HIGH PRESSURE STEAM TURBINE ROTOR FORGING

SECTION REPLACEMENT DUE TO SEVERE CREEP DAMAGE

Robert E. Kilroy Jr. Director of Engineering Mike Jirinec Welding Engineer Damian Parham Turbine Repairs Design Engineer ALSTOM Power Inc. 1200 Willis Rd. Richmond, Virginia 23237

Steve Schrickel Wisconsin Public Service Corporation 1530 N. Bylsby Avenue Green Bay, WI 54307

Don McCann ReGENco, LLC 6609R West Washington Street West Allis, WI 53214 Darryl Rosario Structural Integrity Associates, Inc. 3315 Almaden Expressway, Suite 24 San Jose, CA 95118

Presented at the:

Ninth EPRI Steam Turbine / Generator Conference and Vendor Exposition August 22-24, 2005 Denver, Colorado

Printed by Alstom Power Inc., Turbine and Generator Repair Engineering, Richmond, VA

HIGH PRESSURE STEAM TURBINE ROTOR FORGING SECTION REPLACEMENT DUE TO SEVERE CREEP DAMAGE

Steve Schrickel

Wisconsin Public Service Corporation 1530 N. Bylsby Avenue Green Bay, WI 54307

Mike Jirinec Rob Kilroy

Damian Parham ALSTOM Power, Inc.

1200 Willis Road Richmond, VA 23237

Don McCann

ReGENco, LLC 6609R West Washington Street

West Allis, WI 53214

Darryl Rosario Structural Integrity Associates, Inc.

3315 Almaden Expressway, Suite 24 San Jose, CA 95118

Abstract

During a planned outage in 2004, magnetic particle examination of the HP rotor found evidence of cracking in the balance plane radius. Metallurgical examination revealed the presence of severe creep cracking which led to the rotor being condemned by ReGENco for operation at rated temperature. Review of the inspection results, the performance of material analysis, and the testing of creep rupture rotor material samples for correlation allowed the rotor to be put back into service temporarily at a de-rated output and temperature until a repair was decided upon. In 2005, the unit was again brought down and the TE forging section was replaced and a new forging was join welded to the rotor between the Curtis stage wheel and the first reaction stage. Details of the initial findings, material analysis, and de-rated operation rational as well as the rotor repair rational, repair sequence, welding testing, engineering analyses, and project execution will be discussed and presented in detail.

Unit Background Information

Wisconsin Public Service Corporation (WPCS) J.P.Pulliam Unit # 6 is a 60MW Westinghouse Unit. The unit consists of a tandem-compound high-pressure (HP) rotor, and a double- flow (DFLP) low-pressure rotor. Design inlet conditions are 850 psig and 950°F. This unit was placed in service in 1951 and had accumulated approximately 347,000 operating hours at the end of 2004. The unit was designed for base load operation and continues to operate in this mode today.

Damage / Indications Found

Inspections

During a spring outage in 2004, the HP rotor, Figure 1 was inspected using ultrasonic (UT) and magnetic particle (MT) examination methods. In preparation for the boresonic examination, the bore diameter measured along the length (~170”) of the rotor before honing revealed a bore enlargement beneath the high-pressure and low-pressure balance pistons (HPBP & LPBP) at 124"-138" from the generator end of about 0.23%. The enlargement was checked against the 1994 bore measurements and was found to be real, indicating that creep deformation had occurred in service. The rotor was then boresonic (UT) inspected and numerous UT indications were recorded about 4"-6" radially from the bore surface at the large transition radius of the LPBP face (135"-137.5") as shown in Figure 2, which displays 90"-138" of the rotor. Magnetic particle examination (MT) was performed on the rotor periphery whereupon 63 MT indications were reported on the surface of this transition radius. Visually the indications appeared to be opened cracks in the circumferential-radial plane extending 360° around the circumference of the radius. The center of the 1.3" wide radial band of cracks was about 5.1" radially from the bore surface. The largest MT indication was 0.272" long. Metallurgical replicas revealed that the MT indications were creep cracks, Figure 3. It was then decided to machine 0.187" from the transition radius to determine if the cracks would disappear, and to remove a ring sample for mechanical properties tests. After MT, 72 cracks were recorded; the five largest cracks were between 0.262" and 0.395" long. There was good correlation between UT and MT indications.

Figure 1. HP Rotor Configuration

Figure 2 . HP Rotor UT Data from 90” to 137.5”

AXIAL VS RADIAL PLOTUltrasonic Inspection Data

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WPS Pulliam 6 HP Rotor U191c (90"-138") April 2004

Theta Range: 0 to 360°16-Apr-04Generator end C/L #1 Balance Hole

AXIAL VS CIRCUMFERENTIAL PLOTUltrasonic Inspection Data

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WPS Pulliam 6 HP Rotor U191c (90"-138") April 2004

ReGENco Generator end C/L #1 Balance Hole 16-Apr-04 Depth Range: 0" to 10"

Figure 3. MT Indications in LPBP Transition Radius

Mechanical Property Tests

Essentially the rotor was manufactured from a 2.54%Ni-0.46%Mo-0.04%V alloy steel with only 0.06% Cr. In addition, the sulfur was 0.040% and the phosphorus was 0.032%. The yield and tensile strengths were 77.1 ksi and 99 ksi, respectively. The FATT was 300°F, and the Charpy V-notch energy was 4.5 ft-lbs @ 70°F, 27 ft-lbs @ 300°F and 34 ft-lbs @ 500°F. Temperature and pressure data were recorded by WPSC between 1/1/2001 and 3/12/2004. The temperatures recorded in the exhaust steam line at the HPBP and LPBP were as follows:

HPBP, °F LPBP, °F Minimum 812.04 802.48 Average 882.12 879.01 Maximum 925.26 919.93

Hence, the average temperature range at the BP was probably about 879-882°F. Stress Analysis

The steady state running speed (3600 rpm) stress profile was calculated about the bore of the HPBP and LPBP using the finite element program COSMOS/Works. Stress analyses were made of the rotor in its original condition, and after the ring sample was removed and the radius was re-contoured. Plots of the circumferential and radial stresses are shown in Figures 4 and 5 for the original design. Note that the radial stress at the transition radius was about 35.6 ksi whereas the circumferential stress at the bore was only about 24.9 ksi. This explains why the creep cracks initiated at the transition radius and not at the bore. After the machining modification both stresses decreased a small amount. Fortunately the radial stress decreases axially inward from the radius. For example, at an axial distance of 0.4" the radial stress was about 23 ksi, which indicates that cracks should not develop across the entire LPBP due to a lower stress field.

Creep Cracking

Creep Test Results

A condition assessment of the UT and MT data revealed that the days for continued operation of this rotor in its present condition were limited. To continue in operation in its present condition for a limited amount of time, either the rotor stress level or temperature must be reduced to mitigate creep crack growth. Since the stress level cannot be adequately reduced, the temperature must be reduced. For long-term continued operation of the rotor, it was recommended that the rotor be cut between the impulse wheel and the 1st reaction row, and that a new forging (CrMoV with no bore) be join welded on. However, to allow preparation time for welding the new partial forging to the old rotor; it was recommended that the rotor could be returned to service, but with a lower operating temperature at the LPBP. Since creep crack growth data were not found for this alloy, actual creep rate tests were conducted to get the creep rate as low as possible; specimens were tested at 850°F/30 ksi, 850°F/25 ksi, 800°F/30 ksi, 800°F/25 ksi, and 750°F/30 ksi. The stress levels were based on the radial and circumferential stress results. All initial creep rate tests were run at least 1000 hours. Based on these results it was decided to limit the maximum temperature of the LPBP to 800°F, where the estimated creep rate would be about 0.000065 % per hour. However, the 800°F/30 ksi and 750°F/30 ksi creep rate tests were continued as long as the rotor was in operation to make sure the creep rate did not suddenly increase.

Figure 4: Circumferential Stress at HPBP And LPBP for 3600 rpm

Figure 5: Radial Stress at HPBP and LPBP for 3600 rpm Creep testing results allowed the unit to be returned to service at lower operating temperature conditions for a specified period of time. During this time frame, further analysis showed that a join weld repair represented the best economical and schedule solution and also provided for a permanent repair. This repair involved replacement of the shaft end section and resulted in a sever grove being located between the impulse stage and first reaction blade row. A new shaft end section was reinstalled using deep groove submerged arc welding. The following describe the results of this forged end section replacement.

Join Weld Repair Concept

There are two basic options and/or concepts utilized by ALSTOM in the axial join welding of rotor forging sections; (i) join welding solid forging sections over hollow cavities, and (ii) join welding bored forging sections utilizing piloted bores. The method of join welding solid forging sections utilizes shaft end weld preps that form a hollow cavity after assembly of the two forgings. New turbine rotors have been manufactured by ALSTOM utilizing this method for more than 70 years, see Figures 6 & 7. The method of welding bored forging sections, regardless of whether both forgings have bores or only one forging has a bore, utilizes bore pilots or spigots to ensure alignment of the bores when the two forgings are assembled. Damaged and condemned rotors have been repaired by ALSTOM utilizing this method for more than 30 years.

Figure 6. Model of Standard Solid Forging Join Welding Configuration

Figure 7. Forging Sections Prior to Join Welding

The Pulliam rotor had a bore; therefore the second method of join welding was utilized. The new forging section used to replace the damaged section was a solid forgings without a bore and therefore the final join welded rotor configuration consisted of a blind bore once completed. The HP rotor sever and join weld location would be between the Curtis Stage and the first reaction row, Figure 8.

Figure 8. HP Sever Location and Proposed Join Weld Repair

Join WeldOriginal Rotor SectionNew Forging Section

Proposed Repair Process

The proposed repair process for the HP rotor, new forging, and removed shaft end included the following scope of work: HP Rotor Work • Non-destructive removal and tenon refurbishment of R1A & R1B impulse blades • Non-destructive removal and replacement of reaction row 1-3 blades • Reverse engineering of rotor • Sever removal of damaged rotor section • Face butter layer welding • Shaft end join weld prep machining & NDE • Join weld attachment of new forging section • Post weld heat treatment • Join weld and base material HAZ NDE • Final machining of new forging section • Reassembly of originally removed reaction row and impulse blades • Low Speed Balancing New Forging Work • Rough machining of new forging section (already in stock) • Weld buildup of new forging section in Impulse and dummy piston areas (forgings

were undersized in diameter at these locations • Post weld heat treatment • Weld buildup and base material HAZ NDE • Forging end join weld prep machining & NDE Remove Shaft End Work • Shaft ends join welding groove machining & NDE • Welding of shaft end section • Post weld heat treatment • Base, weld, and HAZ material sample machining removal • Metallurgical testing of samples

The proposed solution for the HP rotor included the forging section removal and replacement as shown in Figure 8 above. Although the design, configuration, and analysis of joining welds have been tested and proven, it was requested by the customer to perform additional analysis. In parallel to the rotor repair, the severed and removed original shaft end material was welded and mechanical properties established. This testing provided information in regards to the weldability of the current rotor forging, as well as the actual mechanical properties to be utilized in the design analysis of the rotor final bore and join weld configuration.

To ensure a repair is performed to mutually acceptable standards of quality and that these standards are met, monitoring of the repair by the customer, the insurer, or a designated third party is sometimes performed. In the case of this HP rotor weld repair, multiple project witness points were defined and agreed upon by WPSC prior to proceeding with the repair. Due to the critical nature of the repair schedule no hold points were imposed during the repair, only witness points. These witness points agreed to were as follows: • Welding Testing and Weld Material Property Results • Replication Evaluation and Results • Join Weld Repair Analysis and Evaluation Results • Join Welding • Join Weld and HAZ NDE • Final Machining Dimensional Inspection • Low Speed Balance

Base Material Welding and Metallurgical Testing

Chemical composition of the HP rotor was required prior to performing the joining weld operations to ensure compatibility of the base metal and weld filler metal. The results of this analysis are presented in Table I. The composition met the respective original ASTM A293 specification requirements, but had significantly high levels of sulfur. This data suggested that there might be an issue with the joining weld on HP rotor.

Table I Chemical Analysis of HP Rotor

Element ASTM A293 (Grades 4,5&6 – wt %)

Actual HP Rotor (wt. %)

Carbon 0.45 max. 0.22 Manganese 1.00 max. .0.64

Phosphorous 0.05 max. 0.040 Sulfur 0.05 max. 0.032 Silicon 0.15 min. 0.21

Chromium 1.25 max. 0.06 Nickel 2.25 min. 2.54

Molybdenum 0.50 max. 0.46 Copper N.S. 0.13

Aluminum N.S. 0.006 Niobium N.S. <0.001

Vanadium 0.12 max. 0.04 Tin N.S. 0.006

Antimony N.S. 0.001 Arsenic N.S. 0.007

To successfully weld repair a rotor, the properties of the base forged material must be identified. These include metallurgical condition (tensile & hardness) and toughness. The metallurgical condition of the rotor forging allows the repair organization to determine the appropriate post weld heat treatment temperature to assure that the rotor base metal properties are not degraded. The toughness of the material is to determine the ductile/brittle transition temperature to determine if any additional metallurgical or welding conditions need to be addressed. Table II lists the mechanical properties for the HP rotor and showed they met ASTM A293 specification requirements.

Table II Mechanical Properties of HP Rotor

Mechanical Properties ASTM A293 HP Rotor Actual Tensile (ksi) 100 95 Yield (ksi) 70 74 Elong. (%) 20 20 RA (%) 40 49 Impact (Ft. lbs) 7 7 – 8 Hardness (HRB) 225-250 205-210 Fracture Toughness (KIc ksi√in)– RT & 250°F

---

54 (138.7)

FATT 50 (°F) --- + 235°F

In many cases, there is not sufficient rotor material to produce a sub-size mock-up weld test using the same welding parameters and heat treatment as those to be applied to the actual repair. If a mock-up is possible, it can be destructively tested and the results evaluated and included in the final repair design analysis calculations. The testing on a mock-up weldment should be appropriate to determine the applicable mechanical properties for the in-service conditions of the equipment. In all cases it is paramount to determine the mechanical properties of the weld metal, heat affected zone, and base metal properties after final post weld heat treatment operations. In this case, the filler material as-welded mechanical properties were already known, but the weldability and heat-treated mechanical properties of the rotor base material was still needed. Although there was base material mechanical property information on previous similar rotor weld repairs including initial mechanical property data, this testing was still requested by the customer and included in the repair design analysis. Mock-up of the HP rotor section was prepared from the severed shaft end portion in the LPDP area. The test join welding was performed in an area just under the original dummy piston seals, Figure 9.

Figure 9. Weld Mock-up Groove Shaft End and Completed Weldment

Mechanical testing and metallographic examinations of the completed test welds were then performed. The mechanical properties of both welds and the adjacent heat-affected base metal were acceptable as shown in Table III. It should be noted, that due to the very poor base material mechanical properties, the HAZ mechanical properties improved, which is typically the case for these older forging materials. Figure 10 is a photo of the polished and etched mock-up groove joining weld showing the initial root weld passes, subsequent weld bead layers, and the base material HAZ.

Table III Mechanical Properties of Join Weld & HAZ

Mechanical Properties HP Rotor HP Weldment Tensile (ksi) 95 110 Yield (ksi) 74 98 Elong. (%) 20 21.5 RA (%) 49 70 Impact (Ft. lbs) 7 – 8 38-40 (HAZ) Hardness (HRB) 205-210 216-247 Fracture Toughness (KIc ksi√in)– RT & 250°F

54 (138.7) 200

FATT 50 (°F) + 235°F -32°F

Figure 10. Polished and Etched Mock-up Groove Joining Weldment

HP Rotor Section Preparation

In order to restore the rotor back to original geometry complete reverse engineering measurements of the rotor were taken. The distance to all axial features from the impulse wheel face to the coupling of the rotor were carefully measured and recorded. The cylindrical features of the ends to be removed were measured with pi tapes, the transition radii were measured with radius gauges, and the more complicated geometric features were measured with a portable coordinate measuring machine (CMM). Since the impulse blades and reaction row blades from the rotor were to be reused, the work began with the non-destructive removal of all the impulse and reaction row blading. The forging section was then severed between the impulse and first reaction row blades, Figure 11. The location of the cut was chosen based on weld access, operational stresses and temperatures. Since the rotor had sustained creep damage, it was advisable to locate the new joining weld in an area free from creep damage. In order to verify this, metallurgical replicas were taken on the severed and machined faced of the existing rotor surface. Analysis of the replicas showed that the microstructure of the parent rotor material in the cut location had a STAGE 1 material creep classification, Figure 12.

Figure 11. Rotor Sever Location (1st Reaction row Blades and Curtis Stage Blades)

Test Site 1

Test Site 2

Test Site 3

Test Site 5

Test Site 6

Test Site 8 Test Site 7 Test Site 4

Figure 12. HP Rotor After initial Prep Machining Showing Replica Test Sites and Corresponding Microstructures (Test Sites 1 & 3, Respectively)

This cut location was also chosen to ensure there was sufficient remaining material to re-attach the new forged end to re-establish the transition radius. In addition, preliminary stress analyses showed this to be a lower stress area. At this point UT examination of the rotor surface was performed and revealed no recordable indications. Because of the age of the rotor and sulfur levels, a butter layer weld build-up was applied prior to the actual join welding, Figure 13.

Figure 13. HP Rotor Butter Layer and Join Weld Layer

Join Weld Material

“Butter Layer” Weld Material

Figure 14. HP Rotor “Butter Layer” Final Machined

Figure 15. Final Machining of the Rotor Butter Layer Surface and Bore for Join Welding

After completion of the butter layer welding operation, the weld build-up was examined using MT & UT techniques, with no recordable indications noted. The weldment face and bore were then final machined and ready for join welding, Figures 14 & 15.

New Forging Section Preparation

The replacement forging section utilized was an ALSTOM ST460TS material, which was originally purchased for a gas turbine high temperature forging section. The forgings were measured during the initial phase of the project and found that they were an

acceptable size in all areas except the Curtis stage and dummy piston areas and required weld build-up, Figure 16. This work was all accomplished prior to the arrival of the rotor. After completion of the welding, the weld build-up was subjected to a post weld heat treatment and non-destructive testing using MT and UT examinations. No recordable indications were observed. Prior to any rough machining and welding, the forging underwent the customary inspections and investigative material lab work to confirm material properties. The mechanical properties were acceptable and the forging sections was rough machined and weld prepped in parallel to the rotor work, Figures 16. Since this forging was solid and did not have a bore, the forging shaft end was machined with pilot fit as well as an oversized cavity to allow final bore machining after completion of the joining weld, Figure 17.

Figure 16. Forging Section After Weld Build-up and Rough Machining

Figure 17. New Forging Prep & Rough Machining Layout

Rotor / Forging Assembly and Join Welding

The rotor and replacement forged end sections were assembled with shrink fits and a welding groove is formed at their intersection Figure 18. The turbine rotor was set up in a submerged arc welding station in the horizontal position. It was supported on rollers at the turbine end and on temporary supports near the severed end. Next the new forging section was set up in the submerged arc welding station also in the horizontal position. It was supported on a second set of rollers. The forging support rollers required careful positioning with respect to the rotor forging to insure concentricity and alignment. If the concentricity were out of tolerance there would not be sufficient material on the new forging end to machine the diameters during final machining. In preparation for joining the two pieces, the end of the rotor was heated to enlarge the female shrink seat and the forging was cooled to reduce the diameter of the male pilot. Once the correct rotor and forging temperatures were reached the new forging was axially slid into position.

Figure 18. Rotor Shaft End Shrink Fit Bore

The axial position of the forging with respect to the rotor was established by the join weld prep interfaces at the bore and alignment was checked with three distance blocks evenly spaced around the rotor at the OD of the weld groove. The rotor forging and new forging were then allowed to equalize in temperature. The rotation of the assembly was started and welding pre-heat began while carefully monitoring the axial temperature gradient, the runout of the forging coupling diameter and the axial position of the forging with respect to the rotor, Figure 19.

Figure 19. Set-up of Turbine Rotor and Forging

A submerged arc welding process was used to join the forging to the rotor. Specially designed “bayonet” welding torches were required to hold the wire in the correct position in the deep, narrow groove. The inter-pass temperature and axial temperature gradient were carefully monitored during the welding process as the entire join weld groove was filled, Figure 20.

Figure 20. Set-up for Join Welding The post weld heat treatment (PWHT), also known as the stress relief cycle, was performed with the rotor in the vertical position with resistance heating, Figure 21. The temperature was increased at a defined ramp up rate, held in a carefully defined soak range, and decreased at a defined ramp down rate. In addition to monitoring the temperature in the soak zone the axial temperature gradients along the rotor were measured and recorded as well.

Figure 21. Post Weld Heat Treatment

Non-Destructive Testing of Weldment

After completion of the welding and heat treatment operations, non-destructive ultrasonic examination of the weldment is performed to determine the acceptability of the weldment for service. The ultrasonic examination requires multiple scan directions to assure complete coverage of the weldment. In support of the ultrasonic examination the rotor was moved to a large lathe and the welded area was machined to a smooth surface, Figure 22. Also, final machining of the bore is completed at this time to provide a smooth bore surface matching the existing bore surface for NDT examinations. The design of the piloted bore allows the pilot features to be machined away along with the first several weld root passes. An ultrasonic inspection of the weld and adjacent area from at least three angles revealed no unacceptable indications. Further wet magnetic particle testing of both the OD and the bore were performed to inspect for any weld fusion line cracking. No issues were identified.

Figure 22. Join Weld UT Prep and Inspection

For most weld repair applications, certain non-destructive metallurgical tests may be performed on the as repaired component as well. Such tests might include hardness, in-place metallography, and/or visual inspection. An example would be hardness testing of the deposited weld to ensure it was adequately heat treated to provide the required strength and hardness testing of the forging base material to ensure it was not softened and lost strength. Post repair metallurgical inspections performed on these rotors consisted of base material, filler material, and forging material visual, magnetic particle, and hardness inspections.

Join Weld Repair Evaluation

Since no standard formulas exist for evaluating weld repairs, each weld repair must be considered and evaluated on a case-by-case basis and not all of the evaluation requirements are derived solely from engineering considerations. Economics and safety are also high priorities to be considered and evaluated. While economic pressures can be high, they should never be allowed to obscure the engineering and safety considerations. The following are 12 standard evaluation factors involving metallurgical and mechanical engineering concerns, inspection requirements, and quality assurance issues. Not all are an absolute requirement for each weld repair scenario, but most are recommended. Each of these factors are discussed and/or defined throughout this paper and the join weld repair analysis and evaluation: 1. Failure Mechanism Determination 2. Perform a complete rotor inspection including a boresonic inspection 3. Determine base material chemical composition 4. Determine base material mechanical properties 5. Perform mechanical and metallurgical tests of a mock-up weld 6. Perform classical engineering mechanics calculations 7. Perform an FEA of the welded design, including SAFER Analysis 8. Perform an ultrasonic inspection of deposited weld 9. Perform a quality assurance audit of the repair process 10. Perform metallurgical testing of the welded component 11. Monitor operational vibration 12. Define a post repair inspection plan In respect to this repair solution for the HP rotor, the proposed join welded forged end section replacement was analyzed to determine stress levels and acceptability for service. The analysis took into consideration the redesigned bore, weld prep configuration, as well as the base material, forging material, filler metal material, and HAZ mechanical properties. Although the concept of the repair was known and agreed to up front, the exact final configuration was unknown due to the unknown material condition at the cut location. Therefore the analysis and final repair solution evolved as the project progressed. The evaluation of the repair options for the HP rotor performed by Structural Integrity Associates and reviewed by Alstom Power, included the following phases: i) Evaluation of the proposed final rotor to forging joining configuration, ii) Evaluation and determination of allowable welding flaw sizes in various areas of

the repair.

Initial analysis concentrated on the stresses due to rotational inertia for both the original rotor and proposed repair configuration, using the ANSYS software program. FE hoop stress results at 3600 rpm are shown in Figures 23a and 23b for the original and repaired

condition. The highest hoop stress of 28.8 ksi was predicted at the bore under the last (L-0) stage blades. The next highest hoop stress location (~ 25 ksi) was predicted at the bore under the dummy piston of the original rotor; this bore hoop stress reduced to about 16 ksi in the boreless new forging. The join weld location was selected at a relatively low stress location with a maximum bore stress of 14.1 ksi; this stress is lower than the 15.6 ksi stress at the same location in the original rotor. First principal stress results are shown in Figures 24a and 24b for the original and repaired condition. The maximum first principal stress predicted for the original rotor was 38.7 ksi at the dummy piston exhaust face fillet radius, consistent with location of the observed creep cracks. In the boreless repaired design, this stress increased very slightly to 39.9 ksi. From the standpoint of rotational stresses, therefore, there is no significant change in stress for the repaired design compared with the original rotor. The use of a boreless (solid) forging does however reduce the near-bore hoop stress under the dummy piston. Overspeed: Stress at the 30% design overspeed can be calculated by using a scale factor = (1+0.3)2 = 1.69, over the stresses at rated speed (3600 rpm), because rotational stresses are proportional to the square of the rotor speed.

Figure 23a. - Original Rotor: Hoop Stresses at 3600 RPM

Figure 23b. - Repaired Rotor: Hoop Stresses at 3600 RPM

28.8 ksi

25 ksi

15.6 ksi

Figure 24a. - Original Rotor: First Principal Stresses at 3600 RPM

Figure 24b. - Repaired Rotor: First Principal Stresses at 3600 RPM

The stress analysis module of the SAFER code was used to determine the distribution of stress and temperature in the rotor body during startup and loading. These stress and temperature distributions are required for processing of near-bore indications and for input to fracture mechanics critical crack size and crack growth calculations. A description of the transient thermal and stress analysis is provided below. FE Model: Using the geometry data described above, a FE model of the repaired rotor was generated in SAFER as shown in Figure 25. This model includes the two Curtis Stages and all 23 reaction stages. Note that the SAFER code cannot exactly simulate the detail outside surface contour of the rotor that is possible with the ANSYS code nonetheless, the focus of the SAFER code is on determination of near-bore stress fields which is relatively unaffected by approximations to the OD surface profile.

Location of Creep Cracking at Max. Stress Location

Simulated Startup: HP rotor steam inlet/outlet conditions and RPM versus time data were provided for several representative startups/shutdowns. From this data it was assessed that the worst combination of high stress and low fracture toughness (corresponding to lower rotor body temperatures) would occur during a cold startup. Data for the selected cold startup is shown in Figure 26. This data was used to develop steam inlet/outlet and RPM profiles for input to the SAFER FE model, as shown in Figure 27. The initial rotor temperature was reported to be 85°F at the start of the transient. Appropriate physical properties for typical (CrMoV) rotor steels were specified for the analysis and heat transfer boundary conditions imposed on the periphery of the rotor were calculated by SAFER. Transient Stress Results: Stress results were processed to identify the worst combination of high stress and low fracture toughness, when brittle failure is most likely to occur. Estimation of fracture toughness versus temperature will be discussed along with fracture mechanics calculations in the next section of this paper. Predicted bore stress, temperature and fracture toughness versus elapsed startup time for the repair weld location is shown in Figure 28; results for the HAZ on the old forging side are shown in Figure 29. The contribution of thermal stress to the total stress state is also shown in these figures. The results show that the highest thermal stress is 16.4 ksi in the weld and 18.5 ksi in the HAZ occurring approximately 300 seconds after startup due to admission of hot steam on a relatively cold rotor (about 97ºF at the bore). The rotor RPM is less than 200 rpm, and the contribution of rotational stress at 300 seconds is therefore very small (less than 0.03 ksi). The worst combination of stress (18.5 ksi) and fracture toughness (28.7 ksi√inch) is therefore predicted for the HAZ on the old forging side. Color contour plots of hoop stress and temperature in the rotor at 300 seconds after startup are shown in Figures 30a and 30b.

X

YZ

Figure 25. - FE Model of the Repaired Rotor using SAFER

Control Stages 1, 2

Reaction Stages 1 - 23

Figure 26. - Steam Temperature, Pressure and RPM versus Time Data for a Cold Startup on 10/8/03

Figure 27. - Steam Temperature, Pressure and RPM versus Time Data for a Cold Startup Simulated

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Pulliam Unit 6 HP Rotor 10/08/03 Startup: Parameters Simulated in SAFER Analysis

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Tin (°F)Tout (°F)Pin (psig)Pout (psig)RPM

3600 RPM Tin=900 °F

Tout=390 °F

Pin=850 psig

Pout=52.5 psig

Start of Simulation10/8/03 4:15:00 AM

End of Simulation (Steady-State)10/08/03 8:35PM

217 °F/hr

80 °F/hr

2000 RPM

1500 RPM

Figure 28. - SAFER Bore Hoop Stress, Temperature and Fracture Toughness versus Time Results at

Weld Center for Cold Startup

Figure 29. - SAFER Bore Hoop Stress, Temperature and Fracture Toughness versus Time Results at Old Forging HAZ for Cold Startup

Weld Center (El.#548) Bore Stress & Temperature vs. Time

-5

0

5

10

15

20

25

-

3,60

0

7,20

0

10,8

00

14,4

00

18,0

00

21,6

00

25,2

00

28,8

00

32,4

00

36,0

00

39,6

00

43,2

00

46,8

00

50,4

00

54,0

00

57,6

00

61,2

00

64,8

00

68,4

00

Time (sec.)

Stre

ss (k

si)

0

100

200

300

400

500

600

700

800

900

Tem

pera

ture

(°F)

, Kic

(ksi√i

nch)

Tot. StressThermal StressDEG.FKic

Old Forging near Weld (El.#582) Bore Stress & Temperature vs. Time

-5

0

5

10

15

20

-

3,60

0

7,20

0

10,8

00

14,4

00

18,0

00

21,6

00

25,2

00

28,8

00

32,4

00

36,0

00

39,6

00

43,2

00

46,8

00

50,4

00

54,0

00

57,6

00

61,2

00

64,8

00

68,4

00

Time (sec.)

Stre

ss (k

si)

0

100

200

300

400

500

600

700

800

900

Tem

pera

ture

(°F)

, Kic

(ksi√i

nch)

Tot. StressThermal StressDEG.FKIC

Figure 30a. - SAFER Color Contour Hoop Stress Results 300 Seconds after Simulated Cold Startup

Figure 30b. - SAFER Color Contour Temperature Results 300 Seconds after Simulated Cold Startup

Linearized net-section membrane, and membrane+bending stresses for comparison with ASME Code allowables were reviewed. A 30% reduction in minimum material strength was applied to account for elevated temperature operation (up to 950ºF). Calculated maximum FE stresses for 30% overspeed are less than 30.5% of typical ASME Code allowables for this material.

16.4 18.5 (Weld) ksi ksi (HAZ)

Critical crack size, for a given set of applied loads, is the maximum tolerable crack size before unstable fracture occurs. Using linear elastic fracture mechanics principles, applied stress intensity factors (KI) were computed as a function of crack size (a), which then permit determination of a critical crack size (acr) for a rotor fracture toughness (KIc); i.e., critical crack size corresponds to an applied KI = KIc. KIc was estimated at 28.7 ksi√inch for the old forging base material. Stress Intensity Factor Calculations: Stress intensity factor (KI) versus crack size (a) calculations were performed using the stress intensity factor models and solutions from the SAFER-PC code. Through-wall stress gradients due to rotational and thermal effects are plotted in Figure 31. As discussed above, the maximum thermal stress at the weld (18.5 ksi) occurs very early on, when rotational stresses are insignificant, and, conversely rotor overspeed conditions are most likely to occur when the turbine is warmed up and thermal stresses are minimal. Therefore, the most limiting scenario is the 30% overspeed design limit (14.1 ksi x 1.32 = 23.8 ksi bore stress at the weld). These 30% overspeed stresses were used to perform the KI versus a calculations; results for a semi-elliptic surface connected flaw are shown in Figure 32. Crack Model: For critical crack size calculations, a range of elliptic surface-connected and buried flaw aspect (depth/length or a/l) ratios were evaluated. Flaw shapes and critical parameters are illustrated in Figure 33. Surface-connected flaws are most limiting because of higher stress intensity factors and high stress at/near the bore. An independent check of these SAFER KI versus the results was performed by comparison with “semi-elliptic axial crack in a cylinder” model results from SI’s pc-CRACK computer code. Critical Crack Size Results: Flaws in the HAZ material adjacent to the old forging (assuming the lowest estimated fracture toughness of 28.7 ksi√inch) are shown in Figure 34. For the weld material with an estimated room temperature fracture toughness of 56.8 ksi√inch, critical flaw sizes are significantly greater than those in Figure 34 (>4 inches radial depth, per the results in Figure 32). As reported by Alstom, the maximum permissible flaw for the weld is 1.5 mm (0.059 inch) in radial size. The minimum allowable radial size (a) for the most limiting (a/l=0.05) surface flaw is 0.529 inch, yielding a safety factor of (0.529/0.059)=8.96, or an allowable crack propagation margin of (0.529-0.059) = 0.47 inch. Based on the choice of join weld location which is downstream of the 2nd control stage, metal temperatures are expected to be below 850ºF where no significant creep crack growth is expected. Conservatively using upper bound fatigue crack growth data from the SAFER code at 850ºF: da/dN = 1.119E-08 ΔK 2.474, with a ΔK approaching the fracture toughness value of 28.7 ksi√inch, yields da/dN = 4.52E-05 inch/cycle. For a projected 20 starts (and 8200 hours) per year, this equates to 0.018 inch of crack growth after 20 years. Even if a multiplier of 10 is applied to account for any unlikely high

temperature creep crack growth, this yields a maximum expected growth of 0.180 inch after 20 years or only 38% of the allowable crack propagation margin. Based on the foregoing critical crack size and crack growth calculations, the weld repaired rotor is qualified for a maximum of 400 starts or 20 years at 20 starts/year for a maximum flaw radial size of 1.5 mm (0.059 inch) near the bore.

Figure 31. - Rotational and Thermal Transient Hoop Stress Gradients used for Fracture Mechanics Calculations

Through-Wall Stresses at Join Weld

-50

-40

-30

-20

-10

0

10

20

0 1 2 3 4 5 6 7 8

Distance from Bore (inches)

Stre

ss (k

si)

ANSYS 3600 RPM (Original)ANSYS 3600 RPM (Repaired)SAFER 3600 RPM (Repaired);El.#548SAFER 3600 RPM (Repaired);El.#582SAFER Therm.Stress @300 sec.(Repaired);El.#548SAFER Therm.Stress @300 sec.(Repaired);El.#582

Figure 32. - Stress Intensity Factor (KI) versus Crack Size Results at 30% Overspeed for a Semi-Elliptic Surface Flaw

Figure 33. - Flaw Location & Shape Parameters Used In Allowable Flaw Size Plots

5

10

15

20

25

30

35

40

45

50

55

60

65

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5Crack Depth (inches)

Cra

ck T

ip S

tres

s In

t. Fa

c. K

(ksi√i

nch)

SAFER (a/l=0.05, 30% OS) SAFER (a/l=0.1, 30% OS)SAFER (a/l=0.2, 30% OS) SAFER (a/l=0.35, 30% OS)SAFER (a/l=0.5, 30% OS) Kic (=28.7 ksi√inch; Old Forging Base Material)Kic (=56.8 ksi√inch; Weld Min.)

Figure 34. - Allowable Flaw Sizes For 30% Overspeed Stresses & Minimum Expected Kic = 28.7 Ksi√Inch (Worst-Case Scenario: Old Forging Properties, 30% Overspeed with Cold Rotor)

Machining of the Forging / Weldment

After completion of all NDE, the rotor was restored to its original geometry by machining. The machining of the Impulse blade Grooves and the shaft end were performed in a large lathe, Figure 35. The coupling milling was performed on a horizontal-boring mill, Figure 36. The entire machined surfaces were magnetic particle inspected and dimensionally inspected after final machining.

Pulliam Unit 6 HP Rotor Repair: Allowable Flaw Sizes; 30% Overspeed (Min. Kic=28.7 ksi√inch)

0.529

0.0630

1

2

3

4

5

6

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.5Aspect Ratio, a/l (depth/length)

Cra

ck R

adia

l Dep

th (i

nch)

: a (S

urfa

ce) o

r 2a

(Bur

ied)

(Depth=0", Kic=28.7) (Depth=1", Kic=28.7) (Depth=2", Kic=28.7) (Depth=3", Kic=28.7) (Depth=4", Kic=28.7)Max. Allowable Flaw (1.5 mm=0.059 in.)

Min. Safety Fac.=8.5

Figure 35. Blade Groove, Dummy Piston and Seal Groove Machining

Figure 36. Spigot fit machining and balance plane hole drilling

Rotor Re-assembly

After completion of all the machining and drilling operations, measurements were taken to assure proper fit. After acceptance, the re-blading of the impulse and reaction blades began. The re-blading consisted of installation of the original refurbished impulse blades and new reaction row blades and shroud covers, Figure 37. NOTE: (Visual and magnetic particle examination of the reaction row 1-3 blades showed indications in the hook radii. Based on these results a decision was made to replace these three rows of blades, keeping the project on schedule). The final scope of work included a low speed balance of the rotor prior to shipment back to site.

Figure 37. Blading Operation

Unit Post Repair Operation

After weld repair and re-blading, the rotor was placed in service in April 2005. To date the rotor is in service and operating normally in all respects with no problems. In regards to unit vibration monitoring, since the station already had continuous vibration monitoring equipment installed on the unit, no special unit vibration collection plan was required. In regards to a post repair inspection plan, no special post repair inspections would be required other than those already required by the OEM. The rotor forging in

the area of the join weld should be blasted, visually inspected, and magnetic particle inspected at the future planned major outage for the equipment. In addition, a boresonic examination on the remaining original rotor forging should be scheduled as originally planned. The repair project for the rotor was completed in 59 days from receipt of rotor until shipment of the rotor back to the plant site. Overall, the salvage of this rotor was a success for all involved and was due to the coordinated efforts between WPSC, ReGENco, Structural Integrity and ALSTOM Power Inc.

Acknowledgements

The authors gratefully acknowledge the assistance of Mr. Steve Schrickel of WPSC, who provided invaluable input with regard to risk assessment to assure project success. The authors would like to thank Mr. Darryl Rosario of SIA, who gave up many hours of sleep to provide quick and accurate turn around of preliminary analysis results to keep the project moving forward. The authors are also indebted to Dr. Don McCann of ReGENCo, for keeping this unit on-line while the preparations for section replacement were being finalized. Finally, the authors would also like to thank Mr. Randy Wilson and the staff at the Alstom Materials Lab for quickly providing all the replication, chemical and mechanical property data in support of this project.

References

1. Jirinec, M. J., Kilroy, R., Prescott, E. and Rosario, D. - “Steam Turbine Rotor Reclamation Via Deep Groove Submerged Arc Join Welding” EPRI Sixth International Welding and Repair Technology for Power Plant Conference, Sandestin, FL June 2004.

2. Kilroy, R.E., Morin, M., Radcliff, M., Sculley, D., Cable, J., & Shriver, D.,

“Shaft Forging Section Replacement By Join Welding of a Large Generator Rotating Field – A Case Study”, Fifth International EPRI Conference on Welding and Repair Technology for Power Plants, Point Clear, Alabama, June 2002.

3. Alstom Power Project Reference Files – 1340072 (2005)