L-PI-09-009, Supplement to Northern States Power Company, a Minnesota Corporation (Nspm), Position on Two Apparent Violations and Preliminary White Findings, EA-08-272 and EA-08-273

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Supplement to Northern States Power Company, a Minnesota Corporation (Nspm), Position on Two Apparent Violations and Preliminary White Findings, EA-08-272 and EA-08-273
ML090150041
Person / Time
Site: Prairie Island  Xcel Energy icon.png
Issue date: 01/12/2009
From: Wadley M
Northern States Power Co, Xcel Energy
To:
Document Control Desk, Office of Nuclear Reactor Regulation
References
EA-08-272, EA-08-273, L-PI-09-009
Download: ML090150041 (45)
v • d • e


Text

Prairie Island Nuclear Generating Plant 1717 Wakonade Drive East ZWelch. MN 55089

.01/12/2009 L-PI-09-009 EA-08-273 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington, DC 20555 Prairie Island Nuclear Generating Plant, Units 1 and 2 Docket Nos. 50-282 and 50-306 License Nos. DPR-42 and DPR-60 Supplement to Northern States Power Company, a Minnesota Corporation (NSPM),

Position on Two Apparent Violations and Preliminary White Findings, EA-08-272 and EA-08-273.

Reference:

1. Letter from NSPM to the NRC, "Northern States Power Company, a Minnesota Corporation (NSPM), Position on Two Apparent Violations and Preliminary White Findings, EA-08-272 and EA-08-273", dated December 5, 2008.

This letter submits additional information regarding the 11 Turbine-Driven Auxiliary Feedwater Pump (TDAFWP) high outboard turbine bearing temperature event at Prairie Island Nuclear Generating Plant (PINGP) during the Unit 1 refueling outage number twenty-five (1 R25).

The original NSPM submittal, dated December 5, 2008, included a bearing analysis report by MPR Associates Inc. (MPR) dated December 1, 2008 that assessed whether the 11 TDAFWP outboard journal bearing would have been capable of operation for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> if the pump had been put into service. The conclusions of the MPR report stated that the bearing would have been able to operate for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> with bearing temperatures less than 300 degrees F, but there was insufficient data to conclusively state that bearing steady state temperature would not have exceeded 300 degrees F.

NSPM has since provided additional information to MPR. This information allowed MPR to perform further analyses of the 11 TDAFWP and determine its partial capability. The additional MPR analysis in combination with calculations of decay heat load and required AFW flow during the period in question (MAAP4 Analysis) was utilized in NSPM engineering evaluation 13630 to evaluate the significance of the high outboard bearing temperature. The following two paragraphs summarize the conclusions drawn from the engineering evaluation.

Document Control Desk Page 2 Engineering evaluation 13630 concludes that the condition of 11 TDAFWP did not significantly increase risk from 3/15/2008, 10:59 AM to 3/23/2008, 5:00 AM (186 hours0.00215 days <br />0.0517 hours <br />3.075397e-4 weeks <br />7.0773e-5 months <br />) because it can be said with a high degree of confidence that the pump could have provided full flow for a short period of time and adequate steam generator water level due to the very small decay heat load which would mainly be dissipated to containment.

Therefore, the plant would have been maintained in a safe condition. AXprobabilisti6 risk assessment (PRA) significance evaluation performed for this issue demonstrates that the dominant risk contributors are associated with fires that can result in a control room evacuation. The results of this engineering evaluation directly impact the results of the significance evaluation in that the increase in plant risk due to the issue would have been negligible during the initial 0% and low power time frames between 3/15/2008 and 3/23/2008. This time window is a large fraction of the overall time window associated with the potential unavailability of the pump. The PRA significance evaluation used an estimated exposure time of 10 days. The exposure time as outlined in this evaluation is 34.67 hours7.75463e-4 days <br />0.0186 hours <br />1.107804e-4 weeks <br />2.54935e-5 months <br />. Using the entire time window of 10 days, the risk contribution due to control room evacuation from a fire was estimated to be less than 3.96E-7 ACDP (change in core damage probability). Considering the reduced time window, the ACDP due to control room evacuation is 34.67 hours7.75463e-4 days <br />0.0186 hours <br />1.107804e-4 weeks <br />2.54935e-5 months <br /> / 240 hour0.00278 days <br />0.0667 hours <br />3.968254e-4 weeks <br />9.132e-5 months <br />

  • 3.96E-7 =

5.72E-8 ACDP.

For the 34.67 hours7.75463e-4 days <br />0.0186 hours <br />1.107804e-4 weeks <br />2.54935e-5 months <br /> of risk exposure, there is reasonable assurance to conclude that 11 TDAFWP would have been able to supply the AFW flow demand shown in the MAAP4 Analysis (Figure 2 of Attachment F in engineering evaluation 13630) between 10%

power and when 11 TDAFWP was proven operable (34% power). This is based on the determination that the bearing is capable of temperatures up to 300°F for a 24-hour period and the available temperature trends support the conclusion that the turbine bearing temperature would not have exceeded 300 0 F. Additionally the majority of the AFW demand is required to restore steam generator water level due to steam generator level shrink after the plant trip and not due to inventory loss. Although this evidence supports that the plant would have been maintained in a safe condition, the ACDP number listed above did not take credit for 11 TDAFWP during this time period of approximately 35 hours4.050926e-4 days <br />0.00972 hours <br />5.787037e-5 weeks <br />1.33175e-5 months <br />. Even with the increased risk, the ACDP for the reduced time frame shows the event was of relatively low safety significance.

Engineering evaluation 13630 is included as enclosure 1 to this letter.

Document Control Desk Page 3 Summary of Commitments This letter contains no new commitments and no revisions to existing commitments.

Michael D. Wadley Site Vice-President Prairie Island Nuclear Generating Plant Northern States Power Company-Minnesota Enclosure cc: Regional Administrator, Region Ill, USNRC Project Manager, Prairie Island Nuclear Generating Plant, USNRC Resident Inspector, Prairie Island Nuclear Generating Plant, USNRC Reactor Projects Branch 4 Chief, Region III, USNRC Senior Risk Analyst, Region Ill, USNRC

ENCLOSURE 1 ENGINEERING EVALUATION 13630, 11 AFW PUMP HIGH OUTBOARD TURBINE BEARING TEMP 1R25 41 pages follow

Engineering Evaluation 13630 11 AFW PUMP HIGH OUTBOARD TURBINE BEARING TEMP 1R25

==

Description:==

The 11 TD AFW Pump (145-201) experienced high outboard turbine bearing temperatures during Unit 1 startup from refueling outage 1R25 in March of 2008 and was not fully operable for a period of time. Unit 1 entered mode 3 on 3/15/08 at 10:59 AM at which time 11 TD AFWP was required by Tech Specs. The condition was discovered on 3/23/2008 at approximately 10:52 AM. The pump was proven operable on 3/24/08 at 3:40 PM after corrective actions were taken. The high bearing temperature was reduced by adding insulation to the outboard side of the turbine and by replacing insulation on the nearby governor.

The purpose of this evaluation is to determine the significance of this event by reviewing power history and calculating the potential decay heat removal requirements during the time period in question and then comparing those results to the capability of the 11 TD AFW Pump to provide flow to the steam generators. The bounding scenario for risk assessment (PRA) that will be evaluated is an Appendix R Control Room Evacuation. 11 TD AFWP is the only AFW pump assumed to be available for this event. The major inputs to this evaluation are a MAAP4 Analysis of TDAFW Demand for an Appendix R Control Room Evacuation Scenario (Attachement F) and bearing analysis provided by MPR Associates Inc. (Attachments C, D, and E).

This evaluation will focus on two main periods of time. The first interval starts when unit 1 entered mode 3 and the 11 TD AFW Pump was required by Tech Specs until the reactor power reached 10%. This represents the time when decay heat would have been minimal and it can be shown that no AFW flow would have been necessary to keep the plant in a safe condition. This evaluation will refer to this as period of time as "Interval 1". The second interval starts just before reactor power exceeded 10% and ends after the pump is tested and proven to be operable. For this period of time it will be shown that the degraded pump would have likely been able to provide adequate AFW flow but due to the limited data available it cannot be credited with a high degree of confidence. This evaluation will refer to this second period of time as "Interval 2". The date and time for these two time periods are shown in Table 1 below and are identified by the bolded boxes in the detailed timeline of related events provided in Attachment A. Selected Operations Log entries are provided in Attachment B as supporting information.

Table 1 - Time Calculator Start Time Stop Time Days Hours Interval 1 3/15/2008 10:59 AM 3/23/2008 5:00 AM 7.75 186.02 Interval 2 3/23/2008 5:00 AM 3/24/2008 3:40 PM 1.44 34.67 1Total 1220.68 Background Information for 11 TD AFW Pump Insulation The configuration of the insulation on the turbine (pump driver) after 1R25 was such that the turbine outboard bearing temperature would have stabilized over a period of time above the manufacture's shutdown limit of 220'F had the pump been required to provide

Engineering Evaluation 13630 11 AFW PUMP HIGH OUTBOARD TURBINE BEARING TEMP 1R25 the full design flowrate to the steam generators. However, the pump in this condition would have been able to either provide full flow for a short period of time or have provided flow at a reduced flowrate and not exceeded the bearing temperature shutdown limit. This will be discussed further in the justification section below and in the attached letters from MPR and Associates.

The manufacturer's shutdown limit for this bearing is 220'F. The high outboard bearing temperature was reduced by adding insulation to the outboard side of the turbine to protect the bearing housing from heat radiating from the side of the turbine case and by replacing degraded insulation on the nearby governor valve. Both of these changes provided better protection to the bearing housing from radiant heat. The pump was successfully run (Surveillance Procedure SP 1103, 11 Turbine-Driven Auxiliary Feedwater Pump Once Every Refueling Shutdown Flow Test) for approximately 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> under the turbine bearing shutdown limits at full pump flow on 3/24/08 after the insulation addition an replacement.

Justification:

It can be shown from the temperature trends included in Attachment G that the pump could have run for at least 12 minutes at full design flow. The pump starts at 9:33 AM on recirculation flow then at 10:00 AM reaches design flow to A steam generator and provides at least 200 GPM (Mainly 215 GPM) for approximately 7.5 minutes. Later B steam generator is fed for 4.5 minutes at approximately 210 gpm with and additional 16 GPM going to A steam generator for a total of 226 GPM. The bearing temperature reached 21 1F at the end of these full flow periods. Therefore it can stated that the pump would have been able to start and provide full flow for at least 12 minutes before exceeding the 220°F limit. This is very conservative because it does not include the amount of time that the pump initially ran on recirculation flow while it was warming up.

The 12 minutes was used as an input to the MAAP4 Analysis (Attachment F). For the sensitivity case the pump was run for 12 minutes and then failed. It is appropriate to use the above data for Interval 1 because it reflects the turbine insulation configuration between the time the reactor went critical and 10% power. The decay heat load prior to the reactor going critical was sufficiently low that AFW flow demand would have never exceeded 40 GPM as stated in the conclusion of Attachment F. Justification for the pump being able to provide at least 40 GPM with the initial insulation configuration is provided in Attachment E which reviews the data from the pump run on March 16, 2008 and associated bearing analysis in Attachment C.

Interval 1 The first period of time starts when Unit 1 entered mode 3 on 3/15/08 at 10:59 AM. The reactor went critical on 3/21/08 at 5:08 AM and reached 10% power 3/23/08 after 5:00 AM. The sensitivity case (10% initial power) provided in Attachment F, MAAP4 Analysis of TDAFW Demand for an Appendix R Control Room Evacuation Scenario, shows that the AFW flow demand after refill (12 minute) is virtually zero due to heat loss

Engineering Evaluation 13630 11 AFW PUMP HIGH OUTBOARD TURBINE BEARING TEMP 1R25 to containment heat sinks. This provides justification that the condition of 11 TD AFW Pump did not significantly increase risk during this time period.

Interval 2 The second interval starts just before reactor power exceeded 10% on 3/23/08 shortly after 5:00 AM and ends after the pump is tested and proven to be operable on 3/24/08 at 3:40PM. The base case (34% initial power) provided in Attachment F, MAAP4 Analysis of TDAFW Demand for an Appendix R Control Room Evacuation Scenario, shows that level would have been restored to the steam generators after 35 minutes with the pump providing full flow and that the AFW flow demand would have decreased to 40 GPM after 43 minutes. Note the majority of the flow demand to restore steam generator level is the result of SG level shrink at trip and not inventory loss.

While there is evidence that the pump would have very likely been able to provide the flow for this scenario it cannot be said for certain that the pump would have been successful. MPR Associates Inc. Letter Dated December 1, 2008 (Attachment C) states that the bearing is capable of 300'F for 24-hours but there is not sufficient data to say for certain that bearing temperature would not have exceeded 300'F. It is reasonable to expect that the pump could have provided full flow for 35 minutes and then provided flow at a reduced rate thereafter and remain under 300'F for the turbine bearing temperature. If a conservative linear trend of 4F /min increase was assumed starting from 220'F it would provide an additional 20 minutes of run time to reach 300'F in addition to the 12 minutes used above for a total of 32 minutes. For the event being evaluated the operators would not have known the bearing temperature (outside control room) and would have likely set flow to meet their procedural guidance and training to restore and maintain level. Therefore they would not have shut the pump down had the 220'F limit been reached.

==

Conclusion:==

Interval 1 The condition of 11 TD AFW Pump did not significantly increase risk from 3/15/2008 10:59 AM to 3/23/2008 5:00 AM (186 hours0.00215 days <br />0.0517 hours <br />3.075397e-4 weeks <br />7.0773e-5 months <br />) because it can be said with a high degree of confidence that the pump could have provided full flow for a short period of time and adequate steam generator level due to the very small decay heat load which would mainly be dissipated to containment. Therefore the plant would have been maintained in a safe condition. A PRA significance evaluation performed for this issue demonstrates that the dominant risk contributors are associated with fires that can result in control room evacuation. The results of this Engineering Evaluation directly impact the results of the significance evaluation in that increase in plant risk due to the issue would have been negligible during the initial no and low power time frames between 3/15/2008 and 3/23/2008. This time window is a large fraction of the overall time window associated with the potential unavailability of the pump. The PRA significance evaluation used an estimated exposure time of 10 days. The exposure time as outlined in this evaluation is 34.67 hours7.75463e-4 days <br />0.0186 hours <br />1.107804e-4 weeks <br />2.54935e-5 months <br />. Using the entire time window of 10 days, the risk contribution due to control room evacuation from a fire was estimated to be less than 3.96E-7 ACDP (change

Engineering Evaluation 13630 11 AFW PUMP HIGH OUTBOARD TURBINE BEARING TEMP 1R25 in core damage probability). Considering the reduced time window, the ACDP due to control room evacuation is 34.67 hours7.75463e-4 days <br />0.0186 hours <br />1.107804e-4 weeks <br />2.54935e-5 months <br /> / 240 hour0.00278 days <br />0.0667 hours <br />3.968254e-4 weeks <br />9.132e-5 months <br />

  • 3.96E-7 = 5.72E-8 ACDP.

Interval 2 There is reasonable assurance to conclude that 11 TD AFW Pump would have been able to supply the AFW flow demand shown in the MAAP4 Analysis (Figure 2 of Attachment F) between 10% power and when 11 TD AFW Pump was proven operable (34% power).

This is based on the determination that the bearing is capable of 300'F for a 24-hour period and the available temperature trends discussed above that support the turbine bearing temperature would not have exceeded 300'F. Additionally the majority of the AFW flow demand is required to restore steam generator level due to SG level shrink after the plant trip and not inventory loss. Although this evidence supports that the plant would have been maintained in a safe condition the ACDP number listed above did not take credit for II TD AFW Pump during this time period of approximately 35 hours4.050926e-4 days <br />0.00972 hours <br />5.787037e-5 weeks <br />1.33175e-5 months <br /> (Interval 2). Even with the increased risk during Interval 2 the ACDP above shows the event was of relatively low safety significance.

Attachments:

Attachment A - March 2008 Timeline of Events for 11 TD AFW Pump Attachment B - Selected Operations Log Entries for 11 TD AFW Pump Attachment C - MPR Associates Inc. Letter Dated December 1, 2008 Attachment D - MPR Associates Inc. Email Dated December 16, 2008 Attachment E - MPR Associates Inc. Email Dated December 17, 2008 Attachment F - MAAP4 Analysis of TDAFW Demand for an Appendix R Control Room Evacuation Scenario Attachment G - Data Trends for SP 1103 Performance March 23, 2008

Attachment A - March 2008 Timeline of Events for 11 TD AFW Pump Engineering Evaluation EC 13630 Reactor Power <10%

(>10% at 6:00 AM)

Discovery, SP 1103, Pump SP 1102, SP 1376, & Shutdown Due to High Temp SP 1330 Completed WR 00033776 completed to WO 00357419 completed to adjust turbine insulation adjust turbine insulation and replace degraded insulation, Pump run to check for steam leaks also completed

Attachment B - Selected Operations Log Entries for 11 TD AFW Pump Engineering Evaluation EC 13630 03/16/2008 Started SP 1376, AFW FLOW PATH VERIFICATION TEST AFTER EACH COLD SH per WO-00283304.

0715 (BJORKMAN, BRIAN, 1LPE & RO )

03/16/2008 Started SP 1102, 11 TURBINE-DRIVEN AFW PUMP MONTHLY TEST per WO-00330247. (BJORKMAN, BRIAN, 1105 1LPE & RO )

03/16/2008 Started SP 1330, 11 TURBINE-DRIVEN AFW TURBINE/PUMP BEARING TEMPERA per WO-00283299.

1254 (BJORKMAN, BRIAN, 1LPE & RO )

03/16/2008 Satisfactorily Completed SP 1330, 11 TURBINE-DRIVEN AFW TURBINE/PUMP BEARING TEMPERA per WO-1442 00283299. Comments: CAP 1131305 and WR 33776 issued for 11 AFWP Turb OB bearing reaching 216 degrees F, above the alert range of 203 degrees F. (JOHNSON, BRIAN J, MISC )

03/16/2008 Satisfactorily Completed SP 1102, 11 TURBINE-DRIVEN AFW PUMP MONTHLY TEST per WO-00330247.

1443 Comments: CAP 1131305 and WR 33776 issued for 11 AFWP Turb OB bearing reaching 216 degrees F, above the alert range of 203 degrees F. (JOHNSON, BRIAN J, MISC )

03/16/2008 Satisfactorily Completed SP 1376, AFW FLOW PATH VERIFICATION TEST AFTER EACH COLD SH per WO-1443 00283304. Comments: CAP 1131305 and WR 33776 issued for 11 AFWP Turb OB bearing reaching 216 degrees F, above the alert range of 203 degrees F. (JOHNSON, BRIAN J, MISC )

03/16/2008 Satisfactorily Completed AFW SP 1102, 1330, and 1376 CAP 1131305 and WR 33776 issued for 11 AFWP 1546 Turb OB bearing reaching 216 degrees F, above the alert range of 203 degrees F. Alert range does impact operability. (KAPITZ, JON K, Outage Manager )

03/23/2008 Started SP 1103, 11 TURBINE-DRIVEN AUX FEEDWATER PUMP ONCE EVERY RE per WO-00153953. (MURPHY, 0857 MICHAEL T, 1LPE & RO) 03/23/2008 145-201, 11 TD AFW PMP taken out of service per WO-00153953, WR-00034049. Component is Inoperable.

1052 OOS Note: Failed SP 1103 due to hi turb outbd bearing temp. (MURPHY, MICHAEL T, 1LPE & RO) 03/23/2008 Received ANNUNC 47010- 0605 11 TDAFW Oil Hi Temp. Stopped 11 TDAFWP per SP 1103 for Outbd Turbine 1052 Bearing Temp 220F. (MURPHY, MICHAEL T, 1LPE & RO )

Attachment B - Selected Operations Log Entries for 11 TD AFW Pump Engineering Evaluation EC 13630 03/23/2008 Elevated temperatures on the 11 TDAFW Pump. Bearing temperature reached 221 degrees. Limit is 220 1059 degrees. Pump was secured. (CARLISLE, DAVID, MISC )

03/23/2008 Unsatisfactorily Completed SP 1103, 11 TURBINE-DRIVEN AUX FEEDWATER PUMP ONCE EVERY RE per WO-1141 00153953. Comments: Pump stopped due to Turb Outbd Bearing Temp exceeding 220F. Issued CAP 1132098 WR 34049. (MURPHY, MICHAEL T, 1LPE & RO )

03/24/2008 Started 145-201, 11 TD AFW PMP per WO-00357419. Per 1C28.1 (HANSON, JEFF A, 1PE & RO )

0225 03/24/2008 Stopped 145-201, 11 TD AFW PMP per WO-00357419 with direction from system engineer. (HANSON, JEFF A, 0232 1PE & RO) 03/24/2008 Started SP 1103, 11 TURBINE-DRIVEN AUX FEEDWATER PUMP ONCE EVERY RE per WO-00316496. (HAREN, 1221 MARK, 1LPE & RO )

03/24/2008 Satisfactorily Completed SP 1103, 11 TURBINE-DRIVEN AUX FEEDWATER PUMP ONCE EVERY RE per WO-1540 00316496. (HAREN, MARK, 1LPE & RO )

03/24/2008 11 TDAFW Pump has been declared operable. Exit TS 3.7.5 Condition B. Reference CAP 1132098. (JOHNSON, 2238 BRIAN J, SM )

03/24/2008 Returned 145-201, 11 TD AFW PMP to service per WO-00153953, WR-00034049. (HORTON, LARRY, 1LPE &

2238 RO)

Attachment C SMPR ASSOCIATES 1"C_

E N G IN E P S December 1, 2008 Mr. Gary Wheelock Xcel Energy Prairie Island Nuclear Plant 1717 Wakonade Drive East Welch, MN 55089-9642

Subject:

Prairie Island Unit One -- Turbine-Driven Auxiliary Feedwater Pump Assessment of High-Temperature Operation of the Outboard Turbine Bearing

Dear Mr. Wheelock:

MPR has assessed whether the outboard journal bearing of the drive turbine for the AFW pump at Prairie Island Unit One would have operated for 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> if the pump had been put into service during the period in late March 2008 when an improper thermal insulation configuration was installed. This assessment was done in support of Xcel's PRA of the consequences of the pump being inoperable due to the improper insulation configuration.

Our assessment concludes that the bearing would operate at temperatures of up to at least 300'F for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. The available test data with the improper insulation installed does not allow a definitive statement to be made about whether the maximum, steady-state bearing temperature would have been less than 300'F with the AFW pump providing steady, maximum feedwater flow rate. However, the test data shows that lesser flow rates could have been provided intermittently without the bearing exceeding 300TF had operation of the pump been necessary.

The core decay heat burden may have been lower than normal during the period of interest because the plant was coming out of an outage. If so, full flow rate capability from the AFW pump may not have been needed had the pump been called into service. If this possibility has not already been explored, we recommend you consider it as a possible path to showing an acceptable outcome using the pump in its degraded condition.

Please see the enclosure for the details and the bases for our conclusions. If you have any questions or if we can be of further assistance, please call.

Sincerely, T. E. Greene Enclosure 320 KING STREET ALEXANDRIA, VA 22314-3230 703-519-0200 FAX: 703-519-0224 http://www.mpr.com

Attachment C

  • MPR ASSOC(ATES NC.

ENG iN'E E R S Enclosure to MPR Letter Dated December 1, 2008 Prairie Island Unit One - Turbine-Driven AFW Pump Turbine Outboard Bearing Operation at High Temperatures

1. Introduction During pump flow testing in late March 2008, the drive turbine for the turbine-driven AFW pump at Prairie Island Unit One experienced high temperatures in the outboard radial bearing.

The pump was secured when operators concluded that the bearing temperature would very likely exceed a 220'F limit The drive turbine was taken out of service and inspected. Plant staff determined that the high temperature condition was due to a change in turbine insulation. The insulation arrangement was revised and the pump was successfully returned to service.

Xcel has determined that the pump was not operable during the time that the improper insulation was installed. They are assessing, by probabilistic methods, the consequences of the pump not being operable. As part of that assessment, information about the ability of the outboard journal bearing to operate at higher-than-normal temperature is needed. MPR has been asked by Xcel to assess the bearing to help determine whether the bearing would have operated for 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> if the pump had been put into service. This letter report summarizes our assessment.

2. Discussion 2.1. Descriptionof Equipment The drive turbine for the AFW pump is a Terry turbine. The turbine is equipped with inboard and outboard radial bearings. The bearings are sleeve-type babbitted journal bearings. Each radial bearing is equipped with double oil rings for lubrication. The oil rings obtain oil from individual oil reservoirs. The oil reservoirs are water cooled by finned-tube cooling coils mounted in each reservoir.

Information provided by Xcel and used in this evaluation includes the following:

Approximate turbine rotor weight (shaft and wheel): 89 lb Pump coupling weight: 54 lb Approximate turbine bearing span, center-to-center: 17 inches Outboard journal diameter: 1.996 inches

Attachment C Bearing overall length: 2.50 inches Bearing central groove width: 0.375 inch Outboard bearing diametric clearance: 0.005 inch Babbitt material: SAE 15 Lubricating oil specification: Mobil DTE Medium Heavy Grade Oil-ring outside diameter: 4.4785 inches Oil-ring inside diameter: 4.3065 inches Oil-ring axial width: 0.185 inches Oil ring submergence: 0.375 inches Oil ring masses (sample of two): 39.184/40.977 grams Turbine operating speed: 3570 rpm 2.2. Assessment of the JournalBearing Mean PressureLoading The mean pressure loading is an important parameter when estimating the performance of a journal bearing. The primary load on the drive turbine's journal bearings is the rotor weight. The rotor is nearly symmetric, as shown in Figure 2-1, so the weight of the rotor should be shared about equally between the inboard and outboard bearings. The weight supported by the bearings includes the weight of the shaft and the wheel, half of the weight of the pump coupling, and the weight of the disk mounted adjacent to the outboard bearing. The weight of the disk is not known, but it is comparable to the pump coupling at the other end of the shaft, so its weight has been estimated to be about the same as the pump coupling weight supported by the turbine shaft.

The total rotor weight is estimated to be about 143 lbf, and the load on each bearing is 71.5 lbf.

I! /I/ft Figure 2-1. Prairie Island AFW Turbine Cros~s Section (from vendor manual, Ref. 1)

The mean pressure loading on the bearing is the supported weight divided by the projected area (i.e., length x diameter). For this bearing, the resulting mean bearing loading is about 14 psi. The Attachment C actual bearing has a central relief groove, as shown in Figure 2-2. Subtracting out the projected area of the central relief groove results in a slightly higher mean pressure loading: about 17 psi.

This is a typical pressure loading value for a Terry turbine, but it is a very light pressure loading for ring-oiled bearings, which generally have loadings about ten times the Terry turbine's bearing loading. This light loading is beneficial for high temperature operation, and is the main reason that Terry turbines are able to operate at high bearing temperatures without wiping the bearings.

Central Relief Groove in the Bearing f25tt 2 STAINLt 5 in Figure 2-2. Bottom half of an outboard bearing from the drive turbine of a Prairie Island AFW pump. The picture shows the central relief groove.

2.3. Assessment of the Bearing OperatingPoint The oil film between the rotating journal and the bearing is an important consideration and can be limiting in some applications. A bearing operating map is a useful starting point for evaluation of the oil filn. The operating map defines an envelope of permissible operation in light of several potentially limiting considerations. Figure 2-3 shows an operating map, taken from Reference 2, for starved journal bearings. The vertical axis is bearing mean pressure. The horizontal axis is the relative speed between the journal surface and the bearing surface. For the bearing of interest, the coordinates, in convenient units, are 17 psi bearing pressure and 3-

Attachment C 9.5 meters/sec rubbing speed. The operating point for the Prairie Island turbine bearing is marked with a red star on the operating map.

Limit lines for minimum oil film thickness RUbBING 3PEk1, tt>"n 5; 10 100IC r -" .E 0 Re  :

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,U TIG .R " BR-.C ZD (jl

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/. !RE'GIOb o.'-JL ~U-O'CR FOIR / URICAI*3 B-ARIG i; *i.

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40

  • .hY 0.92 (0.5 O.1 0.2 0.5 2, Limit lines for journal bearings. (Ref. 2) oxidation oil I

Figure 2-3. Bearing operating map for starved The limit lines on the operating map define three potential limits for the bearing. The rising limit lines on the left side are based on minimum oil-film thickness considerations, with the bearing being vulnerable to failure from metal-to-metal contact if a limit line is exceeded. The top and declining limit lines in the center are based on maximum babbitt temperature, with the bearing being vulnerable to wiping due to inadequate babbitt strength at temperature if a limit line is exceeded. The vertical limit lines to the right are based on oil oxidation due to high temperature.

Figure 2-3 shows that the bearing operating point for this application is well away from the minimum oil-film thickness and babbitt strength operating limits. This suggests that neither of these considerations is likely to be an issue. This is not surprising considering the very light loading on the bearing. The operating point is right at the oxidation limit line for a drip-fed bearing. To operate at speeds higher than the drip-fed limit line, a ring-oil, splash, or pressure-fed lubrication system is needed. Since the bearing is equipped with an oil-ring system, the bearing lube-oil system is appropriate, and just barely needed. This suggests that the system should have margin for abnormal operation.

In the case at hand, the turbine bearing housing was over-insulated and operated at higher-than-expected temperatures. Therefore, additional assessments of the potential for oxidation of the Attachment C lubricating oil, significant reduction in the oil-film thickness and the strength of the babbitt material are warranted.

2.4. Assessment of Oil Oxidation Potential Oxidation of lubricating oil is mainly a long-term consideration. Oil run at high temperature will gradually degrade and eventually reach a point where it must be replaced. The considerations are increased viscosity of the oil and production of acids that might cause corrosion of the bearing material. High-temperature stability is routinely evaluated for turbine oils. The specification sheet for the Mobil DTE medium heavy oil used in the bearing indicates that it has an ASTM D943 stability life of 3500 hours0.0405 days <br />0.972 hours <br />0.00579 weeks <br />0.00133 months <br />. This test heats the oil to 95°C (203'F) and exposes it to pure oxygen. The rating is the time required for the oil to suffer a certain amount of oxidation in the test.

The exposure on test to pure oxygen is more severe than exposure to ordinary air, which is only about 21% oxygen. The test temperature, however, is lower than the range of interest in this evaluation. The oxidation rate of the oil is very sensitive to temperature. Therefore, some assessment of the affect of higher temperature is necessary.

The temperature sensitivity of oil oxidation typically follows the Arhenius law. For typical Arhenius degradations, a 10C increase in temperature halves the life. An oil temperature of 275'F is 40'C above the D943 test temperature, implying a factor of 16 (24) reduction in life (3500/16 = 219 hours0.00253 days <br />0.0608 hours <br />3.621032e-4 weeks <br />8.33295e-5 months <br />). At a temperature of 300'F, 53°C above the test temperature, the reduction factor would be about 40, reducing the oxidation time to 3500/40 = 88 hours0.00102 days <br />0.0244 hours <br />1.455026e-4 weeks <br />3.3484e-5 months <br />. These reduced test times are much longer than the 24-hour period of interest.

An additional source of margin is that the oil is not continuously at the bearing operating temperature. The temperature of the oil drops substantially when the oil re-enters the oil reservoir and is cooled. Therefore, the actual operating time required to reach the test limits would be much longer than the reduced test times estimated above. Therefore, oxidation of the oil and the consequential concerns are not considered to be a concern for bearing operation at temperatures up to 300'F for an operating period of 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> provided that the oil does not have significant, pre-existing oxidation at the start of the postulated event.

2.5. Assessment of Oil-Film Thickness An approximate evaluation of the minimum oil film thickness has been made as an additional check on the conclusions from the operating map discussed in Section 2.3. The approach used involves two steps. The first step is to estimate the oil-ring oil delivery rate using the oil-ring oil-delivery correlations in Reference 3. This provides an estimate of the oil delivery rate. The predicted oil- delivery rate to the bearing is in the range of .016 to .031 gpm. The lower-bound oil delivery rate of .016 gpm is used in this estimate.

The second step is to use the oil-flow estimate and the applied bearing load to estimate the oil-film thickness as a function of temperature using the methods of Reference 4. The method of Reference 4 considers only a bearing geometry in which the length of the bearing is equal to its Attachment C diameter (i.e., L/D=I). The bearing of interest has an overall L/D of 1.25, and an L/D of about t when the central relief groove is subtracted out. While the L/D with the central groove subtracted out is essentially equal to the L/D=I analyzed in the Reference 4, the application is not exact because the bearing of Reference 4 does not have a relief groove. The minimum oil-film thickness will be reduced by the presence of the central relief groove. The reduction is estimated to be about 20% in the high temperature range of interest.

Figure 2-4 shows predicted minimum oil-film thickness as a function of bearing operating temperature based on the methods of Reference 4 with 20% deducted to account for the presence of the central relief groove. The predicted behavior is a steady reduction in minimum film thickness with increasing temperature, without any sudden decreases. The moderating rate of reduction of film thickness with increasing temperature occurs because the oil has a similar characteristic reduction in viscosity. A minimum film thickness limit of 0.0005 inch is shown on the plot. This is a reasonable estimate of the limit for a ground shaft journal and a well-machined bearing babbitt surface. These results indicate that the minimum oil-film thickness should be adequate for operation to temperatures up to at least 300'F.

1-0-Estimated Min Film Thickness - - Min Film Limit]

0.00121 - I 0.001 0.0008 C

0.0006 0

0.0004 0.0002 0

150 175 200 225 250 275 300 325 Bearing Operating Temperature (F)

Figure 2-4. Estimated oil-film thickness as a function of bearing operating temperature.

Attachment C 2.6. Assessment of Oil-Film Temperature Rise.

The bearing operating temperature is measured at the top of the bearing, 45 degrees from top dead center. This is considered a good estimate of the maximum film temperature because the bearing is very lightly loaded and because most of the oil in the oil film is recirculated through the bearing. (The side leakage rate is only about 10% of the total oil flow in the film.)

Reference 6 provides a method of checking this conclusion quantitatively.

As a result of viscous dissipation, the temperature of the oil in the oil film increases as the oil progresses from the inlet through the area of minimum oil-film thickness. The maximum babbitt temperature is expected to occur near the minimum oil-film thickness, and this temperature will be the limiting value for assessment of the babbitt strength. Reference 6 provides an analytical model that estimates the maximum temperature in the oil film knowing its operating eccentricity ratio and the oil temperature at the inlet to the oil film.

The analytical model integrates the Reynolds lubrication equation and the energy equation along the inlet arc of the bearing at the center of the bearing. Reference 6 includes experimental validation of the model predictions. Figure 2-5, from Reference 6, compares the model predictions to the experimental data, plotting the maximum temperature versus the eccentricity ratio'. The 3600-rpm case in Figure 2-5 validates the turbulent-flow model, and the 1800-rpm case validates the laminar-flow model. The bearing of interest in this evaluation operates in laminar flow, so the laminar flow model has been used in this evaluation.

Figure 2-6 shows the results of application of the analytical model of Reference 6 to the AFW turbine bearing for assumed inlet oil temperatures of 263°F and 300'F. The plot shows the predicted maximum bearing temperature as a function of bearing eccentricity ratio. The bearing eccentricity ratios for 263'F and 300'F oil inlet temperatures can be determined from the minimum oil-film predictions in Figure 2-4. The minimum oil-film thickness at 263'F is

.00072 inches, so the eccentricity ratio has a value of 0.71 [1-(.00072 in)/(.005in/2)]. At 300'F, the eccentricity is 0.75 [1-(.00062 in)/(.005 in/2)]. These operating points are marked on the model prediction curves in Figure 2-6. The results show that the predicted increases in oil-film temperature are not significant - only 5 or 61F - as expected. Therefore, the measured bearing temperature should be close to the maximum temperature of the bearing.

The eccentricity ratio is defined as the quantity [1-(minimum oil-film thickness)/(nominal radial clearance)].

7

Attachment C Figure 2-5. Comparison of maximum temperature predictions from the analytical model of Reference 6 to experimental data.

Figure 2-6. Predicted maximum bearing temperature as a function of eccentricity for the AFW turbine outboard bearing at operating temperatures of 263'F and 300'F.

8_

Attachment C 2.7. Assessment of Babbitt MaterialStrength at High Temperature The potential for babbitt metal displacement from high temperature operation is a potential concern. This is what defines the top and top-right limit lines in the operating map of Figure 2-3.

Reference 6 reports a set of tests and analyses that investigate high-temperature wiping of journal bearings under steady loading. Table 2-1 summarizes the test data.

Table 2-1 Summary of Experimental Results from Reference 6.

Mean Pressure Babbitt Type Bearing test Temperature for Start Loading of Babbitt Creep 1050 psi ASTM B23 Grade 2 Thrust bearing, 4820 rpm 268OF 704 psi ASTM B23 Grade 3 Journal bearing, 5000 rpm 305OF 200 psi ASTM B23 Grade 3 Journal bearing with 380OF oscillating motion Reference 6 concludes that the bearing wipes when the maximum bearing pressure exceeds the babbitt material yield strength at temperature, causing the babbitt metal to creep. Figure 2-7, from Reference 6, shows the results of the tests superimposed on the tensile yield strength of the babbitt material. The plot indicates that failure begins when the maximum bearing pressure approaches the yield strength of the metal at the maximum bearing temperature (Note in the figure that the alloy grade numbers in the plot labels are reversed although the chemical compositions are correct.)

8000 02% OFFSET YIELO STREI6THt:

AST U ALLOY 2 (84 6*.eSb.8OW

AS rTV ALLCY CC)~

6000 cc 4000 co 20001 TEMPERAMPUAEF Figure 2-7. Bearing babbitt 0.2% offset tensile yield strength with creep test results superimposed (Ref. 6).

Therefore, for purposes of assessing the potential for wiping of the bearing due to the steady loading, loading less than the babbitt strength at temperature will be considered acceptable.

Attachment C The babbitt metal for the Prairie Island AFW pump bearings is reported to be SAE 15 material.

This babbitt material is a lead-based babbitt having a nominal composition of 83% Lead (Pb),

16% Antimony (Sb), and 1% Tin (Sn). It was widely used in internal combustion engines for crankshaft bearings operating at high temperature with significant dynamic loading until environmental concerns about the use of lead curtailed its use. SAE 15 is also known by its ASTM designation, ASTM B23, Grade 15.

The SAE 15 material is less commonly used in power applications than the more familiar Tin-based babbitt alloys (i.e., ASTM B23, Grades 2 and 3). Table 3 of Reference 7 compares a number of bearing materials and characterizes SAE 15 as having a load capacity of 800-1200 psi and a maximum operating temperature of 300'F. It lists similar load capacity and temperature data for the more familiar tin-based babbitts: 800-1500 psi and 300'F. As previously noted, the mean pressure in the bearing is about 17 psi. The peak pressure is typically 3-4 times the mean pressure (Ref. 6). Therefore, the maximum pressure load on the bearing should be on the order of 70 psi or less. This result and material data from Reference 7 suggest that the babbitt should have acceptable performance at temperatures up to at least 300 0 F.

Table 4.5 of Reference 8 provides compressive yield strength data for a number of common babbitt metals, including the ASTM B23 Grades 2, 3, and 15. Figure 2-8 plots the data for these three babbitt metals. The data show that the Grade 15 material is slightly weaker than the tin babbitts, but not significantly weaker at temperatures around 300'F. The compressive yield strength of the Grade 15 material at 300NF is in excess of 1000 psi and more than ten times the estimated peak pressure in the bearing of interest, 70 psi. Therefore, the compressive strength data indicates that the bearing should not be at risk of wiping at temperatures up to at least 3000 F.

7.0ASTM B23 Grde 3 ASTM B23 G, 2 ASTM 823 GM.ra

... 175 ..

7000- -_____ _

6000 5000 4000 3000 2000 10000 0-50 100 150 200 250 300 350 4V0 Temperature (F)

Figure 2-8. Compressive yield strength as a function of temperature for ASTM 1323 Grades 2, 3, and 15.

- 10-

Attachment C 2.8. Assessment of Babbitt FatigueStrength at High Temperature Bearings are subjected to dynamic loads as well as steady loads. Repeated dynamic loading can result in fatigue damage to the babbitt material at stresses well below the material yield strength.

Therefore, fatigue should be considered.

The AFW steam turbine is an application with relatively steady loading of the bearings compared with applications such as crankshaft bearings in internal combustion engines. The dynamic loads on the turbine bearings are expected to be less than the steady load because otherwise, the journal would be subject to relatively large radial vibrations on the order of the bearing clearances. In Section 2.7, the peak steady pressure on the bearing was estimated to be not more than 70 psi. Therefore, this value will be taken as a bounding estimate of the dynamic loading of the bearing. The most significant vibration frequency in most rotating machines is the operating speed. At an operating speed of 3570 rpm, approximately 5 million cycles of vibration will occur in the 24-hour period of interest.

Figure 2-9 shows fatigue loading data from Reference 5 for tin babbitt material as a function of temperature. The mean load is on the horizontal axis and the alternating load is on the vertical axis. The circles represent the 10-million-cycle life of the babbitt for the specified temperature range. No fatigue failures occurred in 10-million cycles for steady and dynamic loadings within the defined circles having temperature labels.

The babbitt of interest in this evaluation is a lead-based babbitt. The tin-based babbitt fatigue data is considered representative of the lead-based babbitt based on the test results discussed in Reference 9. Figure 2-10, taken from Reference 9, shows the relative fatigue strengths of several bearing materials. The first bearing material in Figure 2-10 has the composition of ASTM B23 Grade 2 tin-based babbitt metal. The second material has the composition of ASTM B23 Grade 15 lead-based babbitt material. The comparison shows that the lead-based material has slightly better fatigue strength. Therefore, use of the tin-babbitt fatigue data of Figure 2-9 for the lead-based babbitt material is reasonable.

For the estimated maximum bearing pressure of 70 psi and the bounding 70 psi dynamic loading, the operating point is well within the limit circle shown on Figure 2-9 for the babbitt temperature range of 250-295'F, and it is very near the origin of the plot. This data indicates that the bearing should not be vulnerable to a fatigue failure if operated at temperatures up to 300'F for a period of about 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. Therefore, the fatigue performance of the bearing is judged to be acceptable for the event of interest.

Attachment C 0

Figu A-6 Effect of load and tenperdtaie on fatigue life of tin (90%0) babbitt. Inside a given loop, no failures occurredup to li loadcycle.s (CGyde)

Figure 2-9. Fatigue strength of tin babbitt metal versus temperature. (Ref. 5)

COMIMONA NAE COUP.

TJI BASE WhITE 3S-"b -MTAL LEAD BSP... 157L-WHIT ME'ZM. ISn IXA.

207. "IN ALLUMINIUM Al'- ACX18n 67' TIN ALUWIBNIUJM f 0 II4I _ _ns OVERLAY PLATED Al-0_2n L 67 TN ALUMINUM Zun ub l a.

LEAD) BROBTW ulA)! ____)E' LEAD BRONZE Cu- IoxPb ___________

TIN -LA RNE sR uLEALAYPLATED z.m-aaX~t- o~ 2 TIN - LEAD BRONZE exsn ig. 2 .4COmPpCITISO1 of Ictilgue F___________

A; __________________ strength e'4 vtirious Joit, frietion ALUMINIUM SILION IeC MeSring rteriots 131 OVERLAY PLATED Al-lu]t -___

AILILINIUM01SYJYDA I1Wn 1P ID 45 60 71 %1D j

'.q 1Q62 Mlor r4t119ýt "'fa

-AJo ýf -1.ray 1s du. to UIt.- biclan too Figure 2-10. Comparison of Bearing Material Fatigue Strength (Ref. 9)

- 12-

Attachment C 2.9. Expected Bearing Operating Temperaturesin the Event of Interest The results of the bearing evaluations in Sections 2.2 - 2.8 indicate that the outboard journal bearing in the AFW pump drive turbine should be able to operate successfully for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> at temperatures up to at least 300'F. An assessment of the likely operating temperatures in the condition of interest (i.e., with the improper insulation configuration) is needed to complete the evaluation.

Figure 2-11 shows the available test data from operation of the AFW pump with the improper insulation configuration. The data is from surveillance testing done by plant operators at Prairie Island, and it is the testing that identified the high bearing temperature concern. The data shows the bearing heating up slowly after the pump is started and run on minimum flow. The bearing temperature curve looks as if the bearing temperature would come to an equilibrium temperature well under 250°F if the pump was left operating on minimum flow.

AC T NA V I f flC r 18,3.3~3 ...

116 667 11 TD AFWP T:JR3 CBRG T D2FUT)

Figure 2-11. Turbine Outboard Bearing Temperature and AFW Flow Rates to the Steam Generators During Testing on March 23, 2008, with Improper Insulation Installed The outboard bearing shows three short-term increases in temperature during the test. Each excursion is associated with increased pump flow rate while the pump supplies feedwater to one or both steam generators. The last peak in the temperature plot is when the bearing temperature reached or approached the 220'F temperature setpoint value, and operators secured the pump.

The increase in temperature during periods of increased pump flow is believed to be the result of increased heat generation in the anti-friction thrust bearing co-located with the outboard bearing due to increased thrust as the turbine is loaded. The journal bearing oil supply and the radiant cooling of the bearing housing must carry away the heat generated by both bearings. The Attachment C bearings and housing will continue to warm until their temperatures are high enough to equilibrate the heat being generated inside the bearing housing with the heat removed by the oil and the heat transferred to the environment.

The periods of significant flow to the steam generators during the test are short: in the range of five to ten minutes. The bearing and housing, which appear to have a thermal time constant of roughly 45 minutes, never have time to approach thermal equilibrium in the high-flow condition.

As a result, a reasonably accurate estimate of the equilibrium temperature in the high-flow condition cannot be made from the data. Therefore, the data cannot be used to definitively say that the bearing would have equilibrated in the high flow rate condition to a temperature less than about 300 0 F.

The data of Figure 2-11 does seem to indicate that the pump can provide high flow rates on the order of 200 gpm for periods of perhaps 10-15 minutes without the bearing temperature getting anywhere near 300NF, and could provide lesser flow rates, in the range of 75 to 100 gpm for longer periods of time without risk of bearing failure. It seems likely that if flow from this pump were urgently needed, operators would have been able to intermittently feed the steam generators with intervening periods of low or minimum flow rate operation to allow the bearing to cool down. Given that the plant was coming out of an outage when the insulation problem was discovered, it may be possible to show that the full flow capability of the pump would not have been needed due to the lower-than-normal decay heat burden present at that time. This possibility should be considered as a possible success path to showing acceptable results.

3. Conclusions
1. The outboard journal bearing would have been able to operate for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> at temperatures of at least 300'F if the AFW pump had been called into service.
2. The available data is not sufficient to say with any certainty that the outboard bearing would have come to an equilibrium temperature less than 300°F had the pump been operated continuously at high feedwater flow rates.
3. The pump very likely could have been operated at lower-than-maximum flow rates and intermittently high flow rates without exceeding 300°F if the pump were intermittently throttled back to minimum flow to allow the bearing to cool.
4. Given that the period of interest is a time when the plant was coming out of an outage, the decay heat load was very likely well below design-basis values. If so, it may be possible to show that full flow capability from the pumps would not have been needed to achieve an acceptable outcome. This possibility should be investigated if it has not already been

,evaluated.

Attachment C

4. References
1. Prairie Island Vendor Manual XH-258-23, "Terry Turbine for AFW Pump," Rev. 12.
2. Neale, M.J., "Bearings-A Tribology Handbook," SAE/Butterworth-Heinemann, 1993.
3. Elwell, R.C., "Ring- and Wick-Oiled Starved Journal Bearings," STLE Tribology Data Handbook, CRC Press, 1997.
4. Connors, H.L., "An Analysis of the Effect of Lubricant Supply Rate on the Performance of the 360 degree Journal Bearing," ASLE Trans, v.5, No. 2, pp. 404-417 (1962).
5. EPRI GS-7352, "Manual of Bearing Failures and Repair in Power Plant Rotating Equipment," 1991.
6. Booser, E. R., Ryan, F. D., and Linkinhoker, C.L., "Maximum Temperature for Hydrodynamic Bearings Under Steady Load," Lubrication Engineering, v. 26, no. 7, July 1970, pp. 226-235.
7. Zeidan, F. Y., and Herbage, B. S., "Fluid Film Bearing Fundamentals and Failure Analysis," Proceedings of the 20th TurbomachinerySymposium, Texas A&M University, College Station, Texas, (1991) pp. 161-186.
8. Khonsari, M. M., and Booser, E. R., "Applied Tribology: Bearing Design and Lubrication," J. Wiley & Sons, (2001).
9. Krzymien, A., and Krzymien, P., "Properties of the Bearing Alloys Used for Slide Bearings of Crank Mechanism," J. of KONES Internal Combustion Engines, v. 11, No. 1-2, (2004), pp. 393-398.

Attachment D Wheelock, Gary From: Huting, Mark Sent: Friday, December 19, 2008 11:30 AM To: Wheelock, Gary

Subject:

FW: PI AFW pump turbine bearing evaluation Original Message -----

From: Greene, Tom [1]

Sent: Tuesday, December 16, 2008 10:57 PM To: Huting, Mark Cc: Harp, Susan; Grant, Will; Zarechnak, Alex

Subject:

PI AFW pump turbine bearing evaluation

Dear Mark:

My letter of December 1, 2008, summarizes our evaluation of the outboard journal bearing of the drive turbine for the Prairie Island Unit One turbine-driven AFW pump. The evaluation was made in support of Xcel's PRA assessment of the significance of an improper insulation installation on the drive turbine. Our evaluation concludes that the bearing would operate for at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> at temperatures up to at least 300F. It also concludes that the available test data does not allow a definitive statement to be made about what the bearing temperature would have been if the pump had been run at full flow until the bearing reached a steady-state condition. Earlier today, we discussed what can be said about what the steady-state temperature of the bearing might be, and you asked that I summarize my thoughts on the issue. My thoughts are as follows:

1. The AFW pump was never run at high flow rate for an extended period in the configuration of concern. This is because the improper insulation was installed during a refueling outage and the issue was discovered during pump testing coming out of that outage. The bearing was still warming up during the early part of the test, and the bearing reached the procedural 220F shutdown limit late in the test after only a few minutes into a high flow run.
2. The temperature history during the test is shown in Figure 2-11 of our report. Around 10:15, the pump is at very low flow, providing about 20 gpm to SG A. The bearing temperature is about 190F and looks to be near its equilibrium temperature for low-flow conditions. The bearing then sees a transient 25 F increase, to about 215 F, as SG B is fed for about 8 minutes. The SG B flow ramped up to about 200 gpm, stayed at that flow for about 5 minutes, and then ramped down to no flow. (SG A was also receiving about 20 gpm during this period.) During the next few minutes, the bearing temperature drops to about 205F while SG A continues to get 20-30,gpm of flow. This data suggests that the steady-state bearing temperature for low-flow conditions (roughly 20 gpm forward flow to the SGs) is between 190F and 205F, probably about 200F.

These results suggest to me that the pump would have been able to provide a SG flow of 50 gpm or so without the steady-state temperature exceeding the normal 220F shutdown limit.

3. My experience with AFW turbine bearings leads me to expect the change in the bearing temperature between minimum-flow operation and full-flow operation would normally be about 40 - 50 F. While it is tempting to add this value to the 200F temperature discussed in item 2 above and estimate the full-flow bearing temperature at about 250F, I believe that would be an underestimate because the normal temperature rise is determined in part by the ease with which heat can be transferred out of the bearing area, and the improper insulation would have tended to reduce heat transfer via one of the normal paths and therefore would have increased the temperature rise above that normally expected. I would expect the bearing temperature at full flow would end up being somewhere above 250F. I would be surprised if it would have reached 300F for the reasons discussed below.
4. Over a period of about 15 minutes after 10:30 (again referring to Figure 2-11 in our report), the flow to SG A is increased in several steps from about 20 gpm to about 100 gpm. Then the AFW flow to SG B is started and stepped up to about 80 gpm. During the next 1

0 ~Attachment D

,6minutes, the combined flow to the two steam generators is about 160 gpm. Over the roughly 20-minute period just discussed, the bearing temperature slowly increased from about 205F to about 220F, at which time the plant operators secured the pump because the bearing temperature was near or at the 220F shutdown set point.

5. The durations of high-flow operation during the test shown in Figure 2-11 are so short that a direct estimate of the steady-state bearing temperature at high-flow conditions is not possible. However, an approximate, indirect estimate can be made as follows.
6. In response to an increase in power dissipation, the bearing and its housing heat up with a first-order (i.e., exponential) characteristic.

This can be seen in Figure 2-11 in the period immediately after start-up, out to about 10:15. The temperature trace from start-up to about 10:15 represents the response of the bearing and housing to a power dissipation increase that results in a bearing temperature rise of 130F: i.e., from 70F to 200F. The initial temperature rise is rapid and the rate of temperature rise steadily decreases over 45 minutes or so as the temperature approaches an asymptotic value, estimated to be about 200F in item 2 above.

7. If the steady-state bearing temperature at full flow were as high as 300F, upon increase in pump flow, the temperature rise above the 200F minimum-flow asymptotic value would have the characteristic of a 10OF exponential. Using 100/130 of the start-up temperature transient as a standard of comparison, none of the temperature rises associated with the flow increases in Figure 2-11 (i.e., those flow increases just after 10:00, at about 10:20, and at about 10:45) appear large enough to be a transient large enough for the bearing to reach 300F. Therefore, it seems to me that the full-flow (i.e.,

160 to 200 gpm) bearing temperature would most likely have been less than 300F if the AFW pump had been allowed to run to steady-state with the improper insulation installed.

Based on items 1 - 7 above, I think you can conclude with high confidence that, had the pump been needed, SG flows of at least 50 gpm could have been provided continuously without exceeding the 220F shutdown limit, and that much higher flows would have been possible, at least intermittently, without failing the bearing. The pump most likely could have run continuously at 160 gpm to 200 gpm without the bearing exceeding 300F, but this is a reasonable inference, based in part on judgment, that doesn't rise to a high-confidence result in my mind.

As we discussed and as I mentioned in my letter of December 1, I expect that full flow capability from the AFW pump may not have been needed during the *period of interest to your evaluation. As I understand it, the insulation of concern was installed during a refueling outage, and the problem was discovered early in the plant start up from that outage before any significant power operations had occurred. With a significant fraction of the core being new fuel having no decay heat, and the remainder of the fuel having sat through the refueling outage, the actual decay heat load during the time of interest should have been well below the design-basis value. Flow sufficient to manage the reduced decay heat load and to cool the reactor coolant system down to decay-heat cut-in conditions probably does not require full flow from the AFW pump. If so, you may be able to gain some PRA margin if it is permissible to take credit for such factors.

I've not run this line of thinking past others here; but tomorrow I will discuss it with a couple of other engineers having relevant background.

In the meantime, if you would like to discuss this issue further, please call.

Regards, Tom Greene 2

Attachment E

_Whee.ock, Gary From: Greene, Tom [tgreene@mpr.com]

Sent: Wednesday, December 17, 2008 3:58 PM To: Wheelock, Gary; Huting, Mark Cc: Harp, Susan; Grant, Will; Zarechnak, Alex

Subject:

Assessment of 11TD AFWP run on 3/16/08 Attachments: SP 1102 3-16-08 data pt 6.jpg; SP 1102 3-16-08 data pt 3.jpg; SP 1102 3-16-08 data pt 5B.jpg SP 1102 3-16-08 SP 1102 3-16-08 SP 1102 3-16-08 data pt 6.jpg ... data pt 3.jpg ... data pt 5B.jpg...

Gary/Mark:

I have had a look at the plots you sent of the IITD AFWP run on 3/16/08.

Here is what I see in the data.

1. The data from about 12:50 to about 13:30 shows the IITD AFWP operating on minimum flow, putting no flow to the steam generators.

(Steam generator feed flow is being provided by the 12 AFWP during this period.) The outboard bearing came to steady state at a temperature of about 212F based on your point 6. The steady-state bearing temperatures with the pump providing forward flow will be higher than this value.

2. The data from your data point 5B, about 12:41, shows IITD AFWP supplying about 35 gpm to the steam generators. The 12 AFWP was off at that time. The bearing temperature was 216F and decreasing. This indicates that the steady-state bearing temperature with 35 gpm of steam generator feed would be less than 216F. Considering this result and item 1, the steady-state temperature would likely have been about 214F.
3. The data from your point 3, about 12:35, shows IITD AFWP supplying about 68 gpm of steam generator flow with the 12 AFWP off. The temperature of the bearing was at 213F and increasing. At that time, the bearing temperature was not near steady state, and judgment suggests that the steady-state temperature would have been above the procedural shutdown limit of 220F.
4. Items 1 - 3 indicate that the maximum steam generator feedwater flow rate for which the steady-state bearing temperature would have been less than the 220F setpoint was more than 35 gpm, less than 68 gpm, and very likely somewhere between 40 and 50 gpm.
5. Using the start-up exponential envelope as a standard of comparison, as was discussed in my note of yesterday, the modest rise of bearing temperature during the period from about 12:14 to about 12:40, a period when 11 TD AFWP was feeding about 68 gpm to the steam generators, suggests that the steady-state bearing temperature at this flow rate, while above 220F would have been well below 300 F.

These results give high confidence that the pump could have delivered at least 40 gpm of steam generator feedwater flow without reaching the alarm point. If much higher flow rates were demanded from the pump, the bearing temperature would exceed 220F. However, at least 68 gpm, and likely much more flow, could be delivered without having the bearing exceed 300F.

If you would like to discuss these observations, please call.

Regards, Tom Greene I

Attachment E - Data Trends for SP 1102,1376, and 1330 Performance March 16, 2008 Engineering Evaluation EC 13630 Point 6 - In reference to Email from Tom Greene dated Wednesday December 17, 2008 1q 2 ItIk __ FT6 - ?I k?

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Attachment E - Data Trends for SP 1102, 1376, and 1330 Performance March 16, 2008 Engineering Evaluation EC 13630 Point 3 - In reference to Email from Tom Greene dated Wednesday December 17, 2008

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Attachment F MAAP4 Analysis of TDAFW Demand For An Appendix R Control Room Evacuation Scenario Introduction This analysis is performed to predict the demand on the Turbine Driven Auxiliary Feedwater (TDAFW) pump in the event of a control room evacuation scenario. For this study, a base case is run that assumes the default decay heat in the MAAP4 parameter file using a steady state core power of 34% at the time of trip. To better refine the risk estimation, an additional study is run that examines the AFW demand for the portion of the power profile that is below 10% power given an AFW pump failure due to overheating.

Inputs and Assumptions The MAAP4 analysis uses the recently updated MAAP 4.0.6 parameter file. The basic sequence of events is as follows;

1. At time zero, the following actions occur;
a. Reactor trip
b. Main feedwater trip and MSIV isolation
c. Reactor coolant pumps are tripped
d. AFW is forced off
2. At twenty-six minutes, the TDAFW pump is started and aligned to one steam generator (referred to as the credited SG). The twenty-six minute delay is to simulate operators performing a manual startup of TDAFW following a control room evacuation.
3. For the base case, the AFW flow is allowed to run at 200 gpm to the credited SG until the nominal SG level is restored. In the sensitivity case, the pump is, forced off after 12 minutes to reflect a pump failure due to overheating.
4. The nominal SG level from the MAAP4 parameter file (ZSEP) is equivalent to a SG level of approximately 48 %NR. It is assumed that the operators will feed the credited SG at the maximum flow rate until this level is achieved.
5. Each case is run for 36 hours4.166667e-4 days <br />0.01 hours <br />5.952381e-5 weeks <br />1.3698e-5 months <br /> demonstrating that the credited SG level equilibrium is achieved.
6. The power ascension history of interest for the MAAP4 analysis is listed in Table 1.

The MAAP4 decay heat model is based on the ANSI/ANS -5.1-1979 standard. It assumes a period of continuous reactor operation followed by a shutdown for which the decay power fraction is calculated as a function of time after shutdown. The model is not setup up to simulate consecutive operating periods (i.e., a long run followed by a refueling shutdown and then a shorter run). Therefore, the analyses are run using the default decay heat parameters.

1

Attachment F The MAAP4 routine calculates a decay heat to power ratio, which depends on the time of irradiation, fuel burnup, and time since shutdown. As documented in the parameter file, the principle parameters are the time of irradiation (TIRRAD) and the average fuel bumup (EXPO). The default values are TIRRAD = 530 days (core average), and EXPO = 31 GWD/MTU. By setting the core power at the time of trip, the decay heat can be modeled for the time period of interest to more closely predict the actual decay heat. The base case will assume a core power of 34% of rated thermal power to estimate the decay heat for the entire period of interest. The sensitivity case will assume a value of 10% power to model the earlier part of the power history.

Note that for the period of time prior to reactor startup on 3/21/2008, the decay heat load is very low. Prior to entry into Mode 3 on 3/15/08, the reactor had been shutdown for 750 hours0.00868 days <br />0.208 hours <br />0.00124 weeks <br />2.85375e-4 months <br />. Decay heat at this point in time is estimated to be on the order of 1.7 MWt (Glasstone - 1981 correlation), which would boiloff approximately 11 gpm of AFW if all the heat were directed into the steam generator liquid.

Therefore, for the time period prior to reactor criticality, the decay heat load would never exceed a 40 gpm demand.

Results Base Case - Power = 34% at time of trip The SG water levels are shown in Figure 1, and the AFW flow profile is shown in Figure 2. An AFW flow of 200 gpm is started twenty-six minutes after the trip.

Level in the credited SG is restored approximately 35 minutes after pump start, and AFW demand begins to decrease. Flow is reduced below 40 gpm at approximately 43 minutes after pump start. Note that the level in the non-credited SG continues to decrease until approximately 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> into the event, providing additional heat removal. AFW demand beyond the initial refill stays below 40 gpm for the duration of the analysis.

Sensitivity Case - Power = 10% at time of trip, AFW Forced Off After 12 Minutes The SG water levels are shown in Figure 3, and the AFW flow profile is shown in Figure 4. An AFW flow of 200 gpm is started twenty-six minutes after the trip, and is forced off after 12 minutes to simulate the failure of the pump. Level in the credited SG is recovered to approximately 32.2 feet above the tubesheet, which is approximately 70% WRS. The level remains stable in both steam generators for the duration of the transient.

The decay power at this initial level does not generate enough heat to require additional AFW flow. Heat losses from the primary system and steam generators to the containment heat sinks are sufficient to reduce the demand for AFW flow.

2

Attachment F This is primarily due to the heat loss to the containment sink seat exceeding the decay heat being produced at this low power level. Figure 5 shows the steam generator pressures decreasing as a result of this heat loss.

Conclusion The majority of the AFW demand is a result of the initial SG level shrink at trip.

The MAAP4 SG level model will use maximum flow until the level is nearly reached. For the base case, this results in a predicted time of AFW flow above 40 gpm of 43 minutes. The actual operation of the AFW pump may result in a lower flowrate, and a corresponding longer time to refill. As can be seen in the AFW flowrate profile after the steam generator level is restored, the demand to meet decay heat needs is less than 40 gpm within two hours of the trip. The non-credited SG level also decreases as it provides some additional decay heat removal, and eventually reaches an equilibrium level before dryout occurs.

In the 10% initial power case, the AFW demand after refill is virtually zero due heat loss to the containment heat sinks. Even with the AFW pump forced off, the SG level stabilizes and the SG pressure (and temperature) decrease. For power levels at or below 10% in this case, the there is sufficient heat loss to the containment to maintain adequate cooling to the primary system even without a continued source of feedwater.

Finally, it is noted that the decay heat load prior to reactor criticality on 3/21/2008 was sufficiently low such that AFW demand would never had exceeded 40 gpm.

3

Attachment F Table 1 - Power History__________

DATE TIME (HRS) POWER () DATE - TIME POWER (%)

03/21/2008 12 0 0.1 03/23/2008 14 50 10.4 03/21/2008 13 1 0.2 03/23/2008 15 51 7.2 03/21/2008 14 2 0.2 03/23/2008 16 52 9.5 03/21/2008 15 3 0.4 03/23/2008 17 53 20.8 03/21/2008 16 4 0.7 03/23/2008 18 54 19.4 03/21/2008 17 5 0.7 03/23/2008 19 55 19.0 03/21/2008 18 6 0.7 03/23/2008 20 56 19.9 03/21/2008 19 7 0.8 03/23/2008 21 57 20.0 03/21/2008 20 8 0.7 03/23/2008 22 58 25.0 03/21/2008 21 9 0.7 03/23/2008 23 59 26.9 03/21/2008 22 10 0.7 03/24/2008 00 60 27.8 03/21/2008 23 11 0.7 03/24/2008 01 61 27.8 03/22/2008 00 12 1.3 03/24/2008 02 62 28.4 03/22/2008 01 13 1.2 03/24/2008 03 63 28.0 03/22/2008 02 14 1.2 03/24/2008 04 64 28.1 03/22/2008 03 15 1.2 03/24/2008 05 65 28.1 03/22/2008 04 16 1.2 03/24/2008 06 66 28.0 03/22/2008 05 17 3.1 03/24/2008 07 67 28.1 03/22/2008 06 18 6.2 03/24/2008 08 68 28.1 03/22/2008 07 19 6.3 03/24/2008 09 69 28.1 03/22/2008 08 20 6.5 03/24/2008 10 70 28.0 03/22/2008 09 21 6.5 03/24/2008 11 71 28.0 03/22/2008 10 22 6.3 03/24/2008 12 72 28.0 03/22/2008 11 23 6.3 03/24/2008 13 73 28.1 03/22/2008 12 24 6.3 03/24/2008 14 74 30.6 03/22/2008 13 25 6.4 03/24/2008 15 75 33.0 03/22/2008 14 26 6.4 03/24/2008 16 76 34.2 03/22/2008 15 27 6.4 03/24/2008 17 77 37.6 03/22/2008 16 28 6.4 03/24/2008 18 78 42.1 03/22/2008 17 29 6.4 03/24/2008 19 79 44.9 03/22/2008 18 30 6.4 03/24/2008 20 80 47.3 03/22/2008 19 31 6.5 03/24/2008 21 81 48.6 03/22/2008 20 32 6.5 03/24/2008 22 82 49.2 03/22/2008 21 33 0.6 03/24/2008 23 83 50.0 03/22/2008 22 34 6.5 03/22/2008 23 35 6.4 03/23/2008 00 36 6.5 _____

03/23/2008 01 37 6.5 _____

03/23/2008 02 38 6.4 03/23/2008 03 39 6.4 03/23/2008 04 40 9.0 03/23/2008 05 41 8.0 03/23/2008 06 42 16.8 03/23/2008 07 43 16.3 03/23/2008 08 44 15.6 03/23/2008 09 45 19.4 03/23/2008 10 46 15.9 03/23/2008 11 47 15.2 032/20 2 48 15.3 03/23/2008 13 49 15.4 4

Attachment F Figure 1 - SG Level - Base Case (Initial Reactor Power = 34%)

Steam Generator Level (Initial Power = 34%)

t ýI, at- iý.Wlt,-, ; `.,j Figure 2 - AFW Flow - Base Case (Initial Reactor Power 34%)

AFW Flow (initial Power 34%)

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Attachment F Figure 3 - SG Level - Initial Reactor Power = 10%, AFW Forced Off Steam Generator Level (Initial Power - 10%, AFW Forced Off)

Figure 4 - AFW Flow - Initial Reactor Power 10%, AFW Forced Off AFW Flow (Initial Power = 10%, AFW Forced Off) 6

Attachment F Figure 5 - SG Pressure - Initial Reactor Power 110%, AFW Forced Off SG Pressure (Initial Power 10%, AFW Forced Off) 7

Attachment G - Data Trends for SP 1103 Performance March 23, 2008 Engineering Evaluation EC 13630 fit IC, A~ _____

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Attachment G - Data Trends for SP 1103 Performance March 23, 2008 Engineering Evaluation EC 13630 t ..... .. ' lC- I ... )

UNIT 1 - PSSA MODE: I AUTO SAS A. ACT St. MI/A POWFR OPF:R aw 12124,,08 F. 10:00:1 3/231018 1 1P2861A 1278.12500 GOOD PSIG 2 1F2861A 201.82880 GOOD GPM 3 1F2871A 0.00000 GOOD GPM 4 1T2867A 169.46875 GOOD DEGF 5 1Y2860D 0.00000 GOOD OPENIOFF

Attachment G - Data Trends for SP 1103 Performance March 23, 2008 Engineering Evaluation EC 13630 UNIT I -PSSA MODE: 1 AUTO A: ACT B: WA POWER OPER T2end NOTtTie

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Attachment G - Data Trends for SP 1103 Performance March 23, 2008 Engineering Evaluation EC 13630

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