heat affected zone, and 4) a weld detail design with thickness and alloy transitions coincident with the weld.

Corrective actions included 1) changes to the shutdown operations using a bypass system to minimize quenching from condensate migration, 2) verification that quenching was not occurring through the use of diagnostic/troubleshooting thermocouples that were installed on the supply and outlet legs of the wye block, 3) piping support adjustments to restore the spring hangers into their normal operating range, and 4) excavation of the weld deposits and soft heat affected zones followed by weld repairs of the west leg pipe from Girth Weld 15 through Girth Weld 21.

In 2001, the welds in the vicinity of the wye block were examined again. The examination of Weld 15 revealed an 11-inch long by 5/8-inch deep crack on the downstream side of the weldment (repaired in 1998). Girth Welds 15 and 16 were excavated, welded, post weld heat treated, reexamined, and returned to service.

In April 2009, 15 lead-the-fleet locations with higher inservice stresses were examined. No crack-like indications were revealed. Girth Welds 13, 15, 16, 17, 18, 20, and 21 had lack of fusion (LOF) indications within the weld volume. Girth Weld 14 had indications of stacked slag, inclusions, and LOF indications of the weld volume. Hardness readings indicated that non-optimal weld metal microstructure may exist in Girth Weld 14. As a result, Girth Weld 14 was excavated part way through the wall thickness and repaired during the scheduled outage.

As of December 2011, the unit had operated somewhat more than 100,000 hours since the 1994 wye fitting replacement. In 2011 wet fluorescent magnetic particle examinations were performed on Girth Welds 13 through 18 and there were no reportable indications. The phased array ultrasonic examinations of Girth Welds 13 through 18 revealed that Girth Weld 14 had an elongated linear midwall reflector, Girth Weld 16 had a 13 mm (0.5-inch) long midwall reflector, Girth Weld 17 had a 13 mm (0.5-inch) long midwall reflector, and Girth Weld 18 had two 13 mm (0.5-inch) long midwall reflectors.

In 2011, hardness testing was performed using a Telebrinell portable system with a 200 HB calibration bar. The hardness results indicated 17 locations of relatively low hardness (ranging from 168 to 189 HB) on the ASTM A182 Grade F91 material.

In 2011, metallurgical replicas were performed on Girth Welds 13 through 18. Examination of the Grades F91-F91 replicas (Girth Welds 14 through 16) revealed very subtle carbide denuded regions adjacent to the weld fusion line, typically in the range of 0.08 to 0.13 mm (0.003 to 0.005 inch) wide. These zones were formed by diffusion during post weld heat treatment and are an indication of excessive temperature exposure. The lack of observed creep cavities was evaluated with the understanding that in Grade 91 material, the creep cavities evolve by growing in size, but do not become aligned or oriented until a large fraction of creep life (typically 95% or more) has been consumed [1]. In addition, the creep void density at the pipe outer surface is usually considerably less than at subsurface locations.

3. Methodology It is assumed that the severe thermal quench events that had occurred during fan-cool shutdowns have been

eliminated. The general methodology for time dependent multiaxial stress estimates and weldment life consumption is discussed below.

3.1 Time Dependent Multiaxial Stress Estimates The high energy piping life consumption (HEPLC) methodology was used to estimate the time dependent

multiaxial stresses. This evaluation includes the redistribution of stresses due to dead weight, pressure, external loading, and thermal loading. The evaluation of redistributed stresses is especially important when nonideal

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conditions, such as malfunctioning supports, exist in the MS piping system. In 1998 and 2000, Cohn and Yee [2, 3] used three-dimensional four node quadrilateral shell elements in the ABAQUS [4] general finite element code to evaluate the time dependent redistributed stresses in a portion of a high energy piping system. Since the ABAQUS evaluation considered a thin-walled pipe with a diameter to thickness ratio of 23 and no geometry discontinuities, the five integration points through wall were considered adequate. The redistributed stresses were evaluated to develop close approximations to the fully elastic and fully redistributed stresses using the CAEPIPE [5] flexibility analysis software.

As discussed by Cohn [6], the long range elastic stresses (defined as the multiaxial stress at time equals zero) are estimated using the as-found piping stress analysis results for the operating stress (dead weight, pressure, external and thermal loads) in the direction axial to the pipe and the Lamé equations [7] for the other two orthogonal directions. The total elastic stress in a weldment is estimated as the combination of the applicable initial base metal stress, residual stress, and weld creep performance factor. The HEPLC also considers the inelastic stress (defined as the fully redistributed stress). The inelastic stress is estimated using the results of the as-found piping analysis sustained load stress analysis in the direction axial to the pipe and the Bailey equations [8] for the other two orthogonal directions. The total inelastic stress in a weldment is estimated as the combination of the applicable fully relaxed base metal stress, residual stress, and weld creep performance factor after creep relaxation. Time dependent stress redistribution of ferritic steels is estimated using the Norton equation [9]. Applicable values for the empirically derived coefficients are based on experimental results for the appropriate material and temperature range. At this point in the process, the methodology to relax the stresses is based on uniaxial test results. As a final step, the uniaxial stresses are adjusted to multiaxial stress conditions by using the ASME Code Case N47-32 [10].

3.2 Weldment Life Consumption Once the effective weldment stress is determined as a function of time, cumulative life consumption for each weldment is calculated for the current unit operating hours. Empirically derived and well characterized stress rupture curves have been determined for the appropriate materials. The Robinson linear life fraction rule [11] is then used to determine the incremental life consumed for the applied stress at each incremental time step. An integrated life consumption value is estimated for each girth weld at its current operating hours. Consequently, critical welds are ranked according to their predicted life consumption due to creep damage. If there are many cold starts for the piping system, an additional set of welds are ranked for possible fatigue damage, considering the as-found thermal expansion multiaxial stresses. This HEPLC methodology has been successfully used over the last 10 years on more than 150 MS or hot reheat piping systems. Using this methodology, successful predictions of three weldment creep failures have been published by Cohn [12]. An additional seven weldment creep failure times have been closely estimated using this approach and each of these weldments has been within the top five ranked weldments for the specific pipe line.

4. Application to Low Hardness Grade 91 Material The application of the general methodology to this specific MS piping system includes discussions of ASME

B31.1 Code piping stress analyses, determination of the time dependent multiaxial stresses, evaluation of applicable creep rupture data, and weldment life consumption.

4.1 Stress Analyses of the MS Piping System An as-designed piping stress analysis was performed for baseline information. The American Society for

Mechanical Engineers (ASME) B31.1 Code sustained load (SL) stresses for this piping system ranged from 21 to 59 MPa (3.0 ksi to 8.6 ksi). The as-designed ASME B31.1 thermal expansion (range) stresses (SE) for this piping

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system ranged from 5.5 to 46 MPa (0.8 ksi to 6.7 ksi). All of the as-designed ASME B31.1 stresses were within the Code stress allowables.

Detailed hot and cold walkdowns in 2004 revealed that thermal expansion displacements of this piping system did not behave as predicted in the design analysis. As illustrated in Figure 3, the field thermal displacements at several supports were 51 to 76 mm (2 to 3 inches) less than expected. Consequently, an as-found piping stress analysis was performed to simulate the effects of the malfunctioning supports (five supports were topped-out off-line). As illustrated in Figure 4, the as-found piping thermal displacements correlated well with the observed field thermal displacements, especially for the pipe thermal displacements at Supports MS-2, MS-4E, MS-4W, MS-6E, MS-6W, MS-31, and MS-7. Considering the 2004 observed displacements, the as-found SL stresses ranged from 15 to 56 MPa (2.2 to 8.1 ksi) and the SE stresses ranged from 3.4 to 43 MPa (0.5 to 6.3 ksi). All of the as-found ASME B31.1 SL and SE stresses were within the Code stress allowables. The recommended spring hanger corrective actions for the five malfunctioning supports were implemented some time after 2004.

There were no hot and cold walkdowns performed in 1998 to directly correspond to the observed Weld 15 crack. However, life consumption evaluations of the 2004 set of hot and cold walkdowns and as-found piping stress analyses revealed that Weld 15 ranked #1 and Weld 16 ranked #2. The third ranked weldment was Grade 11 material, with significantly less consumed life as compared to Weld 15.

Detailed hot and cold walkdowns were also performed in 2010. After hanger adjustments, the 2010 as-found SL stresses ranged from 10 to 31 MPa (1.5 to 4.5 ksi) and the SE stresses ranged from 12 to 34 MPa (1.8 to 4.9 ksi). These 2010 as-found piping stresses are considerably less than the 2004 as-found piping stress results. At the Grade P91 circumferential welds, the 2010 calculated stresses also considered field measured minimum pipe wall thicknesses. 4.2 Time Dependent Multiaxial Stresses

The MS piping system materials are subject to creep during normal operation. It is well documented that the time-dependent creep phenomenon results in redistributed stresses and degraded material properties (e.g., ultimate tensile strength, yield strength, and creep rupture strength). The elastic and inelastic throughwall circumferential stresses due to pressure for Girth Welds 15 and 16 are illustrated in Figure 5. At the outside diameter (OD) surface, the inelastic circumferential stress is about 17.5% greater than the elastic circumferential stress and about 6.7% greater than the elastic midwall circumferential stress. The elastic and inelastic throughwall axial stress due to pressure for Girth Welds 15 and 16 are illustrated in Figure 6. At the OD surface, the inelastic axial stress due to pressure is 17.5% greater than the elastic axial stress and the elastic midwall stress.

At Girth Weld 15, the elastic maximum principal stress in the weldment at the outside surface of the pipe is estimated at about 62.5 MPa (9.1 ksi). This value for the weldment stress includes a time dependent weld creep performance factor (calibrated to historical weldment failures and increases over time) and a factor for residual weldment stress (which decreases over time). The effective weldment inelastic (fully redistributed) maximum principal stress is estimated at about 88.0 MPa (12.8 ksi), about 72% greater than the initial as-designed Code sustained load stress. This weldment stress value also includes time dependent weld creep performance and residual stress factors. It was determined that at these stresses and an operating temperature of 541°C (1005°F), the maximum principal stress has not fully redistributed at 100,000 operating hours, so the maximum principal stress is 84.9 MPa (12.3 ksi). Consequently, the weldment effective stress redistribution curve from 10 to 100,000 hours is illustrated in Figure 7. 4.3 Evaluation of Soft vs Normal Grade 91 Materials

Grade 91 base metal creep rupture data from a variety of sources was compiled for creep rupture tests in the temperature range of 932°F to 1292°F (500°C to 700°C). As a second step, Grade 91 creep rupture data were also compiled from technical papers by Kong, Shingledecker, Brett, and Ryu [13-16] for specimens with base metal and weldment anomalies (e.g., low hardness, high hardness, and low N with high Al). The multiple sets of creep rupture data are presented in Figure 8, considering log stress vs LMP (using a constant coefficient of 30). A conversion of hardness scale [17] for this figure is provided in Table 1.

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This figure indicates that the API 579 Grade 91 minimum curve [18] is reasonable for the basic set of Grade 91 base metal creep rupture data, but the added Grade 91 weldment anomaly creep rupture properties fall significantly below the API 579 Grade 91 minimum curve.

At temperatures in the range where creep and stress rupture strength govern the ASME selection of stresses, the maximum allowable stress value of Grade 91 material (566°C to 649°C (1050°F to 1200°F)) is the minimum value of the following three criteria:

(1) 100% of the average stress to produce a creep rate of 0.01%/1,000 hours, (2) 2/3 of the average stress to cause rupture at the end of 100,000 hours (3) 80% of the minimum stress to cause rupture at the end of 100,000 hours Below 1050°F, the ASME allowable stress criterion is calculated as the ultimate tensile strength/3.5. With

the understanding that the ASME B31.1 Code (Code) [19] does not provide time-dependent design stress allowables, the Code Grade P91 stress allowables at 100,000 hours are also indicated in Figure 8. This figure reveals that at 100,000 operating hours, the Code Grade P91 stress allowable curve captures all of the basic base metal data and most of the low hardness Grade P91 data, but has no safety margin for soft weldments.

However, it is not the intent of the Code that the Code stress allowables be used for creep life estimates. For this paper, a lower bound creep rupture curve was generated slightly below the Grade 91 weldment anomaly creep rupture data points. As illustrated in Figure 9, this lower bound curve has a similar slope to the API 579 and ASME B31.1 curves at applied stresses below the available data. 4.4 Low Hardness Grade 91 Weldment Life Consumption

The superheat outlet header (SHOH) has a design pressure of 14.96 MPa (2,170 psig) and a design temperature of 541°C (1005°F). A review of the plant information data indicated that the truncated average operating pressure is about 13.03 MPa (1,890 psig) and the maximum operating temperature is about 541°C.

The unit is operated about 50% of the time in a standby (low power) mode, which is about 60% of the truncated average operating pressure. This operation is typically called on-load cycling. During this time operating at significantly lower pressures, the life consumption fraction is negligible. From a flexible operation point of view, this is especially interesting because it indicates that any increase in damage resulting from the low load and return to full load cycles is more than offset by the dramatic increase in life associated with the time durations spent at the low load state.

Using the above information and the Grade 91 weldment anomaly creep rupture properties (i.e., lower bound creep rupture curve in Figure 8), incremental weldment life consumption evaluations were performed from the time Girth Welds 15 and 16 were excavated and replaced in 2001. Note that Girth Weld 14 (the upstream wye weldment) was excavated and repaired in 2009. The integrated weldment life consumption evaluation indicated that less than 12% life had been consumed at each of the Grades F91-to-F91 weldments and the remaining lives were at least 20 years if the unit is operated in the future as it had been operated in the past.

The Figure 8 lower bound stress rupture curve indicates that at these operating temperatures and pressures, a 6 to 7% increase in stress results in 50% creep life. Therefore consideration of OD surface inelastic axial and circumferential pressure stresses, as-found pipe displacements, time dependent residual stresses, weldment creep performance factors, and creep redistribution can result in significantly higher weldment effective stress estimates and consequently lower estimates of time to failure. Conversely, the lack of consideration of these significant attributes can result in considerably optimistic time to failure estimates.

5. Conclusions This paper provides a life consumption study on the evaluation of low hardness Grade 91 weldments. The

study reveals that API 579 and ASME B31.1 Grade 91 creep rupture properties are not conservative for the minimum low hardness Grade 91 creep rupture data. It was determined that the creep rupture properties of low hardness Grade 91 weldments can be conservatively evaluated using a lower bound curve on a log stress vs LMP plot.

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In 1998 a surface connected creep crack was found in a Grade F91 wye block to F91 pup transition spool weld. This Weld 15 (on the west leg of the turbine lead piping) was excavated and repaired at that time. In 2001, Weld 15 and Weld 16 (on the east leg of the turbine lead piping) were excavated and repaired. A set of hot and cold walkdowns in 2004 included as-found piping stress analyses and life consumption evaluations of all the piping weldments. Considering several malfunctioning supports at that time, it was determined that Weld 15 had a life consumption Rank #1 and Weld 16 had a life consumption Rank #2.

The as-found piping stress analyses in 2010 indicated that corrective actions on several of the MS piping malfunctioning supports reduced the SL stresses by more than 40%. The life consumption evaluations for the Grade F91-to-F91 weldments indicated that less than 10% life had been consumed at each of the weldments and the remaining lives are at least 20 years if the unit is operated in the future as it has been operated in the past. At an operating temperature of 541°C and 100,000 operating hours, it is possible that the weldment stresses have not been fully redistributed. The lack of consideration of life consumption attributes such as OD surface inelastic axial and circumferential pressure stresses, as-found pipe displacements, time dependent residual stresses, weldment creep performance factors, and creep redistribution may result in considerably longer (overly optimistic) estimates of time to failure.

This case study also demonstrated that 1) troubleshooting monitoring can be effectively used to identify and eliminate unanticipated condensate quench events and 2) low load operation may have a beneficial effect on the longevity of high energy piping system welds.

6. Acknowledgments The support of Intertek APTECH is greatly appreciated and acknowledged.

7. References [1] Isamu Nonaka, Takuya Ito, Fumio Takemasa, Kensuke Saitou, Yoshikazu Miyachi, Akigo Fujita, 2007,

“Full Size Internal Pressure Creep Test for Welded P91 Hot Reheat Elbow,” International Journal of Pressure Vessels and Piping, Volume 84, pp. 97-103.

[2] Cohn, M. J., and Yee, R.K., 1998, “Creep Relaxation Behavior of High Energy Piping,” ASME Pressure Vessels and Piping Conference, San Diego, CA, PVP-Vol. 380, pp. 135-150.

[3] Cohn, M. J., and Yee, R.K., 2000, “Creep Relaxation Behavior of High Energy Piping,” Transactions of the ASME, Journal of Pressure Vessel Technology, 122, pp. 488 – 493.

[4] Hibbitt Karlsson & Sorenson, Inc., 1997, ABAQUS Finite Element Code, Version 5.6, Providence, Rhode Island.

[5] SST Systems, Inc., 1995, “CAEPIPE Computer Code User’s Manual,” Revision 20, Version 3.74. [6] Cohn, M. J., 2006 “A Strategy for Life Management of Main Steam and Hot Reheat Piping Systems,” EPRI

International Conference on Advances in Condition and Remaining Life Assessment for Fossil Power Plants, Louisville, Kentucky.

[7] Lamé 1852, “Lecons sur la théorie … de l’élasticité,” Gauthier-Villars, Paris. [8] Bailey, R. W., 1956, “ Creep Relationships and Their Application to Pipes, Tubes, and Cylindrical Parts

Under Internal Pressure,” Proceeding, Institute of the Mechanical Engineers, 164. [9] Norton, F. H., 1929, Creep of Steel at High Temperatures, McGraw-Hill, New York, p.67. [10] ASME, August 1994, “Class 1 Components in Elevated Temperature Service, Section III, Division I,” Cases

of ASME Boiler and Pressure Vessel, Code Case N47-32. [11] Robinson, E. L., July 1952, “Effect of Temperature Variation on the Long-time Rupture Strength of Steels,”

Transactions of the American Society of Mechanical Engineers, 74, pp. 777-781. [12] Cohn, M. J., 2006 “Life Management Projects for Three Main Steam Piping Systems,” EPRI International

Conference on Advances in Condition and Remaining Life Assessment for Fossil Power Plants, Louisville, Kentucky.

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[13] Kong, B.O., Kin, J.T., Lee, Y.S., and Ryu, S.H., 2007, “Thermal Histories Causing Low Hardness and the Minimum Hardness Replacement in a Mod.9Cr1Mo Steel for Boiler,” EPRI International Conference on Boiler Tube and HRSG Tube Failures and Inspections, Calgary, Alberta, Canada (October 16-18, 2007).

[14] Shingledecker, J. P., Santella, M. L., and Kleuh, R. L., 2007, “Evaluation of Heat-Treatment Temperatures and Corresponding Properties of Improperly Heat-Treated Grade 91,” Proceedings of Industry and Research Experience in the Use of P/T 91 and Other New Steels, European Technology Development (ETD), London, UK (June 20-21, 2007).

[15] Brett, S. J., Bates, J. S., and Thomson, R. C., 2004, “Aluminum Nitride Precipitation in Low Strength Grade 91 Power Plant Steels, Proceedings to the Fourth International Conference on Advances in Materials Technology for Fossil Power Plants, Hilton Head, SC (October 25-28, 2004). ASM-International, Materials Park, OH, 2005. 1183-1197.

[16] Ryu, S. H., Lee, Y. S., Kong, B. O., and Kim, J. T., 2002, “A Study on the Variation of the Hardness and the Creep Rupture Strength with Thermal Histories in a Mod. 9Cr-1Mo Steel,” First Int. Conference on Advanced Structural Steels (ICASS2002,) Organized by NIMS, Tsukuba, Japan (May 22-24, 2002).

[17] B.S. 860, 1967, “Tables for Comparison of Hardness Scales,” British Standards Institute. [18] API 579-1/ASME FFS-1, 2007,Second Edition, June 2007, Table F.31. [19] ASME, 2012, “ASME B31.1-2012 Edition, Power Piping,” ASME Code for Pressure Piping, B31, An

American National Standard, The American Society of Mechanical Engineers, New York., NY.

Table 1

Conversion of Hardness Scales (From BS 860/1967)

Diamond Pyramid Scale HV10   HV30

Brinell Standard Ball (HB)

421 400400 380263 250214 200195 186189 180180 172175 167162 155155 148

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Figure 1. Illustration of the Main Steam Piping System

Figure 2. Illustration of the Replaced Wye Fitting and Adjacent Welds

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Figure 3. As-Designed Piping Displacement Profile

Figure 4. As-Found Piping Displacement Profile

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Figure 5. Circumferential Stress Due to Pressure Elastic and Inelastic

Figure 6. Axial Stress Due to Pressure Elastic and Inelastic

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Figure 7. Weldment Effective Stress Redistribution Curve

Figure 8. Soft vs Normal Grade 91 Materials

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Figure 9. Soft Grade 91 Materials with Lower Bound Stress Rupture Curve

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