ML20149G040
| ML20149G040 | |
| Person / Time | |
|---|---|
| Site: | Peach Bottom  |
| Issue date: | 05/30/1993 |
| From: | Marisa Herrera, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20149G039 | List: |
| References | |
| GE-NE-523-61-04, GE-NE-523-61-4, NUDOCS 9409290340 | |
| Download: ML20149G040 (176) | |
Text
. _ _ _ _ _
GE-NE-523-61-0493 DRF 137-0010-6 FATIGUE EVALUATION OF THE PEACH BOTTOM II AND III REACTOR VESSELS May 1993 Prepared by:
,#1-M.L. lierrera, Principal Engineer Structural Mechanics Projects Approved by:
-...--k, Dr. S. Ranbhiath, Manager Structural Mechanics Projects i
GE NUCLEAR ENERGY San Jose, CA 94o9290340 94o916 PDR ADOCK 0500 7
P
IMPORTANT NOTICE REGARDING CONTENTS OF THIS REPORT The only undertakings of the General Electric Company (GE) respecting information in this document are contained in the purchase order between Philadelphia Electric Company and GE, and nothing contained in this document chall be construed as changing the purchase order. The use of this information by anyone other than Philadelphia E. -ic Company, or for any purpose other than that for which it is intended under such purchasc cder is not authorized; and with respect to any unauthorized use, GE makes no representation or warranty, express or implied, and assumes no liability as to the completeness, accuracy, or usefulness of the information contained in this document, or that its use may not infringe privately owned rights.
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Table of Contents Ses11DIl Eagt 1.
Introductions
1
===2.
Background===
3 2.1 Event Grouping 4
2.2 Severity of Events 4
3.
Analysis Methodology 5
3.1 Existing Event 5
3.2 New Events 6
3.3 Power Rerate 6
4.
Component Evaluation (Without Power Rerate) 7 5.
Power Rerate Evaluation 28 5.1 Power Rerate Scaling Technique 30 5.2 Procedure for Power Rerate Fatigue Evaluation 31 5.3 Component Analysis 32 6.
Summary 34 7.
References 37 Appendix A Vessel Support Skirt Stress Analysis Appendix B Recirculation Inlet Nozzle Fatigue Analysis Appendix C Evaluation of Power Rerate Including Modified Cycles
ACKNOWLEDGMENT The following individuals contributed significantly to the analysis:
Ramon Carey Shane Plaxton William Weitz lb
- 1. INTRODUCTION This report documents the re-evaluation of fatigue for the Peach Bottom Unit 2 and 3 Reactor Vessels. This re-evaluation was performed to incorporate changes in event occurrences at the two plants. In addition, fatigue evaluation considering power rerate in combination with the modified cycles was performed. Stress limits were also evaluated to assure compliance with ASME Code requirements for power rerate conditions and considering new events which were not previously analyzed.
The following recommendations and conclusions are based on the results of the evaluation for the selected components:
Feedwater Nozzle The modified fatigue usage factor is 0.795 considering the modified cycles only. The fatigue usage including the modified cycles and power rerate conditions is 0.894. This fatigue usage occurs in the safe end ahead of the thermal sleeve seals, and is therefore independent of rapid cycling. In order to validate the assumptions made in ta feedwater nozzle analysis, the performance of the thermal sleeve seals must be assessed by either refurbishment / replacement or by obtaining some indication that leakage is limited. This can be done by monitoring leakage in the annulus between the thermal sleeve and nozzle.
Suncort Skirt The modified fatigue usage factor is 0.896 considering the modified cycles only. The fatigue usage factor for 40 years of power rerate and with the modified cycles exceeds 1.0.
However, it is overly conservative to use 40 years of operation at the power rerate conditions since prior to power rerate implementation, the original design basis condition should be used.
The number of years of operation at power rerate can be determined by separating the 40 year design life into years operating at original design basis conditions and number of years operating at power rerate conditions. Results of this calculation showed that the allowable number of years at the power rerate conditions is 21 such that the fatigue usage does not exceed 1.0. In terms of cycles at the rerate condition, this is equivalent to 134 heatup/cooldown + SRV blowdown events,5 extreme heatup/cooldown events and 15 loss of feedwater + HPCI/RCIC events.
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t Refueline Containment Skirt The fatigue usage factor for the refueling containment skirt including the modified cycles is 0.583. The fatigue usage factor with the modified cycles and power rerate is 0.777.
Closure Studs i
J The fatigue usage for the closure stud exceeds 1.0 with and without power rerate considered. Three options are available to address the high usage. The first would be to examine the closure stud analysis and remove any conservatisms. The second is inspect the studs per ASME Code requirements and GE RIC SIL 055 Rev. I recommendations and replace the studs if indications are found. The third is to limit the number of heatup/cooldown events or l
boltup/ unbolt events. The closure stud fatigue usage can be reduced to less than 1.0 if the number of heatup/cooldown events is reduced to 189 (with 120 boltup/ unbolt events) or if the number of boltup/ unbolt events is reduced to 89 (with 216 heatup/cooldown events).
Recirculation Inlet The fatigue usage factor for the recirculation inlet nozzles including the modified cycles is j
0.511. The fatigue usage factor including the modified cycles and power rerate was determined to be 0.549.
Table 1-1 Summary of Fatigue Usage Evaluation Results Modified Cycles Comoonent Orieinal Modified Cveles + Power Rerate Feedwater Nozzle 0.89 0.795 0.894 Vessel Support Skirt 0.55 0.896 0.998 Refueling Containment Skirt 0.33 0.583 0.777 Closure Studs 0.76
>1
>1 Recirculation Inlet Nozzle 0.11 PB 2 0.511 0.549 0.12 PB 3
- Power Rerate for 21 Years Only.
- 2. BACKGROUND Nuclear Reactor Vessels are designed in accordance with the American Society of Mechanical Engineers (ASME)Section III Code. As part of the Section III evaluation, the vessel must meet fatigue design criteria. This criteria states that the cumulative fatigue usage must be less than 1.0 for the expected cyclic loads. The cyclic loads are comprised of mechanical and thermal loads caused by various types of events. The events used in the Peach Bottom 2 and 3 vessel design are given in the thermal cycle diagrams and corresponding design specifications, References 1 and 2, power rerate design specifications and other analyses that have been performed which supersede the original analyses. These documents show the type and number of each event considered in the original fatigue design of the various vessel components.
The original design basis used the best available experience and judgement at the time of the vessel design to specify events. However, with more power plant operation experience, it was noted that actual plant operation differs from that assumed in the design basis. Based on
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operating experience, some events have occurred more frequently than anticipated (e.g.,
startup/ shutdown) and others less frequently (e.g., weekly reduction in power). Therefore, in order to obtain a more realistic fatigue usage factor, a re-evaluation is needed based on actual operating experiences which also considers power rerate conditions. This reconciliation can help justify the modification of cycles and can be an important tool in justifying license renewal.
The original calculations generally contain significant conservatism in the methods used to calculate usage. There are two methods which introduce the majority of the conservatism in the fatigue calculation. It should be noted that when the original calculations were performed, the emphasis was to demonstrate that the fatigue usage was less than 1.0 for the prescribed cycles and that consideration for margin to include additional cycles and license renewal was not included. The two conservatisms generally included in the fatigue calculations are:
1)
Grouping of Events 2)
Severity of Events GE-NE-523-61-0493 2.1. Event Grouping l
In order to simplify and reduce the required effort for the fatigue calculations, events are grouped and included in the fatigue usage calculation by conservatively assuming the same severity as the most severe events in the group. An example of the method is the use of an event with a cool down rate of 400*F/hr to represent a 100*F/hr event. This is obviously conservative since the 400*F/hr event would induce higher stresses and therefore higher fatigue usage.
2.2. Severity Of Events As mentioned earlier, fatigue usage was calculated based on an assumed set of thermal events. In addition to the number of events, the severity (temperature vs. time behavior) was also established using the best available knowledge and judgement.
Reactor experience has shown that generally, the severity is less than was originally established. Since the resulting cyclic stresses are proportional to the temperature range during a particular event, the fatigue usage is over-predicted by using the temperature range in the original design basis.
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- 3. ANALYSIS METHODOLOGY i
In this section, the methodology used to modify the vessel fatigue usage is outlined.
Fatigue evaluation of the vessel is reported in several stress reports. These reports were i
evaluated to determine which components would be the most limiting with respect to the cycle modifications requested. Based on this evaluation and the evaluation performed in Reference 3, and consideration of the cycle changes, the following compeaents were determined to be limiting.
l Table 3-1 Lhniting Fatigue Usage 1
I Comoonent Usage Reference l
Feedwater Nozzle 0.89 4
l Vessel Support Skirt 0.55 5
Refueling Containment Skirt 0.33 5
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Closure Studs 0.76 5
l Recirculation Inlet Nozzle 0.11 PB3 6
0.12 PB2 7
l Note that the CRD Hydraulic Return Nozzle was not included since it has been capped and therefore injection of cold water no longer occurs. Although the feedwater nozzle is limiting with respect to fatigue usage margin, the support skirt, refueling containment skirt, recirculation inlet and closure studs are also significantly impacted by the cycle changes and are therefore included in this evaluation. The recirculation outlet nozzle was not included in this evaluation since it is not significantly impacted by the modified cycles.
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3.1. Existing Event i
If the number of cycles of an existing event is modified, the original or current cumulative fatigue usage is used to determine the new fatigue usage. It should be noted that since the alternating stress for fatigue usage is calculated between two stress states, the total number of cycles must be maintained when performing the event pairing.
As an example, assume event A and event B are paired together to determine the alternating stress. Event A occurs 10 times and event B occurs 20 times. Since there are only 10 events of A, only 10 cycles using events A & B may be used. The remaining 10 events of B must be accounted for in another pairing. If the number of A events is increased to 30, then 20 i
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GE-NE-523-61-0493 cycles of pair A & B are used (limited by B events) and the remaining 10 A events must be used elsewhere.
The modified fatigue usage can be calculated for a new number of cycles (between 2 events) by multiplying the old fatigue usage by the ratio of new cycles to old cycles.
3.2. New Event If the event is not included in the prior fatigue usage calculation, the characteristics of the new event are compared against those analyzed in the original fatigue calculations. If the new event is bounded by an existing analyzed event, the number of new events is added to the number of cycles for the representative events. If the event cannot be bounded by an existing event, then further evaluation is required.
Once all cycle modifications are evaluated, a total summation of the cumulative fatigue usage factor is made. Note, that if the cumulative fatigue usage factor exceeds 1.0, a program to manage and monitor the component is required.
3.3. Power Rerate The calculations discussed in Sections 3.1 and 3.2 are also performed with power rerate considerations. The methodology which includes modifying the stress ranges by the appropriate ratio of rerate pressure and temperature to that of the original basis.
A technique was developed to conservatively scale up the original stresses to account for pressure and temperature increases due to power rerate. Since the shear stress component are typically low, the principal stress directions coincide with the normal stress directions (hoop, axial). Therefore, the magnitude of the principal stress due to pressure is directly proportional to the coolant pressure, and the magnitude of the principal stress due to thermal cycling is proportional to the temperature change during a thermal transient. In addition, the mechanical loads do not change due to power rerate.
Further details on the power rerate stress evaluation is provided in Section 5.
- 4. COMPONENT EVALUATION (Witbout Power Rerate)
In this section, the components listed in Table 3-1 are evaluated to determine the impact of i
the cycle changes given in Reference 8 without power rerate considerations. Thus the following l
information is applicable if power rerate is not implemented. These cycle modifications are l
summarized in Table 4-1.
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In addition, the number of loss of feedwater pump events previously used in the feedwater nozzle evaluation was 48. In this fatigue usage reevaluation, the number of loss of feedwater pump events is modified to 10.
l In the following sections, each component is evaluated to determine the impact of the changes shown in Table 4-1.
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GE-NE-523-61-0493 Table 4-1 Description of Cycle Changes Event Name Existine Event Change 1
Startups (3-4)
Yes Increase from 120 to 216 2
Turbine Roll to 100% (4-5)
Yes Increase from 120 to 216 3
Weekly Imad Reduction (6-7)
Yes Decrease from 2000 to 480 4
Rod Worth Tests (7-9)
Yes Delete event 5
Turbine Trip at 100 % (12-13)
Yes Increase from 40 to 80 6
SRV Blowdowns (14-15)
Yes Increase from 2 to 40 7
Other Scrams (15-16)
Yes Decrease from 147 to 107 1
1 8
Improp. Recirc Starts (17-18)
Yes Increase from 5 to 40 9
Sudden Start Events (18-19)
Yes Increase from 5 to 40 10 Shutdown (19-20, 21-24)
Yes Increase from 118 to 216 11 Excessive Heatups No Add 10 12 Feedwater Temperature No Add 2000 Reduction 13 HPCI/RCIC Injection No Add 20 l
14 Shutdown Cooling In-Service No Add 216 15 Excessive Cooldown Events No Add 10 4.1. Feedwater Nozzle The feedwater nozzle was analyzed for fatigue based on the cycle descriptions given in the thermal cycle diagram and corresponding design specification. The feedwater replacement was l
analyzed to comply with NUREG-0619 requirements on crack growth in 1979, and the results l
are presented in Reference 4. The design specification for the nozzle modification is presented in Reference 9.
The feedwater nozzle is subject to several requirements as stated in NUREG-0619 and the Reference 4 report to assure that cracking in the nozzle blend radius is not a structural integrity concern. For the feedwater nozzle, NUREG-0619 (Table 2) requires that a PT evaluation be -
GE-NE-523-61-0493 performed every 9 refueling outages or 135 startup/ shutdowns whichever comes first and that a UT inspecticn must be performed every other fuel cycle. In addition, the rapid cycling evaluation performed in Reference 4 determined that at fifteen years of service, the seal performance is such that leakage past the seals could be significant. Therefore, in the absence of a program (such as a fatigue / leakage monitor) to assure seal integrity, it is considered prudent to verify the seal performance by inspection.
The design specification provides the cyclic information for the modified nozzle. The number of cycles used in the fatigue analysis differed significantly from that used in the original design analysis. These modifications to the thermal cycle diagram (stated in Section 4.4.1 of Reference 9) are shown in Table 4-2.
Table 4-2 Cycle Modifications In GE Document 22A6647 BL Event Name Change j
1 Startups (3-4)
Increase from 120 to 200 2
less Of Feedwater Pumps (11-12) Increase from 10 to 48 3
Shutdown (19 through 24)
Increase from 118 to 198 4
Hot Standby (21-22) 2600 cycles Note that in Table 4-1, the number of startups were increased from 120 to 216. However, as shown in Table 4-2, the feedwater nozzle has already been analyzed for 200 cycles.
Therefore, only an additional 16 startup cycles must be included in the modified fatigue usage calculation. Likewise, for Shutdown,198 cycles have been already considered. Therefore, an additional 18 shutdown events must be included in the modified fatigue usage calculation.
Table 4-3 from Reference 4 shows the calculation of the fatigue usage factor for the limiting location, node 216. Note that this location is at the safe end ahead of the seals, and is therefore independent of rapid cycling.
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GE-NE-523-61-0493 Table 4-3 Feedwater Nozzle Original Calculation S
Eair Event Pair d # Allow. Cycles
- Desien Cycles Eatieue Usace 1
LFWP.14 116.3 435 130 0.299 Design Hydro 2
Zero Load 96.43 716 62 0.086 LFWP.14 3
Zero lead 72.16 1535 200 0.130 TRCD.14 4
Zero I.oad 69.84 1660 68 0.041 HSB.14 5
TR.14 56.53 2911 436 0.150 HSB.14 6
LFWP9 41.67 7106 144 0.020 HSB.14 7
FHB64.5 36.33 11113 270 0.024 HSB.14 8
TT126 34.92 13118 10 0.0007 HSB.14 9
NOOPM 34.60 13620 1672 0.1227 HSB.14 10 NOOPM 31.46 18557 198 0.0107 SD102.73 6
11 SVB1 6.76
> 10 56462 0.0 NOOPM LFP.14 = less Of Feedwater Pumps at 0.14 minutes, cool-down LFP9 = Imss Of Feedwater Pumps at 9 minutes, warm-up DSNHYDRO = Design Hydro ZEROLOAD = Zero Load TRSD.14 = Turbine Roll at 0.14 minutes, warm-up HSB.14 = Hot Standby at 0.14 minutes, cool-down FHBfA.5 = Feedwater Heater Bypass at 64.5 minutes, warm-up TT126 = Turbine Trip at 126 mmutes, warm-up NOOPM = Normal Operation SD102.73 = Shutdown at 102.73 minutes cool-down SVB1 = Safety Valve Blowdown at I minute, cool-down Below material endurance limit
. - - ~. -. -
-1 GE-NE-523-61-0493 l
For purposes of this evaluation, the number of LOFP events will be reduced to 10.
1 Modification of the cumulative fatigue usage (from Table 4_-3) based on cycle changes (Table 4-1), can be performed by evaluating each of the matched pairs of events (Table 4-3) one-by-one and incorporating any cycle changes.
i 4.1.1. Existine Events Pair 1 1
Pair 1 uses Loss of Feedwater Pumps and Design Hydro for the alternating stress calculation. There are 10 LOFP events and four cooldown events per LOFP event (4x10 = 40 events). In addition,20 HPCI/RCIC events are grouped with LOPP (See Section 4.1.2.3).
There are a total of 130 Design Hydro events. From Reference 4, the number of allowable cycles for this pair is 435. Therefore, the usage due to this pair is:
U1 = 60/435 = 0.138 There are 130-60 = 70 design hydro events remaining.
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l Pair 2 Pair 2 uses Zero lead and Loss of Feedwater Pumps for the alternating stress calculation.
f Since all LOFP events were used in Pair 1, the usage in Reference 4 for Pair 2 is deleted.
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In place of Pair 2 from Reference 4, the remaining design hydro events are paired with the turbine roll cooldown event. There are 256 [216 (startup) + 40 (startup following SRV blowdown)] turbine roll cooldown events. Therefore,70 cycles of turbine roll (cooldown) and design hydro are paired. The number of allowable cycles is 740. Therefore, the fatigue usage for this pair is:
l U2 = 70/740 = 0.095
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l GE-NE-523-61-0493 There are 256 - 70 = 186 turbine roll events remaining for pairing.
Pair 3 l
Pair 3 uses zero load and Turbine Roll for the alternating stress calculation. The number l
of remaining turbine rolls is 186. Since the number of allowable cycles for this event is 1535 (Reference 4), the usage for this pair is:
U3 = 186/1535 = 0.121 The number of possible zero load states is equal to the number of design hydro unloads +
shutdowns + SRV blowdowns = 130 + 216 + 40 = 386. Since 166 zero load events were used in Pair 3, only 386 - 186 = 200 zero load states remain.
Pair 4 Pair 4 uses zero load and Hot Standby (HSB) for the alternating stress calculation. The number of remaining zero load states is 200. Since the number of allowable cycles for this pair is 1660 (Reference 4), the usage is:
U4 = 200/1660 = 0.12 All zero load states have been used in pairing. There are a total of 2600 hot standby events. These HSB events are shared between Pairs 4 through 9. Pair 4 used 200 of the HSB events. Therefore, there are 2600-200 = 2400 HSB events remaining.
Pair 5 Pair 5 uses Turbine Roll and Hot Standby for the alternating stress calculation. The number of Turbine Rolls is increasing from 200 to 216. In addition, the Turbine Generator Trip (warmup), Other Scram", Loss of Feedwater Pumps, and Reactor Overpressure events were grouped into the Turbine roll event in the original fatigue calculation. The number of Turbine Generator Trips is increasing from 40 to 80, and the number of "other scrams" is i
GE-NE-523-61-0493 decreasing from 147 to 107. Table 4-4 summarizes the lumping and modification of cycles for Pair 5.
l Table 4-4 Pair 5 Events Event Old Nm Turbine Roll 200 216 LOFP 48 10 Turbine Trip 40 80 Reactor Overpressure 1
1 Other Scrams 147 107 Turbine Roll after 0
40 l
SRV Blowdown Total 436 454 l
Note that the number of LOFP events is decreased to 10. Also,40 turbine rolls were added I
which occur during startup following an SRV blowdown event.
l Since the total of these events is less than the number of remaining HSB events, the number l
i of cycles for this pair is limited to 454. Subtracting 4$4 from the available 2400 HSB events leaves 1946 HSB events.
O The fatigue usage for Pair 5 can be obtained by multiplying the old fatigue usage by (454/436).
US = (454/436)
- 0.14975 = 0.156 i
Pair 6 l
l Pair 6 uses Loss of Feedwater Pumps (heatup) and Hot Standby for the alternating stress calculation. The number of events is limited by the number of LOFP events. Since there are 10 LOFP events, and 3 heatups per event are considered in this event, the number of LOFP events is 3x10 = 30. In addition, LOFP events are included to consider the 10 excessive heatups and 20 HPCI/RCIC events shown in Table 4-1 (See Section 4.1.2.1). Since the allowable number of cycles for this pair is 7106, the fatigue usage for this pair is:
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GE-NE-523-61-0493 U6 = 60/7106 = 0.008 The number of remaining HSB events is 1946-60 = 1886.
Pair 7 1
Pair 7 uses Feedwater Heater Bypass and Hot Standby for the calculation of alternating 4
stress. In addition, the startup cycles were also grouped with the Feedwater Heater Bypass event. The number of Feedwater Heater Bypass events does not change. However, the number of startups increases from 200 to 216. The number of possible cycles for Pair 7 is limited by the number of Feedwater Heater Bypass + Startups = 286. Since these are paired with HSB, the number of remaining HSB events is 1886 - 286 = 1600. The fatigue usage for Pair 7 is:
U7 = (286/270)
- 0.02430 = 0.026 4
Pair 8 Pair 8 uses Turbine Trip (Event 8-9) and Hot Standby for the alternating stress calculation.
Since the number of turbine trips does not change, the fatigue usage for Pair 8 remains at U8=
0.00076. The number of remaining Hot Standby events is 1600 - 10 = 1590.
Pair 9 Pair 9 uses Normal Operation and Hot Standby (cooldown) to calculate the alternating i
i stress. To account for all remaining hot standby events, the number of cycles for Pair 9 is 1590. In addition,10 events are added to HSB to account for excessive cooldown. This gives a total of 1600 remaining HSB events (1590+ 10 = 1600). The fatigue usage for Pair 9 can be calculated by multiplying the old fatigue usage by (1600/1672).
U9 = (1600/1672)
- 0.12276 = 0.117.
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GE-NE-523-61-0493 Pair 10 Pair 10 uses Normal Operation and Shutdown for calculating the alternating stress. The number of Pair 10 cycles is limited by the number of shutdowns. The old analysis used 198 cycles and is to be increased to 216. The new fatigue usage for Pair 10 is therefore, U10 = (216/198)
- 0.01M7 = 0.012 Pair 11 Pair 11 uses Safety Relief Valve Blowdown and Normal Operation for calculating the alternating stress. Since the alternating stress is below the endurance limit, the fatigue usage factor for this pair is 0.
4.1.2. New Events In this Section, the impact of new events listed in Table 4-1 will be evaluated for its effect on the Feedwater Nozzle.
4.1.2.1. Excessive Heatups in these events, the heatup rate is increased to 160 F/hr from the assumed 100 F/hr. In the Reference 4 analysis, startup events were grouped with Feedwater Heater Bypass. However, for the purpose of this evaluation, the excessive heatup will be grouped with the first 3 heatups of event 11-12 (loss of feedwater pumps). The LOFP heatup begins at 40*F. Heatup to 480*F occurs in 2 minutes. At this point the heatup continues to 561*F in 7 more minutes. It can be seen that these events will produce stresses that will bound the stress produced by a 160 F/hr heatup rate. Figure 1 shows the comparison between the excessive heatup and loss of feedwater pumps event.
Note that if the number of events for LOFP is changed, the fatigue usage for Pair 6 and Pair 9 (calculated in Section 4.1.1) will change. These modifications are discussed below..
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Loss of Feedwater Pumps l
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- Excessive Heattp i
56FF y
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7 s
i 160 F/hr 4(PF i
9 7
9 16 Time (min) o m.
Figtso 1 Comparison of Loss of Foodwater Pumps s,
and Excessive Heattp
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Pair 6 i
Ten excessive heatups are being combined with the existing LOFP events. These were
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incorporated in Pair 6 in Section 4.1.1 l
Pair 9
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Since 10 additional HSB events were used in Pair 6,10 HSB events must be subtracted from Pair 9. This was incorporated in Pair 9 in Section 4.1.1.
l 4.1.2.2. Feedwater Reduction of 20*F 1
In this event, the feedwater flow temperature is assumed to deviate from normal temperatures by 20*F. A approximation of the stress range variation can be obtained by l
assuming a quench of 20*F. The stress caused by this temperature variation is:
a = EaAT/(1-v) = 29(7)(20)/(0.7) = 5.8 ksi This corresponds to an alternating stress of 2.9 ksi. This alternating stress is well below the endurance stress and therefore, the fatigue usage for these events is 0.0.
4.1.2.3. HPCI/RCIC Injection These events are included as part of the IAss of Feedwater Pump event. The additional 20 HPCI/RCIC events are included in Pair 1 and 6 in Section 4.1.1.
4.1.2.4. Shutdown Cooling Based on temperati'r.: traces (Reference 10), the shutdown cooling previously analyzed bounds the observed temperature behavior. Therefore, no change to the fatigue usage is needed.
GE-NE-523-61-0493 4.1.2.5. Excessive Cooldown Events The ten excessive cooldowns can be added to Pair 9. Pair 9 uses the cooldown for the HSB event which is essentially a quench from 546 F to 100 F. This rapid cooldown rate bounds the 160*F/hr excessive cooldown rate. These are included in Pair 9 of Section 4.1.1.
Figure 2 shows the comparison between the excessive cooldown and hot standby events.
4.1.3. Feedwater Nozzle Conclusions Based on the calculations shown in Section 4.1, the fatigue usage factor for the feedwater nozzle remains below 1.0. The modified cumulative fatigue usage for the feedwater nozzle is 0.795. Table 4-5 gives the final revised pairing of events and the resulting fatigue usage.
However, there are some items which must be noted and verified such that these conclusions remain valid. These are discussed below.
Reference 11 gives the results of fatigue usage calculations including the effects of system cycling and rapid cycling. Sheet 7 of Reference 11 shows that locations G, H and I (on ID of Nozzle) are significantly impacted by the rapid cycling. The limiting location used in the calculation above remains the limiting location as long as the thermal sleeve seal refurbishments are performed according to the schedule described in Table 4-6. _.
- - - Excessive Cooldown Hot Stan&y 54@ F 54eP F
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p.__
/
/
x
/
e h
/
160 F/hr Cooldown /g
/
'M 29@ F l e
E i
e 10@ F l
l 0
3 4
7.5 4
Time (n*0 g
Figure 2 Comparison of Hot Stan&y Cooldown and Excessivo Cooldown b,
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GE-NE-523-61-0493 Table 4-5 Feedwater Nozzle Modified Calculation S
Egit Event Pair d # Allow. Cycles
- Desien Cycles Eatigue Usage 1
LFWP.14 116.3 435 60 0.138 Design Hydro 2
Design Hydro 92.03 740 70 0.095 TRCD.14 3
Zero Imd 72.16 1535 186 0.121 TRCD.14 4
Zero Imad 69.84 1660 200 0.121 HSB.14 5
TR.14 56.53 2911 454 0.156 HSB.14 6
LFWP9 41.67 7106 60 0.008 HSB.14 7
FHB64.5 36.33 11113 286 0.026 HSB.14 8
TF126 34.92 13118 10 0.0007 HSB.14 9
NOOPM 34.60 13620 1600 0.117 HSB.14 10 NOOPM 31.46 18557 216 0.012 SD102.73 6
11 SVB1 6.76
> 10 0.0 NOOPM If the seal refurbishments are not performed at the times given in Table 4-6, the rapid cycling fatigue usage could differ from the original calculation. Therefore, it must be verified that the refurbishments have been performed to the above schedule or that the seals are limiting leakage between the thermal sleeve and safe end. If verification of the seal integrity is not performed, then significant fatigue usage is assumed at the nozzle blend radius..
GE-NE-523-61-0493 Table 4-6 Usage Factor Based On Seven Seal Refurbishments Location Seal Refurbishment Year G
H I
12 0.0685 0.0685 0.0523 16 0.0536 0.0536 0.0264 20 0.0848 0.0848 0.0579 24 0.0899 0.0899 0.0647 j
28 0.0858 0.0858 0.0602 32 0.0793 0.0793 0.0526 l
l 36 0.0725 0.0725 0.0449 l
40 0.0663 0.0663 0.0386 l
Total Rapid Cycle 0.601 0.601 0.398 l
Total System Cycle 0.096 0.093 0.082 j
Total Usage 0.697 0.601 0.48 Experience at other BWR plants with the triple sleeve design has shown that the seals are l
performing better than originally assumed in the fatigue calculation. However, the seal performance must be verified on a plant specific basis.
4.2. Vessel Support Skirt The Vessel Support Skirt was analyzed in Reference 5. In Reference 12, an evaluation of the effect of the modified cycles on the suppon skirt fatigue usage was performed. Results of this evaluation demonstrated that the fatigue usage for the support skirt exceeded 1.0 for 40 years. The evaluation in Reference 12 was based on the results from Reference 5 which used conservative methods to evaluate the fatigue usage. In Reference 5, only two events were analyzed. These were,1) Normal Heatup & Cooldown, and 2) Loss of Feedwater Pumps. The Safety Relief Valve Blowdown was considered to be less severe than the Loss of Feedwater Pump and was therefore grouped with the Normal Heatup/Cooldown.
A detailed reevaluation of the vessel support skirt was performed to remove some of the conservatisms included in the original analysis. This reevaluation is summarized in the following sections. The detailed evaluation is included in Appendix A of this report..
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,, -, -7,
,-m-,
l l
I i
The original analysis of this support skirt was performed using a shell analysis with simplified element structures. This method is consistent to provide conservative results.
Therefore, a detailed finite element model was developed to evaluate the stress due to pressure and thermal events.
In We new detailed analysis, three events were evalusted. These three events were: 1) heatup/cooldown, 2) loss of feedwater pump, and 3) excessive heatup/cooldown. Note that the excessive heatup/cooldown event was not evaluated in the original support slit evaluation. The SRV blowdown event was groupsi with the cooldown event due to similarity. The HPCI-RCIC events were conservatively grouped with the loss of feedwater pump.
The thermal stress evaluation was performed using the results of the thermal analysis.
Based on these stress results, the fatigue evaluation was performed using the modified cycles.
Table 4-7 summarizes the stress combination and resulting fatigue usage results. The total fatigue usage for the limiting support skirt location with modified cycles only was calculated to be 0.896.
Table 4-7 Summary of Maximum Calculated Values (All stresses in psi) l i
Limitmg Stress F+Q P+Q+F Ke Salt Cycles Allowable Fatigue Combination Stress Stress Cvcues Usage f
IIU/CD (6.1). HU/CD (26.1) 79862 109813 1.00 106745 246 481 0.511 HU/CD (26.1) - LOFP (41.1) 81174 111552 1.06 114832 10 405 0.025 HU/CD ( 6.1) - EX HU/CD (26.1) 87769 100034 1.39 147522 10 225 0 044 LOFP (16.1) - LOFP (41.1) 91073 123975 1.55 188083 20 124 0.161 LOFP (16.1) - EX HU/CD (6.1) 97904 132705 1.90 246058 10 65 0.154 Total Usage =
0.896 '
GE-NE-523-61-0493 4.3. Refueling Containment Skirt The refueling containment skirt was analyzed in Reference 5. Table 4-8 shows the old and new events used for the fatigue evaluation of the refueling containment skirt.
Table 4-8 Refueling Containment Skirt Events Old # Events New # Events Heatup/Cooldown 120 256 less of Feedwater 10 40 SRV Blowdown 2
40 Design Pressure Test 130 130 Total 282 466 Heatup/Cooldown includes 40 heatups after SRV blowdown. Less of Feedwater includes 20 HPCI/RCTC and 10 excessive heatup/cooldowns.
l The fatigue calculation was performed by grouping all events together, The number of allowable cycles was determined to be 800 in Reference 5. Therefore, the modified fatigue l
usage factor is:
l U = 466/800 = 0.583 l
1 i
l l N -
a GE-NE-523-61-0493 4.4. Vessel Closure Studs The vessel closure studs were evaluated in Reference 5. Event pairs were used depending on bolt location to determine the fatigue usage in the original calculation. However, different event pairs were required in this modification since the number of startup/ shutdowns differs from the number of boltuplunbolt events. Following is the evaluation for the two limiting locations of the stud, bottom outside and bottom inside.
Bottom Outside The event pairs used for the stud bottom outside is shown in Table 4-9.
Table 4-9 Vessel Stud Events Stud Bottom Outside Eair Event New # Events 1
Boltup/ Unbolt 120 2
Heatup/Cooldown 216 3
Design Preload/ Pressure No Effect Test 4
Hydro Boltup/ Unbolt 3+3=6 Stud Bottom Inside Egit Event New # Events 1
Heatup/Cooldown 120 2
Heatup/End of Cooldown 96 3
Boltup/End of Cooldown 120 4
Hydrotest/Zero Load 3
Three additional cycles were added to Stud Bottom Outside Pair 4 since it was determined that the stress cycle due to hydrotest was omitted from the original fatigue calculation.
The modified fatigue usage factors may be determined by multiplying the original fatigue usage by the ratio of new events divided by the number of old events. )
GE-NE-523-61-0493 Ug = 0.353 (no change)
U2 = 216/300 = 0.72 U3 = 0 (No Change)
U4 = 0.009 (6/3) = 0.018 Utotal = 0.353 + 0.720 + 0 + 0.018 = 1.09 l
Bottom Inside j
i f
The event pairs used for the stud bottom inside is shown in Table 4-9. Note that the l
pairing has been performed to consider that there are 120 zero load conditions versus 216 l
heatup/cooldown events.
l I
Using the stress information provided in the original calculations, the fatigue usage may be modified as shown below (all stresses in ksi).
Event Pair Stress Range g
Actual Cveles Allowable Cycles g Heatup-Zero Load 396 198 120 210 0.57 Heatup-End of Cooldown 164 82 96 1000 0.096 Boltup-End of Cooldown 70 35 120 4500 0.027 Hydrotest-Zero Load Unchanged 0.015 Utotal = 0.71 The number of allowable heatup/cooldown events such that the fatigue usage for the bottom outside of the closure stud is less than 1.0 can be determined by reducing the number of heatup/cooldown events (all other events unchanged). The usage per heatup/cooldown event is l
0.72/216 = 0.0033. Therefore, to reduce the usage to less than 1.0, a minimum of l
0.09/0.0033 - 27 heatup/cooldown events must be eliminated. The number of allowable heatup/cooldowns is 216-27 = 189 (for 120 boltup/ unbolt events).
1 1
Similarly, the number of boltup/ unbolts events can be reduced such that the usage remains below 1.0 (all other events unchanged). The usage per boltup/ unbolt event is 0.353/120 =
0.003. The number of boltup/ unbolt events which must be eliminated is 0.09/0.003 = 31.
Therefore, the number of allowable boltup/ unbolt events (with 216 heatup/cooldown events) is 120-31 = 89 (for 216 heatup/cooldown events).
1 1 1 1
t i
GE-NE-523-61-0493 4.5. Recirculation Inlet Nozzle i
The evaluations of the Peach Bottom 2 and 3 Recirculation Inlet Nozzles are presented in References 6 and 7, respectively. Calculations were performed in Reference 12 to include the effects of the modified cycles. In the Reference 12 evaluation, it was noted that some of the l
transients which were not previously analyzed could not be bounded by existing analyzed transients. Therefore, further evaluation of the recirculation inlet nozzles was required to l
include the new unanalyzed transient. The evaluations are summarized in the following sections. The detailed evaluation is contained in Appendix B of this report.
All thermal events which impact the recirculation inlet nozzles were considered and the limiting events were identified. Grouping of the events was then performed to reduce the number of cases required for evaluation. The events evaluated were:
l 1)
Loss of Feedwater Pumps 2)
Normal Operation j
3)
Sudden Start of Pump in Cold Recirculation Loop.
i i
The stress evaluation was performed based on the results of the thermal analysis. Using these 1
stress results, the load combinations and resulting fatigue usage was calculated. Table 4-10 i
summarizes the results for the limiting location in the recirculation inlet nozzle. The maximum l
fatigue usage was determined to be 0.511.
1 s
i
)
i 4
i j l
~
v
Table 4-10 FATIGU E S UMMARY FOR S ECTION 7-0 (All stresses in psi) 17250.
3SM = 51750.
E = 28300000.
KT =
1.350 SM =
STRESS STRESS PEAK P+Q THERMAL P*Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANCE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 SOPCRLHU-ZEROLOAD 43098.
113561.
-71634.
45569.
3.333 261098.
40 100
.400 1 LOFP-CDO-NORMOP2
- 58024, 68402.
-29733.
39869.
2.073 131015.
10 737.
.014 1 LOFP-CDO-SCRAMHUP 57722.
68202.
-35969.
33615.
2.060 129682.
40 761.
.053 1 SCRAMCDN-ZEROLOAD 48385.
58334.
-17317.
41701.
1.424 75990.
180 5003.
.036 1 SCRAMHUP-EXCESSCD 48422.
55657.
29785.
27010.
1.252 65136.
10 9315.
.001 1 SRV8LDWN-ZEROLOAD 44929.
54840.
-15462.
39988.
1.199 59814.
40 13896.
.003 h1SCRAMHUP-ZEROLOAD 39782.
48959.
1.000 44371.
130 63137.
.002 1 NORMOP1 -ZEROLOAD 36143.
44986.
1.000 40564.
130 103190.
.001 1 SHUTDOWN-LOFWHHUP 36162.
42454.
1.000 39308.
80 122360.
.001 1 SHUTDOWN-FWTEMPRD 34350.
42102.
1.000 38226.
136 142336.
.001 2 NORMOP1 -SOPCRLCD 26143.
39441.
1.000 32792.
,40 352028.
.000 1 LOFWHCDN-FWTEMPRD 23617.
32803.
1.000 28210.
.all OT
$a a
m 8
- 5. POWER RERATE CONSIDERATIONS In this section, a summary is presented for the components of interest considering both the modifiec cycles and power rerate conditions. Appendix C of this report contains the details of l
the power rerate calculations. In this section, the term "new" refers to the conditions considering both the modified cycles and power rerate. The term "old" refers to the condition considering only the modified cycles. For example, Salt,new is the alternating stress considering both modified cycles and power rerate conditions.
^
The power rerate conditions which differ from the original requirements are shown in l
Table 5-1.
I I
i i
I I
I 1
I i l
l l
Table 5-1 Power Rerate Conditions l
Operating Condition Affected Region Event Change Descriotion l
Region A Normal Operation Pressure increased from 1000 psig to 1038 psig Regions A, B, and C Normal Operation Temperature increased from 546 F to 551*F Region A Scram Increase following pressure by 38 psig.
Increase corresponding temperatures as shown below.
Original Power Rerate Original Power Rerate Pressure Pressure Temp.
Temp.
(Psig)
(psig)
(*F)
( F) l 1180 1218 573 573 875 913 530 536 1125 1163 561 565 1000 1038 546 551 l
665 738 500 506 930 968 538 542 Region B and C Normal Operation Temperatures change as follows: 522 F to 527*F, 538 F to 542*F, 500*F to 506*F, 546*F to 551*F,512*F to 517 F.
Recirculation Outlet Rated Flow 43,500 to 49,700 gpm Recirculation Inlet Rated Flow 8,700 to 9,940 gpm 6
6 Steam Outlet Rated Flow 3.35x10 to 3.81x10 lb/hr Feedwater 376*F to 387*F 5550 to 5890 gpm j
Other changes to design documents are detailed in Appendix C.
GE-NE-523-61-0493 5.1. Power Rerate Scaling Technique A technique was developed to conservatively scale up the stresses to account for pressure and temperature increases due to power rerate. Since shear stress is typically low in vessel j
calculations, the principle stress directions coincide with the normal stress directions.
Therefore, the magnitude of the principal stress is directly proportional to the coolant pressure, and the magnitude of the principal stress due to thermal cycling is proportional to the temperature change during the thermal event. Since there are usually no changes in the mechanical stresses due to power rerate, the new (power rerate) value for principal stress is:
ototal, new apressure,old * (Puew/Pold)
=
+ athermal,old * (ATnew/ATold)
+amechanical or: ototal, new opressure,old * (SCF)p
=
+ athermal,old * (SCF)T
+ 0 mechanical where (SCF)p
= pressure scaling factor (SCF)T
= thermal scaling factor AT
= (final transient temperature) - (initial transient temperature)
In addition, a scaling factor was also included to consider the effect of increased flow. If increased flow conditions require it, the thermal scaling factor is:
(SCF)T
= (ATnew/ATold) * (SCF)p The stress intensity values are scaled up by the larger factor of (SCF)p and (SCF)T. This is done since the pressure and thermal principal stresses are typically not reported. Therefore, this method of scaling up to obtain the power rerate stress is considered conservative.
GE-NE-523-61-0493 5.2. Procedure for Power Rerate Fatigue Evaluation The following procedure was used to calculate the power rerate cumulative fatigue usage factor (Unew) for the limiting location on the RPV component of interest.
(1)
Multiply the original P+Q+F stress intensity values by the scaling factor for each stress cycle to obtain the power rerate P+Q+F.
(2)
For each of the limiting stress cycle pairs used in the fatigue analysis of the original governing stress report, determine the absolute value of the difference of the power rerate P+Q+F stress intensities (calculated in Step 1). This value is S +Q+F,new-P (3)
Determine the power rerate alternating stress intensity, Salt,new, for each of the original limiting stress cycle pairs as follows:
Salt,new = (0.5)
- Ke,new * (Ec/Ea)
- S +Q+F P
where Ke,new
= simplified elastic-plastic factor
= 1.0, Sn,new < 3Sm,new
=1 + (1-n)/(n(m-1)) * (Sn,new/(3Sm,new) - 1), for 3Sm,new < Sn,new < 3mSm,new
= 1/n, Sn,new > 3mSm,new (Ec/Ea)
= elastic modulus correction factor (4)
Use Salt,new as the new value of the ordinate when entering the applicable design fatigue curve to find the corresponding allowable number of cycles (N,new) for i
each of the limiting stress cycle pairs.
I (5)
Calculate the power rerate incremental fatigue usage factor (U,new = ni/N,new) i i
for each of the limiting stress cycle pairs, where ni s the lesser of the actual i
number of design cycles for each pair. The lesser is used because the value of the P+Q+F stress intensity range for the limiting stress cycle pair is only experienced by the component over the lesser number of cycles.
i j l r
i GE-NE-523-61-0493 (6)
Calculate the power rerate cumulative fatigue usage factor (Unew = EU,new)- If i
Unew < l.0, the ASME Code requirement is met.
5.3. Component Analysis (with Power Rerate)
In this section, the results of the power rerate calculations with the modified cycles is summarized. Details of the calculations can be found in Appendix C of this report.
5.3.1. Closure Studs The resulting fatigue usage for the closure studs is 1.09. This fatigue usage obviously exceeds the ASME Code limit of 1.0. Since these studs are replaceable, an effective managing program would be to either replace the studs or rely on ASME Code Section XI required inspection to monitor the stud integrity.
Since the usage considering power rerate plus modified cycles is essentially the same as that for the modified cycles only (no rerate), the number of allowable heatup/cooldown and boltup/ unbolt events are the same as reported in Section 4.4. The number of allowable heatup/cooldown events is 189 (for 120 boltup/ unbolt events). The number of allowable boltuplunbolt events is 89 (for 216 heatup/cooldown events).
5.3.2. Feedwater Nozzle The location of highest cumulative fatigue usage on the feedwater nozzle is on the outside of the safe end near the upstream end of the thermal sleeve. The fatigue usage factor calculated for the feedwater nozzle considering the modified cycles without power rerate was 0.795 (See Section 4.1). The fatigue usage was due primarily to system cycling and not rapid cycling.
This assumes that the thermal sleeve seal is effective in limiting leakage into the annulus region.
Power rerate is expected to yield no increase in rapid cycling.
Results of the calculations including the power rerate conditions indicate a fatigue usage of 0.894. Therefore, the fatigue usage is less than the 1.0 maximum. !
l l
l GE-NE-523-61-0493 5.3.3. Suonort Skirt The reevaluation of fatigue including power rerate conditions was based on the results presented in Section 4.2 of this report. Details of the power rerate calculations are presented in Appendix C of this report.
Results of the power rerate calculations, including the effects of the modified cycles results in a fatigue usage of 1.09 for 40 years of operation at these conditions. However, it should be noted that this is extremely conservative since operation until power rerate is implemented are at the original design basis conditions. Therefore, it is appropriate to base the fatigue usage factor on the number of years of operation at the original design basis conditions and the number of i
j years at the power rerate conditions. If the operation conditions are differentiated in the fatigue j
usage calculation, the number of years of operation at the power rerate conditions can be determined by setting the resulting fatigue usage factor to 1.0.
When this is done, the number of years of operation at power rerate and original design basis which results in a fatigue usage factor less than 1 is 21 and 19 years, respectively. In terms of cycles at the rerate condition, the following table shows the allowable number of events.
Allowable Number of Cycles Events Orieinal Condition Rerate Condition Total Heatup/Cooldown + SRV Blowdown 122 134 256 Extreme HU/CD 5
5 10 less of Feedwater Pumps + HPCI/RCIC 15 15 30 5.3.4. Recirculation Inlet Nozzle l
The reevaluation of fatigue including the power rerate condition was based on the results l
presented in Section 4.5 of this report. Details of the power rerate calculations are presented in Appendix C of this report. Results of this evaluation showed that the fatigue usage including power rerate and modified cycles was 0.549. l l
- 6.
SUMMARY
A fatigue evaluation of the Peach Bottom Unit 2 and 3 Reactor Vessel was performed. The fatigue usage was modified for changes in existing event frequencies, consideration of new events. Components for evaluation were selected by determining those components with the smallest margin, and those which are affected most by the event revisions. The following conclusions are based on the results of the evaluation for the selected components:
Feedwater Nozzle The modified fatigue usage factor is 0.795 considering the modified cycles only. The fatigue usage including the modified cycles and power rerate conditions is 0.894. This fatigue usage occurs in the safe end ahead of the thermal sleeve seals, and is therefore independent of rapid cycling. It should be noted that the original calculation was based on seal refurbishments occurring at specific times during operation. The first refurbishment was analyzed at 12 years.
Six additional refurbishments over the remaining operating years were also included in the evaluation. In order for the modification of the fatigue usage to be valid, the assumptions in the original calculation must be verified.
In order to validate the assumptions made in the analysis, the seals performance must be assessed by either refurbishment or by obtaining indications that leakage is limited. This can be done by monitoring leakage, temperature and pressure in the annulus between the thermal sleeve and nozzle.
It should be noted that recently a domestic plant was successful in using fatigue monitoring and outside surface Ultrasonic Examination to remove the requirements for I'T of the feedwater nozzle blend radius. In addition, monitoring of the nozzle can give a realistic indication of the fatigue usage since it is independent of event occurrences and relies solely on the conditions at the nozzle blend radius.
Field experience is indicating that the seals are performing better than assumed in the analysis. However, this must be verified on a plant specific basis..
l 1
GE-NE-523-61-0493 If seal performance is not assessed, then the analysis indicates that significant fatigue damage can occur and the limiting location would be in the annulus between the nozzle and thermal sleeve. This would invalidate the limiting fatigue usage of 0.795 and rapid cycling effects would need to be further evaluated.
Support Skirt The modified fatigue usage factor is 0.896 considering the modified cycles only. The fatigue usage factor for 40 years of power rerate and with the modified cycles exceeds 1.0.
However, it is overly conservative to use 40 years of operation at the power rerate conditions since prior to power rerate implementation, the original design basis condition should be used.
The number of years of operation at power rerate can be determined by separating the 40 year design life into years operating at original design basis conditions and number of years operating at power rerate conditions. Results of this calculation showed that the allowable number of years at the power rerate conditions is 21 such that the fatigue usage does not exceed 1.0. In terms of cycles at the rerate condition, this is equivalent to 134 heatup/cooldown + SRV blowdown events,5 extreme heatup/cooldown events and 15 loss of feedwater + HPCI/RCIC events.
Refueling Containment Skirt The fatigue usage factor for the refueling containment skirt including the modified cycles is 0.583. The fatigue usage factor with the modified cycles and power rerate is 0.777.
Closure Studs The fatigue usage for the closure stud exceeds 1.0 with and without power rerate considered. Three options are available to address the high usage. The first would be to examine the closure stud analysis and remove any conservatisms. The second is inspect the studs per ASME Code requirements and GE RIC SIL 055 Rev. I recommendations and replace the studs if indications are found. The third is to reduce the number of heatup/cooldown events or boltup/ unbolt events. The closure stud fatigue usage can be reduced to less than 1.0 if the number of heatup/cooldown events is reduced to 189 (with 120 boltup/cooldown events) or if the number of boltup/ unbolt events is reduced to 89 (with 216 heatup/cooldown events).
Recirculation Inlet The fatigue usage factor for the recirculation inlet nozzles including the modified cycles is
)
]
0.511. The fatigue usage factor including the modified cycles and power rerate was determined i
to be 0.549.
i l
a 3
i l
i i
i 1
4 s
- t.
- 7. REFERENCES 1.
GE Drawing 729E762, " Reactor Thermal Cycles" 2.
GE Drawing 135B9990, " Nozzle Thermal Cycles" 3.
Peach Bottom Units 2 and 3 Reactor Vessel Thermal Cycle Fatigue Assessment, GE Document SASR 85-54, MDE-166-0785, DRF 137-0010 July 1985.
4.
GE Document 22A6647, Stress Report, Reactor Vessel - Feedwater Nozzle,1979.
5.
Peach Bottom Vessel Stress Report, GE Order No. 205-B1156, B&W Contract No 610-0139-51,'1970.
f 6.
GE Document 23A4274, Certified Stress Report, Recirculation Inlet Nozzle and Safe End, 1984.
7.
GE Document DC23A5720, Certified Stress Report, Recirculation Inlet Nozzle and Safe End,1987.
8.
Memorandum, 1-20-2820-213, from Mary Hyslop, " Peach Bottom Reactor Vessel Thermal Cycle Reanalysis", January 2,1992.
9.
GE Document 22A6656, Design Specification, Reactor Vessel (System Cycling).
(
j
- 12. Fatigue Evaluation of the Peach Bottom II and III Reactor Vessel - Phase 1, GE Document i
GE-NE-523-30-0392, DRF 137-0010-5, Rev.1, April 1992.
i l
l
- i y
.e c-.,.
e
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 l
I l
\\
o, i i
APPENDIX A Peach Bottom II and III Vessel Support Skirt Stress Analysis 1
I l
i n
4 A-1 0
l APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 l
Table of Contents Page 1.0 Introduction
1.1 Background
A-3 1.2 Purpose A-3 2.0 Assumptions and Units A-4 3.0 Analysis 3.1 Thermal Analysis A-5 3.2 Stress Analysis A-8 3.3 Fatigue Usage Analysis A-9 4.0 Results A-21 l
5.0 Summary and Conclusions A-36 l
l 6.0 References A-37 l
i l
A-2 w
y
-ew
- - w+c p~?e p-amm-'r-+ + + + -,Mq
'M-yar+--r-
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6
1.0 INTRODUCTION
1.1 BACKGROUND
i The support skirt was originally designed for fatigue to a set of thermal cycles. The j
original stress repon, conducted by Babcock and Wilcox (Reference A-1), showed that this component could withstand the designated number of cycles with a fatigue usage of less than 1.0, as required by the AShE Code (Reference A-2). Based on Reference A-3, operating transients were redefined using actual plant data for Units II and III. As seen in Reference A-3, cycle prediction for the startup, shutdown, and single relief valve blowdown events were higher than the original design basis. In addition several new events were also identified. Those effecting the suppon skirt region are the excessive heatup, excessive cooldown and HPCI-RCIC Injection events. The analysis summarized by this appendix re-evaluates the support skirt fatigue usage by means of a thermal, stress, l
and fatigue analysis applied to a detailed finite element model of the support skirt geometry.
Unless otherwise stated, any section numbers stated refer to those in this appendix.
1.2 PURPOSE This appendix documents the detailed stress analysis of the RPV bottom head - support skirt region that was performed to determine the stresses and the corresponding fatigue usage factor for Peach Bottom Units II and III.
The vessel and support skirt was modeled and analyzed using the general purpose finite element computer code ANSYS version 4.4A. The thermal transients which were analyzed give a temperature distribution for each time step. The thermal transients were analyzed to determine the limiting temperature gradients and the corresponding thermal stresses were calculated. The stresses due to the thermal gradients, as well as the pressure loads and the mechanical loads from the original support skirt analysis (Reference A-1) were used to determine the fatigue usage.
A-3
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 2.0 ASSUMPTIONS AND UNITS The assumptions used in this analysis are summarized below:
- 1. An axisymmetric modelis used since the support skirt and pressure vessel are circumferentially similar.
- 2. The heat transfer coefficients of regions 1,2 and 3 are constant.
- 3. The stub tubes are omitted from the finite element model since their presence has an insignificant effect in the area ofinterest.
- 4. The most severe transients considered were startup, shutdown, excessive heatup,
)
excessive cooldown and the loss of feedwater pump events.
i
- 5. The units for dimensions are inches, stresses are ksi, pressure is psi, time is hours, and heat transfer coceficients are Btu /hr-fP-F.
A-4 1
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 3.0 ANALYSIS The original analysis of the support skirt was performed using a shell analysis with a
)
l simplified element structure. To determine the thermal and pressure stresses in the area of interest, a detailed Snite element model was developed to re-evaluate the stresses in the j
support skin.
3.1 THERMAL ANALYSIS l
3.1.1 MODEL l
The finite element computer code ANSYS Version 4.4A was used to develop the axisymmetric model of the bottom head of the vessel and the vessel support skirt. The I
model dimensions are shown in Figure A-3-1, while the finite element model is shown in hgures A-3-2 and A-3-3. The model is composed of two materials. The support skirt knuckle is composed of SA533, Gr-B, Class I low alloy steel, while the vessel, support skirt, and base plate are composed of SA302,.Gr-B low alloy steel. An air pocket is also included in the region below the knuckle to bottom head junction.
l 3.1.2 THERMAL TRANSIENTS The governing thermal transients for the analysis were bounded by startup, shutdown, excessive heatup, excessive cooldown and the loss of feedwater pump events. The startup and shutdown events were combined and analyzed as one thermal transient, the loss of feedwater pump event was analyzed as another thermal transient and the excessive heatup and excessive cooldown were analyzed as a third event. The safety relief valve blowdown event is very similar to the shutdown event, thus the number of cycles for shutdown event shall be increased to account for the SRV blowdown thermal transient.
To be conservative, the cycles corresponding to the HPCI-RCIC event are grouped with the most limiting transient, the loss of feedwater pump. The thermal cycles are obtained from Reference A-4 (Region C). Figures A-3-4, A-3-5 and A-3-6 show the thermal transients.
A-5
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 The startup and shutdown events were separated into the following load steps.
Step 1. Steady state at 100 F Step 2.100 F to 546 F at 100'F/hr Step 3. Steady state at 546*F i
Step 4. 546 F to 375"F at 100 F/hr i
Step 5. 375*F to 330 F in 10 minutes Step 6. 330*F to 100*F at 100*F/hr The excessive heatup and excessive cooldown events were separated into the following load steps.
Step 1. Steady state at 100 F Step 2.100 F to 546 F at 160*F/hr Step 3. Steady state at 546*F Step 4. 546*F to 375 F at 160 F/hr Step 5. 375'F to 330'F in 10 minutes Step 6. 330 F to 100'F at 160*F/hr The loss of feedwater pump event was separated into the following load steps.
Step 1. Steady state at 522'F Step 2. 522'F to 300*F in 3 minutes 40 seconds i
Step 3. 300 F to 100 F in I hr Step 4. Steady state at 100 F Step 5.100 F to 300'F in 5 minutes Step 6. 300 F to 546 F at 100*F/hr Step 7. Steady state at 546'F 3.1.3 HEAT TRANSFER COEFFICIENTS The heat transfer regions for this model are shown in Figure A-3-7. The primary heat transfer from the vessel is due to conductive and convective heat transfer modes. The convective heat transfer coefficients for each region are described below. A summary of these coefficients for each region is provided in Table A-3-1.
l A-6
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 Region 1 The insulation on the support skin is reflective metal, therefore, based on Reference A-5, the heat transfer coefficient, h, is constant and equals 0.2 Btu /hr-n2-F. The temperature used for region 1 in the thermal analysis was conservatively assumed to be 70 F.
)
Recion 2 1
The heat transfer coefficient for Region 2 is estimated by the original stress report (Reference A-1) to be 500 Btu /hr-fl2-F.
R.ecion 3 The insulation on the outside of the reactor pressure vessel is reflective metal, therefore, based on Reference A-5, the heat transfer coefficient, h, is constant and equals 0.2 Btu /hr-fl2.op, j
Recion 4 The original stress repon assumed perfect insulation in this region. However, based on similar analyses, as well as the formulation given in Reference A-6, the heat transfer coefficient of the bottom head of the vessel is considered to vary as a function of the temperature difference between the film and surrounding temperature. Based on previous analyses, is conservatively taken as 0.21 ATl/3 Region 5 An air pocket was included in the thermal analysis to account for the trapped air as shown in Figure A-3-7. Conductive, convective and radiative heat transfer modes are assumed to exist in this area. A temperature dependent conductive heat transfer coefficient for the air is calculated based on effective convection heat transfer coefficients for both free convection and radiation heat transfer (Refterence A-7). The resulting coefficient is a temperature dependent conduction term, and is not included in Table A-3-1.
A-7
1 APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 i
3.1.4 TEMPERATURE RESULTS j
1 The thennal analysis was performed using ANSYS version 4.4A to obtain the nodal temperature time histories. This temperature output was input into the subsequent stress analysis (described in Section 3.2) which calculated element stresses, strains, and displacements.
j i
The temperature distributions from the ANSYS thermal runs were checked for the mest critical temperature difference across a section of the vessel skin. These sections were identified by node pairs. The times of maximum temperature difference, obtained by inspection of the results, were found to be at the end of startup, at the end of shutdown, at the end of excessive heatup, at the end of excessive cooldown, after the initial cooldown in the loss of feedwater pump event and after the heatup following the loss of feedwater pump event.
Based on the results of this analysis, the limiting section was a horizontal cross section located in the suppon skirt knuckle region, just below the elevation at which the knuckle fonns an inner corner for connection to the bottom head (see Figure A-3-3). This location also corresponds to the limiting section in the original analysis.
3.2 STRESS ANALYSIS 3.2.1 MODEL AND ASSUMPTIONS The finite element model used for this stress analysis was identical to that of the thermal analysis, except the element type used in the analysis was ANSYS STIF 42. In addition to the stresses induced by the temperature gradients found in the previous section, this model was utilized to find the stresses due to internal pressure of the vessel.
3.
2.2 DESCRIPTION
OF MODEL AND ANALYSIS CASES The suppon skirt is supponed at the bottom of the base plate to the pedestal, thus the model was fixed (in both the x and y directions) at that reference location. The nodes at the top of the model were coupled in the venical direction to simulate a generalized plane strain condition.
A-8
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 The thermal stresses were determined for all load steps and iterations. These results were reviewed and the stresses which produce the largest stress range occurred at the end of either a heat up or cooldown transient.
The pressure case was analyzed using 1000 psi applied along the inner face of the bottom head. Additionally, a force was applied at the top of the model to represent the axial stress in the vessel due to internal pressure. Figure A-3-8 shows the boundary conditions and applied load for the pressure case.
Based on the stress output for these transients, as well as the mechanical loads, this analysis generated a combined Snite element stress distribution and subsequent membrane and bending stress for each stress component ( e, o, oxy, o,). The maximum calculated x
y primary + secondary (P+Q) stress was combined with the appropriate finite element peak stress components to find a maximum primary + secondary + peak (P+Q+F) stress value.
The membrane stress was calculated by averaging the stress distribution across the section. The bending stress was found by determining an equivalent linear stress distribution across the skirt thickness as shown in Figure A-3-9. The peak stress is deSned as the stress difference between the finite element surface stress and the membrane plus bending stresses.
3.3 FATIGUE USAGE ANALYSIS The combined P+Q stresses and the combined finite element stresses described in Section 3.2 of this Appendix were used to determine P+Q+F stresses for the fatigue usage calculation. The P+Q+F stress components were calculated using the stress concentration factor:
c(PQF) " CM) OQ-1) + FE where FE is the finite element stress component K is the stress concentration factor (Reference A-1) i This value was then multiplied by the elastic-plastic fatigue factor, K, resulting in a final e
stress range. It should be noted that the finite element analysis does simulate some of the stress concentration, therefore, the c(p9r) is conservative.
A-9
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 The alternating stress intensity for the limiting stress combinations was calculated from this stress range. These alternating stress intensity was then used in conjunction with Table I-9.1 of Reference A-3 to determine the number of allowable cycles. The comparison of actual cycles to this value results in a fatigue usage value. The limiting stress combinations and the calculation of the corresponding fatigue usages are shown in Table A-4-1. The total fatigue usage for the support skirt is calculated to be 0.896.
Because the usage is below 1.0, the ASME Code requirement is satisfied.
l l
A-10
1 APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 Table A-3-1 Summary ofHeat Transfer Coefficients l
l Heat Transfer Heat Transfer Coefficient, h Boundary Fluid Region (Btu /hr-ft2-F)
Temperature ( F) 1 0.2 70 2
0.2 100-545 3
0.2 70 4
0.21 ATIU 160 A-11
l APPENDDC A GE-NE-523-61-0593 DRF 137 0010-6 a
-+-
131.75 R 125.5 R 143.375 REF 133.5 R
/
- 1.25
[
3.75 113.75 e
111.0
,'- 12.5 l
Figure A-3-1 RPV Support Skirt Configuration A-12 1
Figure A-3-2 Finite Element Model ANSYS 4.441
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APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 i
l l
(*F) 546*
S S
100* /hr
(
i 5 0 0 ---
330' 400 ~ 100*/hr 300 --
100* /hr 2 0 0 -.-
100 -
4.46 1.71 167 2.3 7
Time (hr)
Figure A-3-4 Startup and Shutdown l
l A-15
l APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 j
l l
l
(*F) 546*
i S
160* /hr 5
i 500 -
375*
N 400 -- 160*/hr 330*
300 -
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(
200 -
l 100 -
2.79 1.07.167 1,44 Time (hr)
I Figure A-3-5 Excessive Heatup and Excessive Cooldown I
A-16 l
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6
(*F) l 522*
500 100 */hr 400 -
300 -
300*
300*
200 -
0.083 S
S 100 -
1 l
k k
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l Time (hr) i l
Figure A-3-6 Loss of Feedwater Pumps l
A-17 i
l Figure A-3-7 IIcat Transfer Regions e
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APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 l
h PEAK STRESS L,
BENDING STRESS
/
U h
MEMBRANE STRESS 0.
U ACTUAL STRESS DISTRIBUTION EOulVALENT LINEAR STRESS SECTION EQUIVALENT
=
THICKNESS MEMBRANE p
STRESS Figure A-3-9 Linearization Of Stress Distribution Across A Section Thickness l
l l
l 1
l A-20 1
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 4.0 RESULTS The temperature distributions corresponding to the stresses of the load steps which were used in the fatigue calculation are contained in Figures A-4-1 through A-4-6. The 4
temperature distribution figures are (a) at the end of startup, (b) at the end of the shutdown event, (c) at the end of excessive heatup, (d) at the end of the excessive cooldown event,(e) at the end of the cooldown for the loss of feedwater pump (LOFP) event, and (f) at the end of the heatup for the LOFP event respectively. These thermal distributions represented the limiting conditions for each of the events analyzed.
Figure A-4-7 shows the axial thermal stresses at the end of startup for the entire model and includes a magnified view of the stresses in the skirt knuckle region. Figure A-4-8 shows the axial thermal stresses at the end of shutdown. Figure A-4-9 shows the axial thermal stresses at the end of excessive heatup. Figure A-4-10 shows the axial thermal stresses at the end of excessive cooldown. Figure A-4-11 shows the axial thermal stresses at the end of the cooldown for the LOFP event. And Figure A-4-12 shows the axial thermal stresses at the end of the heatup for the LOFP event. Figure A-4-13 shows the axial stresses due to the pressure load. Table A-4-1 summarizes the maximum stresses in 1
the knuckle region (section composed of nodes 260 through 264).
Based on the stress evaluation performed in this analysis, the total fatigue usage factor was calculated to be 0.896. Thus, the total fatigue usage factor is below the ASME allowable of 1.0.
J I
i d
4 4
A-21 4
APPENDIX A GE-NE-323-61-0593 DRF 137-0010-6 i
Table A-4-1 Summary ofMaximum Calculated Values Luniting Stress P+Q PQ+F Ke Salt Cycles Allowable Fatigue Combination Stress Stress Cycles Usage HU/CD (6,1) - HU/CD (26,1) 79862 109813 1.00 106745 246 481 0.511 i
HU/CD (26,1) - LOFP (41,1) 81174 111552 1.06 114832 10 405 0.025 HU/CD ( 6.1) EX HU/CD (26,1) 87769 100034 1.39 147522 10 225 0.044 LOFP (16.1) - LOFP (41.1) 91073 123975 1.55 188083 20 124 0.161 LOFP (16,1) - EX HU/CD (6,1) 97904 132705 1.90 246058 10 65 0.154 Total Usage =
0.896 i
l l
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A-22
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Figure A-4-1 Temperature Distribution - Startup I
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Figure A-4-2 Temperature Distribution - Shutdown 1
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Figure A-4-8 Axial Thermal Stresses - Sliutdown ANSYS 4.4A1 38 43 0
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N ANSYS 4.4A1
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Figure A-4-10 Axial Thermal Stresses - Excessive Cooldown Yi N
ANSYS 4.4A1
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MAY 4 1993
-[aI 17:49:27 l
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?a TIME =17.92 D'
SY (AVG)
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Figure A-4-11 Axial Thermal Stresses - LOFP (Cooldown) 97 i'
7 ANSYS 4.4A1 F /
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MAY 4 1993 1,
18:10:24
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POST 1 STRESS dl STEP =16
/i 9
ITER =1 j
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TIME =1.061 a
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k S GLOB i
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SMN =-27683
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t
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9 $20 5
'x PEACH BOTTOM SUPPORT SKIRT, LOFP 52118
l l
l Figure A-4-12 Axial Thermal Stresses - LOFP (IIcatup) 7 ANSYS 4.4A1 c
i MAY 4 1993 L
T Uf 18:12:35 T
/
W !'
F POST 1 STRESS l
[
STEP =41
.3
,/.
-u i
b' ITER =1
^'
TIME =4.291 N' $
i
/ P rl SY (AVG)
' \\ '<,
S GLOBAL
'\\
DMX =0.188586 u,
'/\\
SMN =-70209 naw+)2 SMN8=-81135 J
- ,' e 19 SMX =38457 m
it l],,
SMX8=42135 g
N 9*f'!
2V
=1 i
y) f DIST=97.006 X
hi'
)/
XF
=G5.875 de
^
l x
YF
=88.188 t"
i;
-70209 0
n ()
-58135 9
-46061 a
o r'
U -33987 M -21913 I'J E -9839
't' 9x 2235 S
26 d
,l 38457 ll WIND =2 0
'l 2V
=1 ll
- DIST=3.673 i
- XF
=114.057 W
mYF
=76.501 6
-70209 o
pl
-58135 5
Nbb87
~
l 2235 7
X dM h6b8h PEACH BOTTOM SUPPORT SKIRT, LOFP 38457
Figure A-4-13 Axial Stresses - Pressure Case uf R~
ANSYS 4.401
[
MAY 4 1993
^
Y I,_
. sx 18:19:57
[
l i
t s'
POST 1 STRESS
- '_ '${
.F-STEP =1 I
N ITER =1 I
i SY (AVG)
'/i v?y t
t S GLOBAL
- n I
4 PA3 DMX =0.08249 SMN =-8015
~
tml SMN8=-9345 SMX =15328 4
SMXB=20681
.u j
h 2V
=1 d
,/
DIST=97.006 lig ' wm e XF
=SS.875 2
g
.s As
/
YF
=88.188 l
d
//
-8015
-5421 o
6 i
-2827 71 j
j h
M g.534 4954 F;
F 7547 y
1 735
~,
c 15328 g
l l
WIND =2 ZV
=1 c
I
- DIST=3.155 y
- XF
=114.093 I
- YF
=7G.14 um l
-8015
?
-5421 8
r l
-2827 5
5.1 5360 7547 X
EM lj735 PEACH BOTTOM SUPPORT SKIRT ANSYS MODEL 15328
APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6 5.0
SUMMARY
AND CONCLUSIONS A detailed stress analysis was performed for the Peach Bottom Units II and III RPV support skirt region. Based on the thermal transients, the bounding thermal stresses were l
used to determine the fatigue usage. In conjunction with the mechanicalloads, the fatigue l
usage for the revised 40 year duty cycle was calculated to be 0.896, less than the allowable value of 1.0.
A-36 1
l
l APPENDIX A GE-NE-523-61-0593 DRF 137-0010-6
6.0 REFERENCES
A-1 Babcock and Wilcox Report No. 610-0139-51, Report #8, Stress Analysis -
Support Skirt, (GE VPF No. 205B1156).
I A-2 AShE Boiler and Pressure Vessel Code,Section III, Nuclear Vessels,1980 Edition.
A-3 Memorandum,1-20-2820-213, from Mary Hyslop, " Peach Bottom Reactor Vessel Thermal Cycle Reanalysis", January 2,1992 l
A-4 GE Drawing No. 729E762, Reactor Thermal Cycles, Revision 0.
A-5 GE Document No. 21 A9477, Revision 6, RPV Base Specification.
A-6 Kreith, F., " Principles of Heat Transfer", Third Edition, Harper and Row Publishers, New York,1973.
A-7 GE Document No. DRF 137-0010, SASR 89-24, " Evaluation of Support Skirt j
Fatigue Usage for Quad Cities Nuclear Station Units 1 and 2", March 1989.
A-8 GE Drawing No. 129388, Support Skirt Assembly and Details, Revision 7.
l A-9 GE Drawing No. 129371, List ofMaterial, Revision 12.
A-10 GE Drawing No. 886D499, Reactor Vessel - Purchase Part Drawings.
1 A-37
GE-NE-523-61-0593 DRF 137-0010-6 May 1993 APPENDII B
Peach Botton Units 2 and 3 Recirculation Inlet Nozzle Fatigue Evaluation 9
GE Nuclear Energy
I 1
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 TABLE OF CONTENTS 1.
SCOPE 2.
SUHMARY AND CONCLUSIONS 3.
DESIGN 4.
ANALYSIS 4.1 Thermal Transient Analysis 4.1.1 Thermal Model 4.1.2 Selection and Description of Thermal Transients 4.1.3 Thermal Boundary Conditions 4.2 Stress Analysis 4.2.1 Primary Stress Analysis Outside the Area of Reinforcement 4.2.2 Primary Plus Secondary Stress Analysis 4.3 Fatigue Analysis 4.3.1 Stress Concentration Factors 4.3.2 Cumulative Usage Factor 5.
RESULTS 5.1 Thermal Transient Analysis 5.2 Primary Plus Secondary Stress 5.3 Fatigue Evaluation 6.
REFERENCES B-2 j
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 FIGURES B-3-1 Recirculation Inlet Nozzle and Thermal Sleeve Geometry for Peach Bottom Unit 2 B-3-2 Recirculation Inlet Nozzle and Thermal Sleeve Geometry i
for Peach Bottom Unit 3 B-4-1 Finite Element Model of Peach Bottom Unit 2 B-4-2 Node Pairs Selected for Temperature Comparison i
B-4-3 Sudden Start of Pump in Cold Recirculation Loop Transient Thermal Model Diagram B-4-4 Thormal Boundaries B-4-5 Primary Plus Secondary Stress Analysis Areas of Interest B-4-6 Resolution of Stress Distribution Across a Section B-3
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 TABLES B-2-1 Sammary of Stress and Usage Results for Peach Bottom Unit 2 Recirculation Inlet Nozzle B-3-1 Recirculation Inlet Nozzle and Safe End Material Properties
)
B-4-1 Thermal Transients Applicable to the Recirculation Inlet Nozzle l
B-4-2 Constant Mechanical and Hydraulic Scale Factors B-4-3 Expressions to Determine Variable Mechanical and Hydraulic Scale Factors B-4-4 Variable Mechanical and Hydraulic Thermal Stress Scaling Factors B-4-5 Themal Stress Scaling Factors B-4-6 "ffsetive Stress Concentration Factors used in Fatigue Analysis B-5-1 stal P+Q Stress Ranges 1
I l
B-5-2 Primary, Secondary, and Peak Stresses for Section 3-I B-5-3 Primary, Secondary, and Peak Stresses for Section 3-0 B-5-4 Primary, Secondary, and Peak Stresses for Section 4-I l
t B-5-5 Primary, Secondary, and Peak Stresses for Section 4-0 B-5-6 Primary, Secondary, and Peak Stresses for Section 6-I B-5-7 Primary, Secondary, and Peak Stresses for Section 6-0 l
l B-5-8 Primary, Secondary, and Peak Stresses for Section 7-I l
I l
l l
1 l
l B-4
Appendix B GErNE-523-61-0593 DRF 137-0010-6 TABLES (continued)
B-5-9 Primary, Secondary, and Peak Stresses for Section 7-0 B-5-10 Fatigue Usage Factors
]
B-5-11 Fatigue Summary for Section 3-I B-5-12 Fatigue Summary for Section 3-0 B-5-13 Fatigue Summary for Section 4-I B-5-13-1 P+Q Intensity Ranges for Section 4-I
)
B-5-13-2 Peak Intensity Ranges for Section 4-I B-5-14 Fatigue Summary for Section 4-0 B-5-15 Fatigue Summary for Section 6-I B-5-16 Fatigue Summary for Section 6-0 B-5-17 Fatigue Summary for Section 7-I B-5-17-1 P+Q Intensity Ranges for Section 7-I B-5-17-2 Peak Intensity Ranges for Section 7-I B-5-18 Fatigue Summary for Section 7-0 B-5-18-1 P+Q Intensity Ranges for Secti.on 7-0 B-5-18-2 Peak Intensity Ranges for Section 7-0 B-5
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 1.
SCOPE This appendix documents an ASME Boiler and Pressure Vessel Code,Section III, analysis of a recirculation inlet nozzle, safe end, and thermal sleeve for the Peach Bottom Units 2 and 3 BWR reactor pressure l
vessel. The analysis was necessitated as a result of new thermal i
transients and changes in number of cycles of occurrence of some thermal transients provided in Memorandum 1-20-2820-213 [B-1) based on plant i
operating experience not accounted for in the current stress report (B-2).
The affect of power rerate is not accounted for in this appendix.
Design specification 23A4158, Revision 3 (B-3) is currently applicable.
2.
SUMMARY
AND CONCLUSIONS The results of the analysis are summarized in Table B-2-1.
These results show that the analyzed components (nozzle, safe end, and thermal sleeve) meet the appropriate ASME Code,Section III allowables for the conditions defined in the scope of this appendix for Peach Bottom Units 2 and 3.
B-6
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 TABLE B-2-1
SUMMARY
OF STRESS AND USAGE RISULTS FOR PEACH BOTTOM UNIT 2 RECIRCULATION INLET NOZZLE l
e I
i i
i I
l l
l l
l ALLOWABLE l l
MAXIMUM l
l l
l l
lREF. l VALUE VALUE l
l l
(ksi) l (ksi) lPAGE l l
CATEGORY LOCATION CONDITION t
f f
f f
1 I
i i
i i
i i
i lPL+PB+Ql l
l lB-31 l Level A&B l
Nozzle 51.9 80.1 l
l l
lB-31 l l68.5/40.2 l
51.8 Safe End l
l l
l Clad l
lB-31 l 42.1 50.1 l
1 l
l 1
I i
i i
i i
i i
I l
Fatigue lLevelA&B&C l Nozzle l
l lB-40 l 0.002 1.0 l
Usage l
l Safe End l
0.511 l
lB-40l 1.0 l
l l
l 0.107 l
lB-40 l Factor Clad 1.0 l
l l
l 1
I i
- With/without thermal bending B-7
]
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 t
3.
DESIGN 3.1 DescriDtion of the Recirculation Inlet Nozzle Assembly.
f The recirculation inlet nozzle and safe end assembly connects the f
reactor pressure vessel with the recirculation piping. The nozzle is joined to the vessel by a weld.
A welded single thermal sleeve is used to connect the safe end to the jet riser.
The safe end and thermal sleeve are made of corrosion resistant stainless steel and the nozzle is a low alloy j
steel with stainless steel cladding.
Dimensions and materials are documented in [B-3] for Unit 2 and [B-4] for Unit 3 and provided in 1
i Figures B-3-1 and B-3-2, respectively.
The primary difference in the i
geometry of Unit 3 components as compared to Unit 2 components is the new transition piece of Unit 3, added to transition from a smaller inner diameter safe end to the thermal sleeve.
Table B-3-1 also summaries some material properties applicable to Units 2 and 3; the material properties to 1
which these analyses are based were taken from (B-5].
i l
f TABLE B-3-1 KATERIAL PROPERTIES FOR PECIACULATION INLET NOZZLE ANALYSIS 1
l DESIGN STRESS MINIMUM YIELD MINIMUM TENSILE l
INTENSITY, Sm STRESS, Sy STRESS, Su LOCATION MATERIAL iksil AT 575'F iksi)
(ksi) l Nozzle SA-508 26.70 50 80 3
Forging Class 2 Safe End SA-182 17.25 30 70 Grade F316 5
Pipe SA-358 17.25 30 75 Type 316 Thermal SA-312 16.70 30 75 Sleeve Type 304 Cladding SA-240 16.70 30 75 Type 304 or ASTM A371 Type ER308 B-8 f
N Appendix B GE-NE-523-61-0593 DRF 137-0010-6 l
i t
1
]
- ASME CODE LOW ALLOY STEEL. -
L 125 50 R MIN SECT 3 NOZ2LE SA 508 CL U CL ASS 1
_10.18 MAX 2 2.19
=
9 51 WN 2.94 7.31 -
3 50 R
- 613 - ~ - 19 STN STL REPL ACEMENT PIPE l
ISUPPLIED BY OTHERS)
EXISTIE STN CLADONG 12"NPS SCH 100 STL BUTTER (50 R TYPE 316 REPLACEMENT 4D 4.25 R 2 25 R ExtSilm STN STt A~
p THERMAL SLEEVE f/
SA 312 TYPE 304 7 c' ~ww
//
ll j;
13 87 e
[I SRt 25.00 0 10 75 e oe It507 e 10.02 e 110660 e g
l --.
.f 12.8 4 l
q MIN e NON CODE 25 MN ti.D ONLU 13*.
l SECT gygg7 gg BOUNDARY LOW-AltOY SR1 = LOCATION FOR CALCULATION STL NOZZLE OF STRESS RULE NDEX
(
NOTE: NOZZLE & THERMAL SLEEVE DiuENSONS
/
V ARE REFERENCE DIMENSONS N NCHES DIMENSIONS IN PARAENTHSES ARE AS-BULT E xlS T!NG 1
MlN e DMENStONS FROM APPLCABLE FDDR'S g; f ygx 125 e 1
I
. VIEW A i
FIGURE B-3-1 RECIRCULATION INLET NOZZLE AND THERMAL SLEEVE GEOMETRY FOR PEACH BOTTOM UNIT 2 B-9
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 16.14 10.0D
[ WELD I
i ASf1E CODE SECTION !!!
.19 SST CLAD -
BOUNDRY e
MEW SST LW ALLOY REPLACEMENT 0.738 nga SUTTER -
- P IPE W ALL STEEL 1250 PSIC NOZZLE BY OTHERS OSSIGN PRESSURE SA 508 CL 2 l
4 W
Eunie Sn N
TMiglAL SLEEVE -i
-a
/[
l l
54 312 TTPE 304 0.69 1500 PSIC NEW DESIGN PftESSURE TRANSITION 11.507 DIA PIECE 12* NPS 10.75 DJA I
'I*
2 OIA 10.02 DIA ASME CODE ROUNDRV ER RT LT l
FIGURE B-3-2 RECIRCULATION INLET NOZZLE AND THERMAL SLEEVE GEOMETRY FOR PEACH BOTTOM UNIT 3 i
I B-10 l
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 3.2 structural Desian Reouirements The structural adequacy of the recirculation inlet nozzle and safe e r. d assembly is demonstrated by the structural analysis results being w2 thin the allowable design criteria and limits.
New thermal cycling as discussed in (B-1), in addition to, the loading as defined in (B-3) will be considered for the structural evaluation.
l l
4.
ANALYSES comparing geometry, stresses, and usage factors of Unit 2 and 3 nozzle components using sources (B-2),(B-3),(B-4), and [B-6) indicated that Unit 2 would be the bounding case to analyze.
The analyses were performed in accordance with Subsection NB of (B-5), as required by (B-3).
Thermal transient analysis consisted of evaluating the thermal events listed in Table B-4-1.
The stress analysis combined the effects of thermal, l
mechanical, pressure, and hydraulic loading. A detailed finite element I
model (FEM), shown in Figure B-4-1, of Unit 2 nozzle, safe end, and thermal l
sleeve was developed for use in the thermal and stress analyses. A fatigue i
i i
ovaluation was performed using the cumulative damage method and the I
stresses obtained from finite element analysis (FEA).
4.1 Thermal Transient Analysis ANSYS FEA was used to determine and evaluate the temperature distributions in the nozzle and safe end caused by thermal transients.
ANSYS output FILE 04.DAT supplied temperature-time history for each node of the model. A POSTl option in ANSYS was used to specify node locations and compute a temperature difference between nodal pairs at each progressive time step (iteration).
Critical times were chosen on the basis of i
comparing temperature differences between preselected nodal pairs to l
determine at what times during the transient the thermal gradients are greatest. The nodal pairs, shown in Figure B-4-2, were located primarily along or across the high stressed regions of the FEM.
B-11 i
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 iN. ~~
/
r i
i ll
,"~~
- L.____ '!
1 t
-l i
[
.___.__ j i
i l
1 I
I i
l I
-d i
j y,/
r i
11 i
i
.\\~
,I l [%.. l
/
-J,N i I
Wg * \\,
<N f~1 N
l
\\ \\,/ N. %j 'y
/,/
ENTIRE MODEL
/
'a
/
\\ /
f
/ l/
\\
W
,//
7)
~
//
\\
\\
l' jy l
j h
\\
k
\\.f' z
w nummmmusum m usammusum X
l PJ SAFE END ENLARCED VIEW
/
"l l
s#
,p j
~j t
I
/
t
~
)
I
=3 7
^
5
==
i j
\\.
w%,
f__
l FIGURE B-4-1 FINITE ELEMENT MODEL OF PEACH BOTTOM UNIT 2 RECIRCULATION INLET NOZZLE B-12 I
l I
s
l l
l Appendix B GE-NE-523-61-0593 DRF 137-0010-6 P
ll i
N0 5 3
5 7
I N/
s sis Ts hos 697 470 g
A h
NODE PAIRS SELECTED FOR TEMPERATURE COMPARISON FIGURE B-4-2 B -13
-- _ - - - - ~
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 All of the thermal events that affect the recirculation inlet nozzle were considered but not all were analyzed; only the bounding events were analyzed. This is discussed further in Section 4.3 of this appendix.
The effects of the less severe thermal transients were included in the total fatigue usage factor as described in Section 4.3.
Along with the FEM, the thermal analysis required supplying boundary (fluid bulk) temperatures, film coefficice.ts, material properties, and solution time steps as input.
4.1.1 Thermal Model The FEM used for the analysis reported in (B-2) with a refined mesh in the region of the full radius in the safe end was used for this analysis. The detailed FEM was generated in ANSYS using [B-7) two-dimensional axisymmetric isoparametric thermal solid elements called STIF55. A mesh plot of the FEM is shown in Figure B-4-1.
A portion of the reactor vessel was modeled.
The curvature of the vessel portion modeled rotated 360' forms a spherical surface area; however, the actual vessel geometry at the recirculation nozzle location is cylindrical. The effect on the temperature solutions in the regions of interest by using a spherical surface area to approximate the actual cylindrical surface area of the vessel is insignificant. The thermal properties used for the vessel, nozzle forging, safe end, and recirculation piping were input for various temperatures between and including 0*F and 600*F so that an interpolation would be performed to determine the temperature solutions.
These properties were selected from [B-5) and [B-8).
4.1.2 Selection and Descriotion of Thermal Transient Thermal events that impact the recirculation inlet nozzle are specified in [B-3) and listed in Table B-4-1.
Some events produce no significant thermal stresses in the nozzle assembly and would have no bearing on the maximum range of stress intensity; therefore, a thermal analysis of each event was not necessary.
Also, some transients have more severe degrees or rates of temperature change than other transients and therefore bound the conditions of other transients.
The bounded transients would not be analyzed because the thermal stresses produced would be less severe as ccepared to another transient. The thermal transients selected for thermal analysis are listed below.
B-14
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 1.
Loss of Feedwater Pumps (LOFP) - Event 3 2.
Normal Operation (NORMOP) - Event 7 3.
Sudden Start of Pump in Cold Recirculation Loop (SOPCRL) - Event 18 The LOFP transient and the NORMOP transient were analyzed previously (B-2].
The SOPCRL transient was analyzed as illustrated in Figure B-4-3.
4.1.3 Thermal Boundary Conditions The thermal boundary conditions are based on the requirements (B-3).
Figure B-4-4 defines the thermal boundary regions. The summary below describes the basis for the heat transfer tilm coefficients "h" and the bulk fluid temperatures for each region. All regions not defined by a region number are considered to be insulated.
Heat Transfer coefficient Region:
Basis:
1 Conventionally calculated "h" values for 2
forced and natural convection conditions 3
4 --
5 "h" value per Section 4.3.4 of [B-3)
Bulk Fluid Temperature Region:
Basis:
1 Recirculation water temperature 3
Reactor water temperature 5
100*F per Section 4.3.4 of [B-3)
The forced and natural convection heat transfer film coefficients were computed based on equations provided in Appendix 10 of [B-2]. Three conditions apply:
(1) forced convection in a pipe, (2) natural convection in an annulus or for the annular surfaces, and (3) natural convection for all other surfaces flat or circular.
B-15
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 100% FLOW 100 % -
600 -
522 F f
500-e RECIRCULATION 400_
7 TEMP.
/
50 % -
300-
/
/
200 -
/
130 F 100 -
/
0%-
0
=
A A
I l
522 F 3
E 4
u_
w REACTOR m
p TEMP.
m m
w ae y.,
I II TIME (min)
O
.65 1.0 NOTES:
STEPS ARE MODELED AS.083 W4. (5 SEC.) RAWS
--- % RATED FLOW TN RECRC. NCZZLE TEW. OF FLLO FIGURE B-4-3 SUDDEN START OF PUMP IN COLD RECIRCULATION LOOP TRANSIENT THERMAL MODEL DIAGRAM B-16
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 r
l t
re 5
l
/
\\
N
/
//
//'
\\ \\W
=4'gsb,-
/
\\
\\
\\/
.a;;;;
- i' j
b
=
=
=
x
-~
FIGURE B-4-4 THERMAL BOUNDARIES B-17
l l
l Appendix B GE-NE-523-61-0593 DRF 137-0010-6 i
l TABLE B-4-1 THERMAL TRANSIENTS APPLICABLE TO THE I
RECIRCULATION INLET NOZZLE CONSIDERED TRANSIENT TRANSIENT I.D.
IN fB-21 Design Hydro Test DESHYDRO Yes Loss of Feedwater Pumps LOFP Yes Start up START-UP Yes Normal Operation NORMOP1 Yes Shutdown SHUTDOWN Yes Scram SCRAM Yes Loss of Feedwater Heaters LOFWH Yes Partial Feedwater Heater Bypass PFWHB Yes Safety Relief Valve Blowdown SRVBLDWN No l
Sudden Start of Pumps in Cold SOPCRL No Recirculation Loop Excessive Heat up EXCESSHU No Excessive Cool down EXCESSCD No Feedwater Temperature Reduction FWTEMPRD No Shutdown In Service SHUTDNSV No HPCI/RCIC Injections HPRC-INJ No I
l B-18
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 i
4.2 Stress Analvsig A stress analysis was performed to show compliance with Subsection NB-3220 of [B-5) for the region of a nozzle outside the " area of reinforcement".
FEA was used to evaluate secondary and peak stresses due to thermal gradients of the SOPCRL transient; these results were combined with the FFI. stress results due to thermal gradients, internal pressure, reacticas from attached piping, and thermal sleeve loads.
The resulta not ar.alyzed in th'.s appendix were analyzed in [B-2).
The original area of reinforcement analysis found in [B-9) showed compliance oith Subsection NB-3332 through NB-3338 of [B-10).
The effect ct ther:nal loads from the Sudden Start of Pump in Cold Recirculation Loop ttsnaient can be shown to not change the outcome of the area of reinforcement fatigue usage factor meeting allowable limits, as revaluated in Section 4.2 of [B-2), by removing the conservatism of the lump sum method used in [B-9) and applying the cumulative method based on the outcome of the results of fatigue usage outside the area of reinforcement recorded in this appendix.
4.2.1 Primary Stress Analysis outside the Area of Reinforcement The analysis that was previously performed [B-2) is still I
applicable because the mechanical and pressure loads have not changed. The primary stress results in Table B-2-1 of this appendix reflect the results presented in Section 5.1 of [B-2).
4.2.2 Primary Plus Secondary Stress Analysis l
Primary plus secondary (P+Q) and peak stress analysis is required I
for Level A and B events according to Subsection NB-3222 and NB-3223, respectively of [B-5).
Subsection NB-3224 indicates that Level C events need not satisfy Secondary and Peak stress requirements; the Sudden Start of Pump in Cold Recirculation Loop and the Shutdown Cooling in Service thermal transients are Level C events.
The FEM described in Section 4.1.1 was used to evaluate P+Q and peak stresses; however, the thermal element STIF55 was replaced with the two-dimensional axisymmetric isoparametric structural element STIF42.
B-19
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 The P+Q stresses were considered in three separate cases mechanical load stresses, pressure stresses, and thermal stresses.
The mechanical loads stress results, pressure load stress results, and LOTP and NOP. MOP thermal stress results were obtained from Appendix 20 of [B-2).
The SOPCRL thermal stress results are found in Section 5.2 of this appendix.
Select locations were chosen to evaluate P+Q and peak stresses.
The FEA component stresses for each case were resolved into membrane, bending, and peak stresses across the selected locations of interest.
Stress intensity ranges for the three cases were summed up to yield the total P+Q stress intensity range.
4.2.2.1 Mechanical Stresses Mechanical stresses were evaluated by FEA in [B-2), and the 1
results were used in this appendix.
The nozzle, safe end, and thermal i
sleeve were loaded by mechanical loads specified in Tables 1 and 2 of
[B-3).
Each load was applied to the FEA as a unit load (e.g. 1,000 psi and 100,000 lb).
Actual stresses are obtained by multiplying unit stresses by appropriate scaling factors. The scaling factors were used to scale for mechanical stresses and hydraulic load stresses.
Scaling factors in Table B-4-4 were computed from the expressions found in Table B-4-3. Tables B-4-2 and B-4-4 were used to calculate mechanical and hydraulic load stresses respectively for each transient.
Tables B-4-2 and B-4-3 were obtained from [B-2).
4.2.2.2 Pressure Stresses The pressure stresses analysis was performed previously in [B-2).
The reactor pressure scaling factors and the adder pressure scaling factors i
were determined based on information provided in specifications [B-3).
B-20
~
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 l
4.2.2.3 Thermal Stresses
(
Thermal stresses for each event listed in Table B-4-1 were determined by scaling the stress of the appropriate " reference" transient.
The scaling factors listed in Table B-4-5 were determined based on the temperature change associated with a transient with respect to a reference transient.
The reference transients are those selected for analysis in Section 4.1.2.
Stress results for LOFP and NORMOP thermal transients were i
l obtained from [B-2).
Thermal stresses were determined at each critical time using the FEM described earlier.
The applied boundary condition at the end surface nodes of the vessel material allows movement along the vessel radial direction but restricts movement in the tangential (hoop) direction. The boundary conditions at the end of the recirculation pipe and thermal sleeve are the generalized plane strain conditions.
The i
elastic material properties input to the FEM were varied with temperature, except for Young's modulus.
Young's modulus was set at the 70*F value consistent with the value used in the appropriate fatigue curve in Appendix I of (B-5).
l l
l B-21
TABLE B-4-2 CONSTANT MECHANICAL AND HYDRAULIC SCALE FACTORS SE SE TS TS TS TS DELTA CASE SHEAR MOMENT SHEAR MOMENT AXIAL TORSION PRESS PRESS I
I CONSTANT LOADS
.000
.000
.020
.042
.114
.000
.000
.000 i
E TABLE B-4-3 EXPRESSIONS TO DETERMINE VARIABLE MECHANICAL AND HYDRAULIC SCALE FACTORS
$a SE SE TS TS TS TS DELTA CASE LOADS IDENT.
SHEAR MOMENT SHEAR MOMENT AXIAL TORSION PRESS ALL CONDITIONS TS+SE DEAD WGT.
.201
.901
.015
.0146 0
0 0
o NORMALOP ONLY TS+SE OBE
.512
.924
.005
.0127
.005
.0072 0
Y FACTOR *(T-70)/(575-70)
SE THERMAL EXPN.
.458 1.099 0
0 0
0 0
FACTOR * % FLOW /100 TS HYDRAULICS O
O
.087
.085
.215 0
0 E
FACTOR * % FLOW /100 DELTA PRESS O
O O
O O
O
.26 "a
t ALL CONDITIONS TS THERMAL LOADS O
O O
O O
O O
E W
e.
NOTES:
- 1. Scale factors do not apply to ZEROLAOD condition 8
O t
t; Y
8 a;
a
I TABLE B-4-4 VARIABLE MECHANICAL AND HYDRAULIC THERMAL STRESS SCALING FACTORS TEMP PRESS SE SE TS TS TS TS DELTA TRANSIENT
- F
% FLOW (pai)
SHEAR MONENT SHEAR MOMENT AXIAL TORSION PRESS PRESS i
I LOFP-CD0 300 0
875
.4096 1.402
.0150
.0146
.0000
.0000
.8750
.0000 l
LOFP-CD7 300 7
875
.4096 1.402
.0211
.0206
.0151
.0000
.8750
.0182 LOFP-NUO 546 50 875
.6327 1.937
.0585
.0571
.1075
.0000
.8750
.1300 i
START-UP 552 50 1050
.6381 1.950
.0585
.0571
.1075
.0000 1.050
.1300 MORMOP1 528 100 1050
.6164' 1.898
.1020
.0996
.2150
.0000 1.050
.2600 l
f NORMOP2 528 100 1050 1.128 2.822
.1070
.1123
.2200
.0072 1.050
.2600 SHUTDOWN 100 50 0
.2282
.9663
.0585
.0571
.1075
.0000 0.000
.1300 E
SCRAMCDN 300 100 950
.4096 1.402
.1020
.0996
.2150
.0000
.9500
.2600 m
l SORAMHUP 528 100 1125
.6164 1.898
.1020
.0996
.2150
.0000 1.125
.2600 LOFWHCDM 496 12 1050
.5874 1.828
.0254
.0248
.0258
.0000 1.050
.0312 O
LOFWHHUP 528 100-1050
.6164 1.898
.1020
.0996
.2150
.0000 1.050
.2600 1
PFWHBCDN 518 100 1050
.6073 1.876
.1020
.0996
.2150
.0000 1.050
.2600 Y
PFWHBCDN 528 100 1050
.6164 1.897
.1020
.0996
.2150
.0000 1.050
.2600 C
SOPCRLCD 130
'60 1050
.2550 1.032
.0672
.0146
.1290
.0000 1.050
.1560 Y
SOPCRLHU 522 100 1050
.6109 1.885
.1020
.0996
.2150
.0000 1.050
.2600 7
EXCESSCD 100 50 0
.2282
.9663
.0585
.0571
.1075
.0000
.0000
.1300 C EXCESSHU 552 50 1050
.6381 1.950
.0585
.0571
.1075
.0000 1.050
.1300 j
I FWTEMPRD 528 100 1050
.6164 1.898
.1020
.0996-
.2150
.0000 1.050
.2600 SHUTDMSV 130 60 1050
.2550 1.032
.0672
.0146
.1290
.0000 1.050
.1560 o
SRVBLDWN 100 100 0
.2282
.9663
.1020
.0996
.2150
.0000 1.050
.2600 HPRC-INJ 300 7
875
.4096 1.402
.0211-
.0206
.0151
.0000
.8750
.0182 w
DESHYDRO 100 7
1250
.2282
.9663
.0211
.0206
.0151
.0000 1.250
.0182 U
i ZEROLOAD 0
0 0
'O.
O.
O.
O.
O.
O.
O.
O.
E O
?
NOTES:
i t
4 a
N l
t n.
--...a
4 i
i Appendix B GE-NE-523-61-0593 DRF 137-0010-6 TABLE B-4-5 THERMAL STRESS SCALING FACTORS AT REFERENCE AT SCALING TRANSIENT I.D.
'F
'F FACTORS COMMENTS CYCLES i
i DESHYDRO O
O O.0 130 1.O A,J 30 LOFP-CD 1.0 B
30 LOFP-HU START-UP 100/hr 360/hr 0.50 C,K 216 1.0 60000 NORMOP1 NORMOP2 1.0 D
10 SHUTDOWN 100/hr 4560/hr 0.50 E,K 216 SCRAMCDN 128 228 0.56 A,E 180 SCRAMHUP 152 200 O.76 B,C 180 LOFWHCDN 32 228 0.14 A,E 10 LOFWHHUP 32 200 0.16 B,C 10 PFWHBCDN 10 228 0.04 A,E 70 PFWHBHUP 10 200 0.05 B,C 70 SRVBLDWN 100/hr 4560/hr 0.50 E,K 40 1.O A
4O SOPCRLCD 1.0 B
40 SOPCRLHU EXCESSCD 160/hr 100/hr 1.6 A,F 10 EXCESSHU 160/hr 100/hr 1.6 B,G 10 FWTEMPRD 20 392 0.05 H
2000 SHUTDNSV 130 392 0.33 H
216 1.0 E
20 HPRC-INJ 0.0 I
390 ZEROLOAD COMMENTS:
A - CD or CDN stands for cool down E - scaled from LOFP-CD B - HU or HUP stands for heat up F - scaled from SHUTDOWN C - scaled from LOFP-HU G - scaled from START-UP D - NORMOP1 plus seismic loads H - scaled from SOPCRLCD I - DESHYDRO+ SHUTDOWN +SRVBLDWN=130&216+40=390 cycles J - consist of LOFP-CD0 at 0% flow and LOFP-CD7 at 7% flow K - 50% is used when stresses are porportional to ramp rate B-24
- _.1
Appendix B G E-NE-52 3-61-0 5 9 3 DRF 137-0010-6 4.2.2.4 Selected Locations for Stress Evaluation The sections identified in Figure B-4-5 are the locations selected for stress evaluation. These sections were chosen based on a review of stress distributions due to thermal transients, pressure loading, and interaction loading looking for the higher stressed areas.
This review was performed in a previous nozzle analysis. Of these locations, sections 3,4,6,and 7 were selected for stress evaluation because their P+Q stress range and usage factors were higher than the other locations accounting for the nozzle, safe end, and cladding material.
4.2.2.5 Resolution of Membrane, Bendino, and Peak Stresses FEA stresses were resolved into membrane, bending, and peak stresses across locations selected for stress evaluation. The section membrane stress was determined by averaging the FEA stresses across the section. The bending stress was determined by first subtracting the membrane stress from the section stress distribution, then finding moments associated with this modified stress distribution.
The moments were found assuming linear variation in stress between elements. The equivalent linear bending stress for the section was found using the beam approximation according to the following relation:
6(IM )
g b"
2 t
where, {&b} is the equivalent bending stress at the surface; {t} is the finite element section thickness; and {Mi} equals the mesment associated with each finite element stress after the membrane stress is removed. The peak stress at the surface was calculated as the difference between the finite element surface stress and the membrane plus bending stress. This technique is illustrated in Figure B-4-6.
To perform linearization, a PC computer program was used that required the following inputs (1) the number of points evaluated at the section - number of elements plus two, (2) the radius to each point - inner and outer surfaces plus element centroids, and (3) the value of component stress at each point along the section.
B-25
~
l Appendix B GE-NE-523-61-0593 DRF 137-0010-6 I
f
/,
N<
y
//
b b$ 4g/fl,'
\\W
\\
x N/
1
\\
r i
s l
O@g x
o l
l FIGURE B-4-5 PRIMARY PLUS SECONDARY STRESS ANALYSIS AREAS OF INTEREST l
B-26
{
i i
I l
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 W_
w f
l l
PEAK STRESS
/
BENDING STRESS MEMBRANE STRESS
/2 l
C SECTION THICKNESS
'v i
FINITE ELEMENT STRESS l
DISTRIBUTION I
- - - EQUIVALENT LINEAR STRESS
- - - - EQUIVALENT MEMBRANE STRESS FIGURE B-4-6 RESOLUTION OF STRESS DISTRIBUTION ACROSS j
A SECTION B-27 l
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 4.3 Faticue Analysis The cumulative fatigue analysis was performed to satisfy the requirements of Subsection NB-3222.4(e) of the ASME Code at the selected locations of the FEM accounting for each of the thermal cycles from Table B-4-5.
The fatique usage was calculated from the FEA results at sections 3,4,6, and 7; these sections are the more critical locations based on the results of [B-2).
A cumulative fatigue analysis versus a lump sum fatigue analysis was performed to remove inherent conservatism due to the lump sum analysis method.
Primary plus secondary and peak thermal stress ranges were calculated at each section.
Similarly, mechanical, pressure, and hydraulic stress ranges were calculated using the stresses determined in Section i
4.2.2 of this appendix. Respective mechanical, pressure, hydraulic, and thermal stress ranges were combined by absolute sum to yield total primary plus secondary stress ranges and total peak stress ranges. The peak l
stresses were determined in Section 4.2.2.5 of this appendix.
Alternating stresses were calculated according to the following equation:
Sa = b*
EE (P+Q)K
+F 2
E t
where, (Sa} is the alternating stress intensity; {P+Q) is the total primary plus secondary stress intensity range; {F) is the peak stress intensity range; {Kt} is the stress concentration factor; (Ke} is the simplified elastic plastic factor; and {Ec/E) is the Young's modulus correction factor. The values of {Kt} are discussed in Section 4.3.1. of this j
1 appendix. The values of {Ke) were determined in accordance with Subsection NB-3222.4(e) of the ASME Code. The values of (Ec/E) were all 1.0 because I
j the same value of Young's modulus used to generate the appropriate fatigue curve from Appendix I of the ASME Code was used in the FEA.
l I
B-28 l
I
Appendix B GE-NE-523-61-0593 DRF 137-0010-6 4.3.1 Stress concentration Factors The stress concentration factors were developed to account for local structural discontinuities not totally accounted for using FEA.
In Sections 4.3.1 and 5.5.2 of [B-2), effective stress concentration values were determined for the nozzle at the sections considered not modeled in enough detail to accurately determine local stresses due to discontinuities. These effective stress concentration factors, which are provided in Table B-4-6, were used to magnify the P+Q stress ranges used in the fatigue evaluation. The effective values were calculated by comparing theoretical stress concentrations with FEA stress concentrations. The FEA stress concentrations were based on mechanical load cases of axial, moment, and shear. The values recorded in Table B-4-6 are an average of the three cases. The effective values of {KTeff} were used as values for {Kt) in the equation found in Section 4.3 of this appendix.
TABLE B-4-6 EFFECTIVE STRESS CONCENTRATION FACTORS USED IN FATIGUE ANALYSIS I
( }
Sec..on K,f 2-I 1.28 Safe End Loads 3-0 1.35 Safe End Loads 4-I 2.41 Safe End Loads 7-0 1.3S Thermal Sleeve Loads NOTES:
(1) I = inside surface, O = outside surface (2) Average of three cases (3) K for section 6-I is conservatively assumed as 4.0 B-29
I o
Appendix B GE-NE-523-61-0593 DRF 137-0010-6
}
4.3.2 Cumulative Usace Factor In accordance with Subsection NB-3222.4(e)(5) of the ASHE Code, j
the cumulative damage method was used to obtain cumulative usage factors j
i i
that must not exceed 1.0.
1 Events were paired and ranked based on the magnitude of the alternating stress range (determined as described in Section 4.3).
The maximum number of available cycles common to both events in a given range were assigned and remaining cycles were applied when appropriate in the order of the highest to lowest alternating stress range. The allowable number of cycles for a given stress range was determined by using fatigue curves in Figure I-9.2.1 of [B-5) for stainless steel and Figure I-9.1 of (B-5) for low alloy steel.
The actual number of cycles at a given alternating stress range divided by the allowable number of cycles at a given alternating stress range defines a usage factor.
5.
RESULTS The results of the stress and fatigue analysis performed in the scope of this appendix are as follows:
5.1 Thermal Transient Analysis The critical times "t"
for the sudden Start of Pump in Cold Recirculation Loop (SOPCRL) transient are listed below.
(1) SOPCRL Cool Down t=
.65 minute (2) SOPCRL Heat Up t = 1.16 minute i
l In addition, section 7 was evaluated at critical times of t=.281 minute for cool down and t=.733 minute for heat up.
B-30 1
i
4 1
1 Appendix B GE-NE-523-61-0593 DRF 137-0010-6 i
4 j
5.2 Primary Plus Secondary Stress j
Primary plus secondary stress ranges were calculated for sections j
3,4,6, and 7 of the nozzle and safe end identified in Figure B-4-5.
Total 1
l thermal stress intensities were calculated for cool down and heat up of the f
SOPCRL transient.
Total thermal stress intensity ranges, and pressure and mechanical stress intensity ranges were added to find the total P+Q stress
\\
ranges. When the total P+Q ranges exceeded 3*Sm, simplified elastic 4
plastic analysis was performed in accordance with Subsection NB-3228.3 of I
e l
the Code.
Limiting primary plus secondary stress range for each section can be found in Table B-5-1.
P+Q stresses (unit stresses for mechanical l
and pressure loads) for given loads (mechanical,. pressure, or thermal) at I;
each section of interest are provided in Tables B-5-2 through 8-5-9.
Additional P+Q analysis information can be found in Tables B-5-13-1,2 and 4
)
B-5-17-1,2 and B-5-18-1,2 for sections 4-I, 7-I, and 7-0 respectively.
"I" 4
stands for inside surface and "O" stands for outside surface.
1 TABLE B-5-1 TOTAL P+Q STRESS RANGES 4
j P+Q P+Q-Qb j
SECTION SURFACE LIMITING THERMAL CASES fosil fosi)
ALLOWABLE j '
3 I
NORMOP2 - SHUTDNSV 33335 51800 I
51800 i
3 O
SCRAMHUP - ZEROLOAD 45201 b
l 4
I NORMOP2 - ZEROLOAD 54318 53097 55200 l
51800 j
i 4
0 NORMOP2 - ZEROLOAD 26219 50100 6
I NORMOP2 - ZEROLOAD 42072 q
1 80100 I
6 O
NORMOP2 - ZEROLOAD S1913 4
i j
7 I
LOFP-CD0 - SCRAMHUP 68457 40194 51800 7
O LOFP-CD0 - NORMOP2 68402 39869
- 51800 J
1 4
a 23A4274 Rev. 2 has P+Q-Qb equal to 46100 which also meets allowable b 3*Sm was based on an average temperature (400+550)/2 due to the specific range of operating condition.
i B-31 4
j i
i
(
3
TABLE B-5-2 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 3-I l
MECH (or)
PRIMARY STRESSES P + U _ STRESSES P+Q+F STRESSES THERMAL hADL TANG HOOP RADL TANG HOOP RADL TANG HOOP SHEAR-SE 0.
D.
-2350.
-235.
O.
-2374.
-175.
O.
4748.
573.
O.
4728.
545.
MOMNT-SE 0.
O.
-6932.
2922.
O.
-7163.
2898.
SHEAR-TS 0.
O.
-15418.
5853.
O.
-15887.
5806.
MOMNT-TS 0.
AXIAL-TS 0.
O.
2059.
-1034.
O.
2126.
-1032.
TORSN-TS 0.
O.
O.
O.
O.
O.
O.
-1000.
2000.
6438.
-1000.
2074.
5390.
PRESSURE
-1000.
-1000.
3168.
5407.
-1000.
3285.
5454.
ADDRPRES
-1000.
O.
13967.
12384.
O.
16309.
15214.
LOFP-CD0 0.
O.
-3255.
-2406.
O.
-3712.
-2968.
LOFP-NUO 0.
O.
-1628.
-1203.
O.
-1856.
-1484.
START-UP 0.
O.
22.
153.
O.
18.
148.
NORMOP1 0.
NORMOP2 0.
O.
22.
153.
O.
18.
148.
O.
6984.
6192.
O.
8155.
7607.
SHUTDOWN O.
? SCRAMCDN 0.
O.
7822.
6935.
O.
9133.
8520.
O.
-2474.
-1829.
O.
-2821.
-2256.
M SCRAMHUP 0.
O.
1955.
1734.
O.
2283.
2130.
LOFWHCDN 0.
l LOFWHHUP 0.
D.
-521.
-385.
O.
-594.
-475.
O.
559.
495.
O.
652.
609.
PFWHBCDN 0.
SOPCRLCD 0.
2620.
7113.
O.
58736.
61841.
O.
72344.
77266.
SOPCRLHU 0.
-820.
-431.
O.
-12971.
-17279.
O.
-19071.
-24339.
O.
11174.
9907.
O.
13048.
12171.
EXCESSCD 0.
O.
-2605.
-1925.
O.
-2970.
-2374.
EXCESSHU 0.
FWTEMPRD 0.
131.
356.
O.
2937.
3092.
O.
3617.
3863.
l SHUTDNSV 0.
865.
2347.
O.
19383.
20408.
O.
23874.
25498.
O.
6984.
6192.
O.
8155.
7607.
SRVBLDNN 0.
ZEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
- Mechanical load and pressure stresses are unit stresses 4
L 1
TABLE B-5-3 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 3-0 MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP O.
-12529.
-2705.
O.
-12911.
-2676.
SHEAR-SE 0.
O.
14352.
2669.
O.
14714.
2670.
MOMNT-SE 0.
O.
6248.
6164.
O.
6403.
6131.
SHEAR-TS 0.
O.
13565.
12989.
O.
13934.
12925.
MOMNT-TS 0.
O.
-1967.
-1823.
O.
-2019.
-1815.
AXIAL-TS 0.
TORSN-TS 0.
O.
-26.
26.
O.
-24.
24.
O.
3528.
5772.
O.
3602.
5811.
PRESSURE 0.
O.
2413.
4373.
O.
2463.
4417.
ADDRPRES 0.
O.
-12974.
-10597.
O.
-10843.
-7588.
LOFP-CD7 0.
LOFP-NUO 0.
O.
3041.
2440.
O.
2674.
1863.
START-UP O.
O.
1521.
1220.
O.
1337.
932.
O.
-24.
287.
O.
-35.
279.
NORMOP1 0.
O.
-24.
287.
O.
-35.
279.
MORMOP2 0.
SHUTDOWN 0.
O.
-6487.
-5299.
O.
-5422.
-3794.
? SCRAMCDN 0.
D.
-7265.
-5934.
O.
-6072.
-4249.
D.
2311.
1854.
O.
2032.
1416.
$ SCRAMHUP 0.
O.
-1816.
-1484.
O.
-1518.
-1062.
LOFWHCDN 0.
O.
487.
390.
O.
428.
298.
LOFWHHUP 0.
PFWHBCDN 0.
O.
-519.
-424.
O.
-434.
-304.
SOPCRLCD 0.
2620.
7113.
O.
-53496.
-47615.
O.
-40236.
-31560.
SOPCRLHU 0.
-820.
-431.
O.
11331.
16847.
O.
4948.
9894.
EXCESSCD 0.
O.
-10379.
-8478.
O.
-8675.
-6070.
EXCESSHU 0.
O.
2434.
1952.
O.
2139.
1491.
FWTEMPRD 0.
131.
356.
O.
-2675.
-2381.
O.
-2012.
-1578.
SHUTDMSV 0.
865.
2347.
O.
-17654.
-15713.
O.
-13278.
-10415.
O.
-6487.
-5299.
O.
-5422.
-3794.
SRVBLDWN 0.
EEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
NOTE: Mechanical load and pressure stresses are unit stresses 1
i m...
d TABLE B-5-4 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 4 - I h
MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP O.
-10077.
-838.
O.
-16458.
-2133.
SHEAR-SE 0.
MOMNT-SE 0.
O.
9085.
748.
O.
14476.
1840.
O.
11573.
3203.
O.
28864.
6143.
SHEAR-TS D.
O.
37082.
9686.
O.
90653.
18606.
MOMNT-TS 0.
O.
-5535.
-1381.
O.
-13111.
-2650.
AXIAL-TS 0.
O.
-4389.
4368.
O.
10644.
-10643.
TORSN-TS 0.
I PRESSURE
-1000.
-1000.
5006.
5452.
-1000.
10581.
6411.
O.
-2451.
1699.
O.
-8888.
941.
ADDRPRES-O.
LOFP-CD0 0.
595.
-2909.
O.
11412.
8789.
O.
39509.
16687.
LOFP-NUO 0.
-218.
1663.
O.
-4103.
-1484.
O.
-11096.
-3537.
START-UP 0.
-109.
832.
O.
-2052.
-742.
O.
-5548.
-1769.
NORMOP1 0.
-77.
980.
O.
-1298.
630.
O.
-1478.
558.
NORMOP2 0.
-77.
980.
O.
-1298.
630.
O.
-1478.
558.
SHUTDOWN 0.
298.
-1455.
D.
5706.
4395.
O.
19755.
8344.
y SCRAMCDN 0.
333.
-1629.
O.
6391.
4922.
O.
22125.
9345.
y SCRAMHUP 0.
-166.
1264.
O.
-3118.
-1128.
O.
-8433.
-2688.
LOFWHCDN 0.
83.
-407.
O.
1598.
1230.
O.
5531.
2336.
LOFWHHUP 0.
-35.
266.
O.
-656.
-237.
O.
-1775.
-566.
PFWH8CDN 0.
24.
-116.
O.
456.
351.
O.
1580.
667.
SOPCRLCD 0.
743.
-13614.
O.
15549.
-7159.
O.
29666.
-5066.
SOPCRLHU 0.
-223.
59.
O.
-7529.
706.
O.
-11142.
-405.
EXCESSCD 0.
477.
-2328.
O.
9129.
7032.
O.
31608.
13350.
EXCESSHU 0.
-174.
2661.
O.
-3283.
-1187.
O.
-8877.
-2830.
FWTEMPRD 0.
37.
-681.
O.
777.
-358.
O.
1483.
-253.
SHUTDMSV 0.
245.
-4493.
O.
5131.
-2362.
O.
9790.
-1672.
SRVBLDWN 0.
298.
-1455.
O.
5706.
4395.
O.
19755.
8344.
EEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
NOTE:
Mechanical load and pressure stresses are unit stresses i
4 h
I i
m
w
.-. - -. - -. ~.... ~.
-. - ~.. ~.. -,
- ~
- ~. -
TABLE B-5-5 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 4 - 0 MECH (or)
PRIMARY STRESSES P +Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP SHEAR-SE 0.
O.
-4145.
1832.
O.
-4455.
1553.
O.
4311.
-1123.
O.
4523.
-942.
MOMNT-SE 0.
O.
-372.
-467.
O.
1093.
31.
SHEAR-TS 0.
O.
-7770.
-6360.
O.
-2896.
-4404.
MOMNT-TS 0.
I O.
1108.
922.
O.
415.
637.
AXIAL-TS 0.
O.
-1930.
1944.
O.
-2302.
2303.
TORSN-TS 0.
O.
-641.
2677.
O.
-158.
2919.
PRESSURE 0.
O.
2355.
2977.
O.
1765.
2767.
ADDRPRES 0.
LOFP-CD0 0.
O.
-10229.
-14578.
O.
-5520.
-10921.
O.
3670.
4794.
O.
2430.
3888.
LOFP-NUD 0.
D.
1835.
2397.
O.
1215.
1944.
START-UP 0.
O.
1143.
1320.
O.
978.
1271.
NORKOP1 0.
NORMOP2 0.
O.
1143.
1320.
O.
978.
1271.
O.
-5115.
-7289.
O.
-2760.
-5461.
SHUTDOWN 0.
y SCRAMCDN 0.
O.
-5728.
-8164.
O.
-3091.
-6116.
O.
2789.
3643.
O.
1847.
2955.
y SCRAMHUP 0.
O.
-1432.
-2041.
O.
-773.
-1529.
LOFWHCDN 0.
LOFWHHUP D.
O.
587.
767.
O.
389.
622.
PFWHBCDN 0.
O.
-409.
-583.
O.
-221.
-437.
SOPCRLCD 0.
743.
-13614.
O.
-14064.
-20068.
O.
-7355.
-17265.
SOPCRLHU 0.
-223.
59.
O.
7083.
-589.
O.
7250.
121.
EXCESSCD 0.
O.
-8184.
-11662.
O.
-4416.
-8738.
O.
2936.
3835.
O.
1944.
3110.
EXCESSHU 0.
FWTEMPRD 0.
37.
-681.
O.
-703.
-1003.
D.
-368.
-863.
SHUTDMSV 0.
245.
-4493.
O.
-4641.
-6622.
O.
-2427.
40.
O.
-5115.
-7289.
O.
-2760.
-5461.
SRV8LDWN 0.
ZEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
NOTE: Mechanical load and pressure stresses are unit stresses l
I 1
TA8LE B-5-6 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 6 - I MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP O.
-9734.
-2793.
O.
-9824.
-2485.
SHEAR-SE 0.
O.
5777.
1239.
O.
5813.
1089.
MOMNT-SE 0.
O.
822.
-589.
O.
798.
-553.
SHEAR-TS 0.
O.
10906.
2190.
O.
10973.
1908.
MOMNT-TS 0.
O.
-2097.
-437.
O.
-2101.
-386.
AXIAL-TS 0.
O.
-3113.
3113.
O.
-3015.
3015.
TORSN-TS 0.
PRESSURE 0.
-1000.
638.
3552.
-1000.
672.
3365.
ADDRPRES 0.
D.
151.
-67.
O.
155.
-64.
0.'
318.
-4849.
O.
-11697.
-16066.
LOFP-CD0 0.
O.
-13610.
-12490.
O.
-29609.
-27397.
i LOFP-NUO 0.
START-UP 0.
O.
-6805.
-6245.
O.
-14805.
-13699.
l O.
-10520.
-11521.
O.
-28706.
-28512.
NORMOP1 0.
O.
-10520.
-11521.
O.
-28706.
-28512.
NORMOP2 0.
SHUTDOWN 0.
O.
159.
-2425.
O.
-5849.
-8033.
7 SCRAMCDN 0.
O.
178.
-2715.
O.
-6550.
-8997.
O.
-10344.
-9492.
O.
-22503.
-20822.
y SCRAMHUP 0.
LOFWHCDN 0.
O.
45.
-679.
O.
-1638.
-2249.
O.
-2178.
-1998.
O.
-4737.
-4384.
LOFWHHUP 0.
O.
13.
-194.
O.
-468.
-643.
PFWH8CDN 0.
1 SOPCRLCD 0.
-924.
1262.
O.
-12663.
-12621.
O.
-30350.
-27733.
SOPCRLHU 0.
-882.
809.
O.
-11301.
-12740.
O.
-28945.
-28050.
EXCESSCD 0.
D.
254.
-3880.
O.
-9358.
-12853.
O.
-10888.
-9992.
O.
-23688.
-21918.
EXCESSHU O.
FWTEMPRD 0.
-46.
63.
O.
-633.
-631.
O.
--1518.
-1387.
SHUTDNSV 0.
-305.
416.
O.
-4179.
-4165.
O.
-10016.
-9152.
O.
159.
-2425.
O.
-5849.
~8033.
SRV8LDWN 0.
EEROLOAD 0.
O.
O.
O.
O.
.O.
O.
O.
O.
t NOTE: Mechanical load and pressure stresses are unit stresses i
i t
~.......... ~.
~. -.
TABLE B-5-7 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 6 - O MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL EM HOOP RADL TANG HOOP SHEAR-SE 0.
O.
-13561.
-3307.
O.
-13239.
-3108.
O.
7569.
1472.
O.
7407.
1372.
MOMNT-SE 0.
O.
601.
-593.
O.
626.
-576.
SHEAR-TS 0.
O.
15770.
2998.
O.
15455.
2805.
MOMNT-TS 0.
AXIAL-TS 0.
O.
-2283.
-439.
O.
-2225.
-406.
O.
-3714.
3714.
O.
-3660.
3660.
TORSN-TS 0.
O.
3707.
3925.
O.
3683.
3800.
PRESSURE 0.
O.
-130.
-126.
O.
-125.
-122.
ADDRPRES 0.
O.
-905.
6696.
O.
-7454.
575.
LOFP-CD0 0.
LOFP-NUO O.
O.
10822.
13942.
O.
2545.
5985.
START-UP 0.
O.
- 5411, 6971.
O.
1273.
2993.
NORMOP1 0.
O.
8031.
14850.
O.
-1563.
5736.
O.
8031.
14850.
O.
-1563.
5736.
NORMOP2 0.
SHUTDOWN 0.
O.
-453.
3348.
O.
-3727.
288.
y SCRAMCDN 0.
O.
-507.
3750.
O.
-4174.
322.
y SCRAMHUP 0.
O.
8225.
10596.
O.
1934.
4549.
O.
-127.
937.
O.
-1044.
81.
LOFWHCDN 0.
LOFWHHUP 0.
O.
1732.
2231.
O.
407.
958.
PFWHBCDN 0.
O.
-36.
268.
O.
-298.
23.
SOPCRLCD 0.
-924.
1262.
O.
10815.
15144.
O.
461.
5008.
SOPCRLHU 0.
-882.
809.
O.
9538.
14358.
O.
-788.
4146.
EXCESSCD 0.
O.
-725.
5357.
O.
-5963.
461.
D.
8658.
11154.
O.
2037.
4789.
EXCESSHU 0.
FWTEMPRD 0.
-46.
63.
O.
541.
757.
O.
23.
250.
SHUTDNSV 0.
-305.
416.
O.
3569.
4998.
O.
152.
1653.
SRVBLDWN 0.
O.
-453, 3348.
O.
-3727.
288.
EEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
NOTE: Mechanical load and pressure stresses are unit stresses I
~
TABLE B-5-8 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 7 - I (SOPCRL-CD, t=.281 min; SOPCRL-HU, t=.733 min)
MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP O.
-1592.
979.
O.
-1271.
1085.
SHEAR-SE 0.
MOMNT-SE 0.
O.
4112.
-685.
O.
3433.
-943.
O.
10768.
-1957.
O.
5977.
-4152.
SHEAR-TS 0.
O.
-2599.
-7100.
D.
-10767.
-11193.
MOMNT-TS 0.
O.
705.
1185.
O.
1743.
1711.
AXIAL-TS 0.
O.
10882.
-10937.
O.
13379.
-13378.
TORSN-TS 0.
-1000.
-4150.
2854.
-1000.
-3844.
2940.
PRESSURE
-1000.
-1000.
12758.
8257.
-1000.
11194.
8039.
ADDRPRES
-1000.
LOFP-CD0 0.
-610.
27376.
O.
-32223.
20188.
O.
-27111.
22443.
LOFP-NUO 0.
107.
-5532.
O.
6894.
-3795.
O.
5735.
-4375.
START-UP 0.
54.
-2766.
O.
3447.
-1898.
O.
2868.
-2188.
l NORMOP1 0.
-10.
903.
O.
-1226.
449.
O.
-1008.
535.
NORMOP2 0.
-10.
903.
O.
-1226.
449.
O.
-1008.
535.
4 7 SHUTDOWN 0.
-305.
13688.
O.
-16112.
10094.
O.
-13556.
11222.
y SCRAMCDN 0.
-342.
15331.
O.
-18045.
11305.
O.
-15182.
12568.
SCRAMHUP 0.
81.
-4204.
O.
5239.
-2884.
O.
4359.
-3324.
j LOFWHCDN 0.
-85.
3833.
O.
-4511.
2826.
O.
-3796.
3142.
LOFWHHUP 0.
17.
-885.
O.
1103.
-607.
O.
918.
-700.
PFWHBCDM 0.
-24.
1095.
O.
-1289.
808.
O.
-1084.
898.
SOPCRLCD 0.
447.
37034.
O.
14761.
64309.
O.
21353.
71288.
SOPCRLHU 0.
-1821.
38190.
O.
-73454.
-4288.
O.
-96775.
-27251.
EXCESSCD 0.
-488.
21901.
O.
-25779.
16150.
O.
-21690.
17955.
EXCESSHU O.
86.
-4426.
O.
5515.
-3037.
O.
'4589.
-3501.
FWTEMPRD 0.
22.
1852.
O.
738.
3215.
O.
1068.
3564.
t SHUTDNSV 0.
148.
12221.
O.
4871.
21222.
O.
7046.
23525.
SRVBLDWN 0.
-305.
13688.
O.
-16112.
10094.
O.
-13556.
11222.
EEROLOAD 0.
O.
O.
O.
D.
O.
O.
O.
O.
NOTE:
Mechanical load and pressure stresses are unit stresses i
....=
TABLE B-5-9 PRIMARY, SECONDARY, AND PEAK STRESSES FOR SECTION 7 - 0 (SOPCRL-CD, t=.281 min; SOPCRL-HU, t=.733 min)
MECH (or)
PRIMARY STRESSES P+Q STRESSES P+Q+F STRESSES THERMAL RADL TANG HOOP RADL TANG HOOP RADL TANG HOOP SHEAR-SE 0.
O.
1408.
2054.
O.
2675.
2142.
MOMNT-SE 0.
O.
-3838.
-3331.
O.
-6332.
-3515.
SHEAR-TS 0.
O.
-63712.
-25431.
O.
-90541.
-28705.
MOMNT-TS 0.
O.
-134070.
-49201.
O. -187519.
-56583.
AXIAL-TS 0.
O.
17325.
6447.
O.
24259.
7409.
TORSN-TS 0.
O.
22147.
-22444.
O.
28503.
-28503.
PRESSURE
-1000.
-1000.
1514.
4239.
-1000.
2965.
4235.
ADDRPRES 0.
O.
2140.
4S10.
O.
424.
4949.
LOFP-CDD 0.
-610.
27376.
O.
30313.
34127.
O.
47986.
35664.
LOFP-NUO 0.
107.
-5532.
O.
-6533.
-7171.
O.
-10248.
-7529.
START-UP 0.
54.
-2766.
O.
-3267.
-3586.
O.
-5124.
-3765.
NORMOP1 0.
-10.
903.
D.
1180.
1335.
O.
1863.
1375.
NORMOP2 0.
-10.
903.
O.
1180.
1335.
O.
1863.
1375.
7 SHUTDOWN 0.
-305.
13688.
O.
15157.
17064.
O.
23993.
17832.
$ SCRAMCDN 0.
-342.
15331.
O.
16975.
19111.
O.
26872.
19972.
SCRAMHUP 0.
81.
-4204.
O.
-4965.
-5450.
O.
-7788.
-5722.
LOFWHCDN 0.
-85.
3833.
O.
4244.
4778.
O.
6718.
4993.
LOFWHHUP 0.
17.
-885.
O.
-1045.
-1147.
O.
-1640.
-1205.
PFWHBCDN 0.
-24.
1095.
O.
1213.
1365.
O.
1919.
1427.
SOPCRLCD 0.
447.
37034.
O.
-13867.
9758.
O.
-8029.
16936.
SOPCRLHU 0.
-1821.
38190.
O.
69813.
80667.
O.
53489.
63176.
EXCESSCD 0,
-488.
21901.
O.
24251.
27302.
O.
38389.
28531.
EXCESSHU 0.
86.
-4426.
O.
-5227.
-5738.
O.
-8198.
-6024.
FWTEMPRD 0.
22.
1852.
O.
-693.
488.
O.
401.
847.
SHUTDNSV 0.
148.
12221.
O.
-4576.
3220.
O.
2650.
5589.
SRVBLDWN 0.
-305.
13688.
O.
15157.
17064.
O.
23993.
17832.
ZEROLOAD 0.
O.
O.
O.
O.
O.
O.
O.
O.
NOTE: Mechanical load and pressure stresses are unit stresses
....~
i
)
l Appendix B GE-NE-523-61-0593 DRF 137-0010-6 5.3 Faticue Evaluation The results of the cumulative fatigue usage factors for sections 3,4,6, and 7 are presented in this section.
The fatigue usage factors were calculated in the order of the limiting combinations of the events listed in Table B-4-4, then summed to give the total forty year life usage factor.
Table B-5-10 is a summary of the results of the fatigue evaluation.
Tables B-5-11 through B-5-18 provide more detail of the analysis.
r TABLE B-5-10 FATIGUE USAGE FACTORS SECTION SUFFACE USAGE ALLOWABLE 3
1 0.086 1.0 3
0 0.100 1.0 4
I 0.095 1.0 4
0 0.000 1.0 6
I 0.107 1,0 6
0 0.002 1.0 7
I 0.225 1.0 7
0 0.511 1.0 I
B-40
-.n TABLE B-5-11 FATIGUE
SUMMARY
F0R SECTION 3-I SM = 17250.
3SM = 51750.
E = 28300000.
KT =
1.000 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 3 SOPCRLCD-SOPCRLHU 22437.
81469.
71576.
9893.
2.914 151405.
40 467.
.086 1 NORMOP2 -SHUTDNSV 4585.
33335.
1.000 18960.
g y
.086 ti
=
TABLE B-5-12 FATIGUE
SUMMARY
F0R SECTION 3-0 R,
SM = 17250.
3SM = 51750.
E = 28300000.
KT =
1.350 7
STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 NORMOP2 -SOPCRLCD 20345.
93023.
-56116.
42195.
3.333 188947.
10 241.
.042 1 SOPCRLCD-SOPCRLHU 10356.
84238.
68267.
22851.
3.093 146271.
30 519.
.058 o
1 SOPCRLHU-EEROLOAD 13507.
53700.
-12151.
43189.
1.126 37825.
10 150698.
.000 4
1.000 30924.
180 509373.
.000 1 SCRAMHUP-EEROLOAD 16648.
45201.
p" 1.000 30797.
330 525658.
.001 1 FWTEMPRD-EEROLOAD 16293.
45300.
1 LOFP-NUO-SHUTDNSV 10027.
40688.
1.000 25357.
S o
.100
?.
t a
4
-m
TABLE B-5-13 FATIGUE
SUMMARY
F0R SECTION 4-I SM = 17250.
3SM = 51750."
E = 28300000.
KT =
2.410 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 NORMOP2 -ZEROLOAD 119655.
54318.
1221.
53097.
1.165 101376.
10 1722.
.006 1 SOPCRLHU-ZEROLOAD 102583.
46238.
1.000 74410.
40 5445.
.007 1 SCRAMHUP-ZEROLOAD 99469.
42451.
1.000 70960.
180 6594.
.027 E
i 1 SCRAMCDN-EEROLOAD 98760.
38082.
1.000 68421.
180 7638.
.024 S
1 LOFP-CDO-EEROLOAD 99348.
36667.
1.000 68008.
50 7827.
.006 s
1 EXCESSHU-EEROLOAD 91420.
39863.
1.000 65642.
10 9029.
.001 1 NORMOP1 -EEROLOAD 90714.
40180.
1.000 65447.
50 9138.
.005 1 NORMOP1 -EXCESSCD 92829.
34523.
1.000 63676.
10 10250.
.001 0
Y 1 LOFP-NUO-SOPCRLCD 87776.
37010.
1.000 62393.
30 11317.
.003 2 NORMOP1 -SOPCRLCD 82959.
40395.
1.000 61677.
10 11970.
.001 h
1 NORMOP1 -SHUTDOWN 79573.
31100.
1.000 55336.
216 20301.
.011 1 START-UP-SRVBLDWN 70600.
26238.
1.Ord 48419.
40 39876.
.001 u
E 7 1 NORMOP1 -SHUTDNSV 56497.
24644.
1 000 40571.
216 103101.
.002 1.000 22059.
{
$ 1 NORMOP1 -LOFWHCDN 32605.
11512.
g
.095 w
- 3*Sm based on an average temperature of 475'F for the E
m specific range of operating condition is 55200.
U, 8
5 m
l 4
4
r i
TABLE B-5-13-1 P+Q INYENSTIY RANGES F0R SECTION 4-I RANGE STRESS STRESS THERMAL P+Q MECHANICAL P+0 MECH-ADDER P+0 NO STATE 1 STATE 2 RL L-H HR R-L L-H H-R R-L L-H H-R o
E 1 NORMOP2 ZER0 LOAD 1298.
-1928.
630.
50600. -40445.
12038.
2420.
1792.
628 E
2 NORMOP2 SOPCRLCD 16847. -24636.
7789.
29933.
26871.
3478.
2420.
1792.
628.
E 3 NORMOP2 EXCESSCO 10427.
-4025.
-6402.
35814.
27241.
10014.
2420.
1792.
628.
4 LOFP-CD0 NORMOP2
-12710.
4551.
8159.
27765.
24056.
4811.
2420.
1792.
628 5 NORMOP2 SHUTDOWN 7004.
-3239.
-3765.
35814.
- 27241, 10014.
2420.
1792.
628.
n 6 SOPCRLMU ZEROLDAD 7529.
-8235.
706.
36289.
27382.
10727.
2420.
1792.
628.
7 7 SOPCRLHU EXCESSCD 16658. -10332.
-6326.
21502.
14178.
8703.
2420.
1792.
628.
8 SCRAMHUP ZER0 LOAD 3118.
-1990.
-1128.
36913.
27575.
11225.
2420.
1792.
628.
i 9 SOPCRLCD SOPCRLHU
-23078.
30943.
-7865.
'15622.
13808.
2167.
2420.
1792.
628.
E 10 SCRAMMUP EXCESSCD 12247.
-4087.
-8160.
22126.
14370.
9201.
2420.
1792.
628 Y
11 SCRAMCDN ZEROLDAD
-6391.
1469.
4922.
29272.
21451.
- 9551, 2420.
1792.
628 7
12 LOFP-CD0 ZER0 LOAD
-11412.
2623.
8789.
22835.
16389.
7227.
2420.
1792.
628.
Y 13 LOFP-CD0 SOPCRLHU
-18941.
10858.
8083.
13454.
10994.
3500.
2420.
1792.
628.
O 14 EXCESSHU ZER0 LOAD 3283.
-2096.
-1187.
34160.
25661.
9878.
- 2420, 1792.
628 15 NORMOP1 ZER0 LOAD 1298.
-1928.
630.
36462.
27541.
In741.
2420.
1792.
628.
16 LOFP-NUO ZEROLDAD 4103.
-2619.
-1484.
32937.
25425.
8734.
2420.
1792.
628.
17 LOFWHHUP ZER0 LOAD 656.
-419.
-237.
36462.
27541.
10741.
2420.
1792.
628.
5 18 FWTEMPRD ZER0 LOAD
-777.
1135.
-358.
36462.
27541.
10741.
2420.
1792.
628
~
19 PFWH8CDN ZER0 LOAD
-456.
105.
351.
36171.
27274.
10717.-
2420.
1792.
628.
20 SHUTDOWN SOPCRLHU
-13235.
9546.
3689.
21502.
14178.
0703.
2420.
1792.
628.
2 21 EXCESSCD EXCESSHU
-12412.
4193.
8219.
.19374.
12456.
7854.
2420.
1792.
628 O
^
~ 22708.
-7159, 20667.
13574.
8560.
2420.
1792.
628.
5 22 50PCRLCD ZEROLDAD
-15549.
23 SCRAMMUP SOPCRLCD 18667. -24698.
6031.
16246.
14001.
2665.
2420.
1792.
628.
24 NORMOP1 EXCESSCD 10427.
-4025.
-6402.
21676.
14337.
8717.
2420.
1792.
628.
25 LOFP-NUO EXCESSCD-13232.
-4716.
-8516.
18150.
12220.
6711.
2420.
1792.
628 26 LOFWHHUP EXCESSCD 9785.
-2516.
-7269.
21676.
14337.
8717.
2420.
1792.
628 27 LOFP-CD0 SCRAMHUP
-14530.
4613.
9917.
14078.
11186.
3998.
2420.
1792.
628.
28 START-UP ZER0 LOAD 2052.
-1310.
-742.
34160.
25661.
9878.
2420.
1792.
628.
29 LOFP NUO SOPCRLCD 1%52. -25327.
5675.
14938.
12547.
2636.
2420.
1792.
628 30 NORMOP2 SHUTDNSV 6429.
-9421.
2992.
29933.
26871.
3478.
2420.
1792.
628.
i
~
-- - - +
w-a---+
u- --
--r
. i -
..m
_m..
_.__.m__
.-_-m..__.___._m._
m_
TABLE B-5-13-2 PEAK INiENSIiY RANGES F0R SECiION 4-I RANGE STRESS STRESS THERMAL PEAK MECHANICAL PEAK MECH-ADDER PEAK NO STATE 1 STATE 2 RL LH
.H-R R-L L-H H-R R-L L-H H-R
- p e
E a
1 NORMOP2 ZEROLDAD 180.
-108.
-72.
39580.
32346.
7450.
3459.
2881.
578.
E 2 NORMOP2 SOPCRLCD 14297. -12132.
-2165.
22610.
- 18450, 4376.
3459.
2881.
5 78.
3 NORMOP2 EXCESSCD 22659
-16269.
-6390.
27193.
22226 5183.
3459.
2881.
5 78.
4 LOFP-CD0 NORMOP2
-28277.
20307.
7970.
23488.
19329 4375.
3459.
2881.
578.
o
~
5 NORMOP2 SHUIDOWN 14229. -10208.
-4021.
27193.
22226.
5183.
3459.
2881.
578.
?
6 SOPCRLMU ZER0 LOAD 3613.
-2502.
-1111.
30316.
24801.
5515.
3459.
2881.
578.
Y 7 SOPCRLHU EXCESSCD 26092. -18663.
-7429.
17929.
14681.
3248.
3459.
2881.
5 78.
8 SCRAMMUP ZEROLDAD 5315.
-3755.
-1560.
30839.
25231.
5608.
- 3459, 2881.
578.
E 9 SOPCRLCD SOPCRLHU
-17730.
14526.
3204.
13346.
10905.
2441.
3459.
2881.
5 78.
Y*
10 SCRAMMUP EXCESSCD 27794. -19916.
-7878.
18452.
15111.
3341.
3459.
2881.
578.
Y 11 SCRANCDN ZER0 LOAD
-15734 11311.
4423.
25870.
21239.
4631.
3459.
2881.
5 78.
12 LOFP-CD0 ZER0 LOAD
-28097.
20199.
7898.
16091..
13017.
3075.
3459.
2881.
578.
8 13 LOFP-CD0 SOPCRLHU
-31710.
22701.
9009.
14224.
11784 2440.
3459.
2881.
578.
8 14 EXCESSHU ZEROLDAD 5594.
-3951.
-1643.
26160.
21281.
4879.
3459.
2881.
578.
15 NORM 0P1 ZER0LDAD 180.
-108.
- 72.
30421.
24885.
5536.
- 3459, 2881.
5 78.
ts 16 LOFP-NU0 ZEROLDAD 6993.
-4940.
-2053.
25079.
20389.
4690.
3459.
2881.
578.
5 17 LOFWHHUP ZEROLDAD 1119.
-790.
-329.
30421.
24885.
5536.
3459.
2881.
578.
~
18 FUTEMPRO ZEROLDAD
-706.
601.
-105.
30421.
24885.
5536.
3459.
2881.
578.
19 PFWH8 CON ZER0 LOAD
-1124.
808.
316.
30244.
24744.
5500.
3459.
2881.
578.
20 SHUTDOWN SOPCRLHu
-17662.
12602.
5060.
17929.
14681.
3248.
3459.
2881.
578.
8 21 EXCESSCD EXCESSHU
-28073.
20112.
7961.
13772.
11160.
2612.
3459.
2881.
578.
22 SOPCRLCD 2EROLDAD
-14117.
12024.
2093.
16970.
13896.
3074.
3459.
2881.
578.
23 SCRAMMUP SOPCRLCD 19432. -15779.
-3653.
13869.
11335.
-2534.
3459.
2881.
578.
l 24 NORMOP1 ENCESSCD 22659. -16269.
-6390.
18034.
14765.
3269.
3459.
2881.
578.
25 LOFP-NUO EXCESSCD 29472. -21101.
-8371.
12692.
10269.
2423.
3459.
2881.
578.
26 LOFWHHUP EXCESSCD 23598. -16951.
-6647.
18034.
14765.
3269.
3459.
2881.
578.
27 LOFP-CD0 SCRAM 4UP
-33412.
23954.
9458.
14748.
12215.
2533.
3459.
2881.
-578.
28 START-UP ZEROLDAD 34%.
-2469.
-1027.
26160.
21281.
4879.
- 3459, 2881.
578.
+
i l
i
+
TABLE B-5-14 FATIGUE
SUMMARY
F0R SECTION 4-O SM = 17250.
3SM = 51750.
E= 28300000.
KT =
1.000 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANCE RANCE BENDING RANCE KE INTENSITY CYCLES CYCLES FACTOR 2 NORMOP2 -ZEROLOAD 239.
26219.
1.000 13229.
.000 m
3o O.
M
=
0 t4 J.
O I*
7 O
W u
t3 I
u 7
O O
p H
i
.m_
~~m._..
m.
_ _. _ _ _ _... _ _ _ _ _ _... _. _ _ _ _. _.. _ - ~
f TABLE B-5-15 FATIGUE
SUMMARY
F0R SECTION 6-I SM = 17250.
3SM = 51750.
E = 28300000.
KT =
4.000 STRESS STRESS PEAK P4Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 NORMOP2 -EEROLCAD 144649.
42072.
1.000 93361.
10 2309.
.004 1 LOFP-NUO-EEROLOAD 118226.
34021.
1.000 76124.
30 4970.
.006 1 SOPCRLHU-EEROLOAD 114504.
32231.
1.000 73367.
40 5764.
.007 1 NORMOP1 -EEROLOAD 113090.
31578.
1.000 72334.
440 6103.
.072 3
S 1 PFWHBCDN-SOPCRLCD 85815.
22846.
1.000 54331.
40 22273.
.002 1 NORMOP1 -SHUTDOWN 81342.
23018.
1.000 52180.
216 27320.
.008 1.000 50568.
10 32017.
.000 1 NORMOP1 -EXCESSCD 78023.
23113.
1 EXCESSHU-SRVBLDWN 73403.
22178.
1.000 47790.
10 42600.
.000 g
1.000 47232.
30 45207.
.001 1 NORMOP1 -SRVELDWN 73910.
20554.
f 1.000 39664.
216 116523.
.002 1 NORMOP1 -SHUTDNSV 62602.
16727.
1 LOFP-CD7-NORNOP1 60522.
18100.
1.000 39311.
50 122306.
.000 g
1 NORMOP1 -SCRAMCDN 60862.
16454.
1.000 38658.
180 133936.
.001 w
1 NORMOP1 -LOFWHCDN 56342.
13275.
1.000 34808.
10 242528.
.000 g
1.000 31779.
30 428211.
.000 g
?
1 NORMOP1 -PFWHBCDN 52096.
11462.
1 NORMOP1 -FWTEMPRD 49105.
10601.
1.000 29853.
2000 666066.
.003 g
1 NORMOP1 -LOFWHHUP 42796.
9056.
1.000 25926.
w
.107 q
t:
TABLE B-5-16 FATIGUE
SUMMARY
F0R SECTION 6-O o
SM = 26700.
3SM = 80100.
E = 30000000.
KT =
1.000 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR i
l 1 NORMOP2 -EEROLOAD
-7380.
51913.
1.000 22266.
10 58722.
.000 1 SCRAMHUP-EEROLOAD
-6886.
38212.
1.000 15663.
180 253679.
.001 1 EXCESSHU-EEROLOAD
-7210.
38096.
1.000 15443.
10 270616.
.000 1 NORMOP1 -EEROLOAD
-7061.
37740.
1.000 15339.
320 279053.
.001 1 LOFP-NUO-EXCESSCD
-3367.
28558.
1.000 12596.
10 933730.
.000 1 LOFP-NUO-SHUTDOWN
-5331.
28286.
1.000 11478.
.002
TABLE B-5-17 FATIGUE
SUMMARY
F0R SECTION 7-I SM = 17250.
3SM = 51750.
E = 28300000.
KT =
1.000 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 SOPCRLCD-SOPCRLHU 30039.
94781.
-85947.
8834.
3.333 208034.
40 183.
.219
>f 2 LOFP-CDO-SCRAMHUP
-4898.
68457.
-28263.
40194.
2.076 65980.
50 8844.
.006 2 SCRAMHUP-EXCESSCD
-4231.
65279.
23378.
41901.
1.871 57123.
10 17383.
.001 s
i S
2 LOFP-NUO-SHUTDOWN
-2969.
49218.
1.000 23124.
u
.225 o
M a
0 a
'T
=
k 8
~
Sa t1 t:
7 8
5 os 1
3 l
.m
~-
r rw rv-
~,
TABLE B-5-17-1 P+0 iNiENSTIY RANGES F0R SECiION 7-1 I
RANGE STRESS STRESS THERMAL P+0 MECHANICAL P+0 MECH-ADDER P+Q No STATE 1 STATE 2 R-L LH M-R R-L L-H H-R R-L L-H H-R e....
............... eeeeeee........
- e
.. ee******
eeeeeeen........................
1 SOPCRLCD SOPCRLMU
-88215.
1%18.
68597.
6161.
63t2.
2669.
405.
498.
472.
2 SOPCRLHU ZEROLOAD 73454. -69166.
-4288.
17117.
20987.
9504.
405.
498.
472.
3 SOPCRLMU SHUTDNSV 78325.
52815. -25510.
6161.
6342.
2669.
405.
498.
472.
4 LOFP-NUO SOPCRLMU
-80348.
79855.
493.
3243.
2913.
2449.
405.
498.
4 72.
5 SOPCRLHU EXCESSHU 78969. -77718.
-1251.
2754.
1763.
1789.
405.
498.
472.
6 START-UP SOPCRLMU
-76907 7;511.
2390.
2754.
1763.
1789.
405.
498.
472.
a.
7 SCRAMMUP SOPCRLMU
-786th Ti?89.
1404.
298.
602.
303.
405.
498.
472.
E 8 NORMOP2 SOPCRLHU
-72228.
e491.
4737.
4845.
6105.
1333.
405.
498.
472.
9 LOFWHNUP SOPCRLMU
-74557.
70876.
3681.
62.
77.
14.
405.
498.
472.
10 SOPCRLHU FWTEMPRD 74192. --66689.
-7503.
62.
77.
14, 405.
498.
472.
o 11 LOFWMCDN SOPCRLMU
-68943.
61829.
7114.
4572.
2766.
3085.
405.
498.
4 72.
7 12 NORMOP1 SOPCRLMU
-72228.
67491.
4737.
62.
77.
14.
405.
498.
472.
E 13 PFWHBCDN SOPCRLHO
-72165.
67069.
5096.
43.
52.
10.
405.
498.
472.
14 SHUIDOWN SOPCRLHU
-57342.
42960.
14382.
10138.
14127.
6768 405.
498.
472.
u 15 SOPCRLCD ZER0 LOAD
-14761. -49548.
64309.
10956.
14645.
6835.
405.
498.
472.
Y 16 LOFP-NUO SOPCRLCD 7867.
60237. -68104.
5451.
6%7.
2249.
405.
498.
472.
m 17 SCRAMHUP SOPCRLCD 9522.
57671. -67193.
6460.
6943.
2972.
405.
498.
472.
8 18 SOPCRLCD EXCES$NU
-9246.
-58100.
67346.
4%2.
5818.
1589.
405.
498.
472.
O 19 NORMOP2 SOPCRLCD 15987.
47873. -63860.
11006.
12446.
4002.
405.
498.
4 72.
20 START-UP SOPCRLCD 11314.
54893. -66207.
4%2.
5818.
1589.
405.
498.
472.
21 LOFWHHUP SOPCRLCD 13658.
51258. -64916.
6223.
6418.
2683.
405.
498.
472.
U 22 NORMOP1 SOPCRLCD 15987.
47873. -63860.
6223.
6418.
2683.
405.
498.
472.
5 23 SOPCRLMU SRV9tDWN 57342. -42960. -14382.
4387.
5391.
1004.
405.
498.
472.
e 24 PFWHSCDN SOPCRLCD 16050.
47451. -63501.
6118.
6289.
2659.
405.
498.
472.
25 LOFP-CD0 SCRAMMUP 37462. -60534.
23072.
8043.
7425.
4941.
405.
498.
4 72.
26 LOFP-CD0 LOFP-NUO 39117. -63100.
23983.
4998.
4522.
2303.
405.
498.
472.
O 27 SOPCRLCD FWiEMPRD
-14023. -47071.
61094.
6223.
6418.
2683.
405.
498.
472.
5 28 LOFWHCDN SOPCRLCD 19272.
42211. -61483.
6069.
5862.
2302.
405.
498.
472.
29 LDFP-CD0 EXCESSHU 37738. -60963.
23225.
5612.
5823.
2991.
405.
498.
472.
30 LOFP-CD0 ZER0 LOAD 32223. -52411.
20188.
9373.
14164.
4867.
405.
498.
472.
31 SCRAMCDN SOPCRLMU
-55409.
39816.
15593.
2622.
3535.
913.
405.
498.
472.
32 SOPCRLHU EXCESSCD 47675. -27237. -20438.
10138.
14127.
6768.
405.
498.
472.
33 SCRAMMUP EXCESSCD
-31018.
50052. -19034 10436.
14728.
7072.
405.
498.
472.
4 34 LOFP-NUO EXCESSCD
-32673.
52618. -19945.
7392.
11825.
4433.
405.
498.
472.
35 EXCESSCO EXCESSHU 31294. -50481.
19187.
8005.
13127.
5122.
405.
498.
472.
36 LOFP-CD0 START-UP 35670. -57756.
22086.
5612.
5823.
2991.
405.
498.
472.
i 37 LOFP-CD0 NORMOP2 30997. -50736.
19739.
12589.
12928.
5970.
'405.
498.
472.
38 SHUTDOWN SOPCRLCD 30873.
23342. -54215.
4197.
8168.
4703.
405.
498.
4 72.
39 LOFP-CD0 LOFWHHUP 33326. -54121.
20795.
7806.
6900.
4652.
405.
498.
4 72.
40 START-UP EXCESSCD
-29226.
47274. -18048.
8005.
13127.
5122.
405.
498.
472.
41 NORMOP2 EXCESSCD
-24553.
40254
-15701.
14983.
20231.
8101.
405.
498.
4 72.
-- _, ~. _ - - -. - - -. - -. _.. _ -.
, - - +
+e-
--,---e-.-
--~----w-
-w
TABLE B-5-17-2 PEAK INTENSITY RANGES F0R SECTION 7-I RANGE STRESS STRESS THERMAL PEAK MECHANICAL PEAK MECH-ADDER PEAK NO STATE 1 STATE 2 R-L L-H H-R RL L-H H-R R-L LH H-R 1
t 1 SOPCRLCD SOPCRLMU
-29913.
-29.
29942.
1806.
1056.
750.
557.
281.
276.
2 SOPCRLHU 2EROLDAD 23321.
-358.
-22%3.
3729.
2287.
1443.
557.
281.
276.
'O 3 SOPCRLHU SHUTDNSV 254 %.
-230.
-25266.
1806.
1056.
750.
557.
281.
276.
l 4 LOFP NUO SOPCRLHU
-22162.
-221.
22383.
966.
581.
385.
557.
281.
276.
5 SOPCRLHU EXCESSHU 22395.
104. -22499.
923.
549.
374.
557.
281.
276.
5 6 START-UP SOPCRLMU
-22742.
69.
22673.
923.
- 549, 374.
557.
281.
276.
7 7 SCRAMMUP SOPCRLHU
-22441.
-82.
22523.
34.
23, 10.
557.
281.
276.
8 NORM 0P2 SOPCRLHU
-23539.
490.
23049.
- 953, 608.
380.
557.
281.
276.
9 LOFWHHUP SOPCRLHU
-23136.
266.
.22870.
11.
7.
4.
557.
281.
276.
o 10 SOPCRLHU FWTEMPRD 23651.
-339.
-23312.
11, 7.
4.
557.
281.
276.
7 11 LOFWNCDN SOPCRLHU
-24036.
757.
23279.
1578.
938.
641.
557.
281.
276.
g 12 NORM 0P1 SOPCRLHU
-23539.
490.
23049.
11.
7.
4.
557 281.
276.
i 13 PFWHBCDN SOPCRLHU
-23526.
4 73.
23053.
7.
5.
3.
557.
281.
276.
E 14 SHUTDOWN SOPCRLHU
-25877.
1786..
24091.
1938.
1216.
722.
557.
281.
276.
'7 15 SOPCRLCD 2ER0 LOAD
-6592.
-387.
6979.
1923.
1230.
693.
557.
281.
276.
i 16 LOFP-NU0 SOPCRLCD 7751.
-192.
-7559.
1241.
742.
499.
557.
281.
276.
17 SCRAMMUP SOPCRLCD 7472.
-53.
-7419.
1840.
1080.
760.
557.
281.
276.
8 18 SOPCRLCD EXCESSHU
-7518.
75.
7443.
1198.
711.
487.
557.
281.
276.
19 NORMOP2 SOPCRLCD 6374.
519.
-6893.
2759.
1665.
1130.
557.
281.
276.
20 START-UP SOPCRLCD 7171.
98.
-7269.
1198.
'.711.
487.
557.
281.
276.
ts 21 LOFWHHUP SOPCRLCD 6777.
295.
-7072.
1817.
1063.
754.
557.
281.
276.
4 22 NORMOP1 SOPCRLCD 6374.
519.
6893.
1817.
1063.
754.
557.
281.
276.
g 23 SOPCRLMU SRV8LDWN 25877.
-1786. -24091.
747.
469.
278.
557.
281.
276.
u 24 PFWH8CDN SOPCRLCD 6387.
502.
-6889.
1799.
1052.
747.
557.
281.
276.
Y 25 LOFP-CD0 SCRAMMUP
-5992.
- 3297, 2695.
2221.
1341.
880.
557.
281.
276.
8 26 LOFP-CD0 LOFP NUO
-6271.
- 3436, 2835.
1305.
789.
516.
557.
281.
276.
g 27 SOPCRLCD FW1EMPRO
-6262.
-368.
6630.
1817.
1063.
754.
557.
281.
276.
8*
28 LOFWHCDM SOPCRLCD 5877.
-786.
-6663.
.1233.
777.
456.
557.
281.
276.
29 LOFP-CD0 EXCESSHU
-6038.
3319.
2719.
1369.
834.
535.
557.
281.
276.
30 LOFP-CD0 2EROLDAD
-5112.
2857.
2255.
1542.
M9.
5 73.
- 557, 281.
276.
31 SCRAMCDN SOPCRLHU
-26184.
1958.
24226.
423.
269.
155.
557.
281.
276.
32 SOPCRLHU EXCESSCD 27410.
-2642. -24768.
1938.
1216.
722.
557.
281.
276.
33 SCRAMMUP EXCESSCD 4%9.
-2724.
-2245.
1972.
1239.
733.
- 557, 281.
276.
34 LOFP-NUO EXCESSCD 5248.
-2863.
-2385.
1057.
688.
'369.
557.
281.
276.
35 EXCESSCD EXCESSHU
-5015.
2746.
2269.
1121.
733.
388.
- 557, 281.
276.
36 LOFP-CD0 START-UP
-5691.
3146.
2545.
1369.
834.
535.
557.
281.
276.
37 LOFP-CD0 NORMOP2
-4894.
2725.
2169.
3140.
1926.
1249.
557.
281.
276.
38 SHUTDOWN SOPCRLCD 4036.
1815.
-5851.
826.
506.
320.
557.
281.
276.
39 LOFP-CD0 LOFWHHUP
-5297.
2949.
2348.
2198.
1324.
874.
557.
281.
276.
J
TABLE B-5-18 FATIGUE
SUMMARY
FOR SECTION 7-O SM = 17250.
3SM = 51750.
E = 28300000.
KT =
1.350 STRESS STRESS PEAK P+Q THERMAL P+Q-QT ALTERNATING DESIGN ALLOWABLE USAGE STATE 1 STATE 2 RANGE RANGE BENDING RANGE KE INTENSITY CYCLES CYCLES FACTOR 1 SOPCRLHU-EEROLOAD 43098.
113561.
-71634.
45569.
3.333 261098.
40 100.
.400 y
1 LOFP-CDO-NORMOP2 58024.
68402.
-29733.
39869.
2.073 131015.
10 737.
.014 9
r 1 LOFP-CDO-SCRAMHUP 57722.
68202.
-35969.
33615.
2.060 129682.
40 761.
.053 1 SCRAMCDN-EEROLOAD 48385.
58334.
-17317.
41701.
1.424 75990.
180 5003.
.036 1 SCRAMHUP-EXCESSCD 48422.
55657.
29785.
27010.
1.252 65136.
10 9315.
.001 M
1 SRVBLDWN-EEROLOAD 44929.
54840.
-15462.
39988.
1.199 59814.
40 13896.
.003 5
1 SCRAMHUP-EEROLOAD 39782.
48959.
1.000 44371.
130 63137.
.002 a
1.000 40564.
130 103190.
.001 y
1 NORMOP1 -EEROLOAD 36143.
44986.
1 SHUTDOWN-LOFWHHUP 36162.
42454.
1.000 39308.
80 122360.
.001 1.000 38226.
136 142336.
.001 1 SHUTDOWN-FWTEMPRD 34350.
42102.
2 NORMOP1 -SOPCRLCD
- 26143, 39441.
1.000 32792.
40 352028.
.000 0
1.000 28210.
a 1 LOFWHCDN-FWTEMPRD 23617.
32803.
y
[n
.511 O
o to t3 C
Y 8
\\
I l
I i
TABLE B-5-18-1 P+Q lNiENSTIY RANGES F0R SECi10N 7-0 RANGE STRESS STRESS THERMAL P+Q MECHANICAL P+Q MECH-ADDER P+0 NO STATE 1 STATE 2 R-L L-H H-R R-L L-H H-R R-L L-H H-R e....
...**e...............................................................
i 1 SOPCRLHU ZER0 LOAD
-69813. -10854.
80667.
34868.
19524.
23088.
8880.
5570.
3310.
2 SOPCRLCD SOPCRLMU 83680. -12771. -70909.
19101.
10390.
9663.
8880.
5570.
3310.
3 SOPCRLHU SHUTDNSV
-74389.
-3058.
77447.
19101.
10390.
9663.
8880.
5570.
3310.
'o 4 LOFP-NUO SOPCRLHO 76346.
11492. -87838.
11280 7267.
5611.
8880.
5570.
3310.
- E 3 SOPCRLHU EXCESSHU
-75040. -11365.
86405.
10898.
6800.
4749.
8880.
5570.
3310.
5.
6 START-UP SOPCRLMU 73080.
11173. -84253.
10898.
6800.
4749.
8880.
5570.
3310.
p 7 LOFWHCDN SOPCRLMU 65569.
10320. -75889.
18928.
11925.
8118.
8880.
5570.
3310.
8 NORMOP2 SOPCRLMU 68633.
10699. -79332.
6592.
2454.
5129.
8880.
5570.
3310.
9 SCRAMMUP SOPCRLMU 74778.
11339. -86117.
246.
215.
448.
8880.
5570.
3310.
o 10 SHUTDOWN SOPCRLHU 54656.
8947. -63603.
17315.
10324.
13824 8880.
5570.
3310.
7 11 LOFWHNUP SOPCRLHU 70858.
10956. -81814.
58.
10.
55.
8880.
5570.
3310.
12 LOFP-CD0 NORMOP2
-29133.
-3659.
32792.
30389.
16805.
17022.
8880.
5570.
3310.
7' 13 SOPCRLHU FWTEMPRD
-70506.
-%73.
80179.
58.
10.
55.
8880.
5570.
3310.
C 14 LOFP-CD0 SCRAMMUP
-35278.
-4299.
39577.
24043.
14566.
12340.
8880.
5570.
3310.
y to 15 NORMOP1 SOPCRLMU 68633.
10699. -79332.
58.
10.
55.
8880.
5570.
3310.
m En 16 PFWH8CDN SOPCRLHU 68600.
10702. -79302.
40.
7.
37.
8880.
5570.
3310.
Y 17 LOFP-CD0 LOFWHNUP
-31358.
-3916.
35274.
23855.
14361.
11947.
8880.
5570.
3310.
S 18 LOFP-CD0 FWTEMPRD
-31006.
-2633.
33639.
23855.
14361.
11947.
8880.
5570.
3310.
g 19 LOFP-CD0 NORMOP1
-29133.
-3659.
32792.
23855.
14361.
11947.
8880.
5570.
3310.
20 LOFP-CD0 PFWHBCDN
-29100.
-3662.
32762.
23758.
14344.
11855.
8880.
5570.
3310.
o 21 SOPCRLMU EXCESSCD
-45562.
-7803.
53365.
17315.
10324.
13824.
8880.
5570.
3310.
4 22 LOFP-CD0 SOPCRLHU 39500.
7040. -46540.
23797.
14351.
11892.
8880.
5570.
3310 y
23 LOFP CD0 SOPCRLCD
-44180.
19811.
24369.
7972.
4536.
5330.
8880.
5570.
3310.
w 24 LOFP-CD0 LOFP-NUO
-36846.
-4452.
41298.
12978.
7165.
6717.
8880.
5570.
3310.
Y 25 SCRANCDN ZER0 LOAD
-16975.
-2136.
19111.
32479.
18877.
20542.
8880.
5570.
3310.
8 26 LOFP-CD0 EXCESSHU
-35540.
-4325.
39865.
13475.
7652.
7688.
8880.
5570.
3310.
g 27 SOPCRLHU SRVSLDWN
-54656.
-8947.
63603.
4065.
713.
3846.
8880.
5570.
3310 i*
28 LOFP-CD0 START-UP
-33580.
-4133.
- 37713, 13475.
7652.
7688.
8880.
5570.
3310 29 NORMOP2 EXCESSCD 23071.
2896. -25%7.
23906.
12778.
18953.
8880.
5570.
3310 30 SCRAMMUP EXCESSCD 29216.
3536. -32752.
17561.
10538.
14271.
8880.
5570.
3310 31 SCRAMCDN SOPCRLCD
-30842.
21489.
9353.
17215.
10288.
8164.
8880.
5570.
3310.
32 SOPCRLCD EXCESSCD 38118. -20574. -17544.
%10.
7147.
8343.
8880.
5570.
3310.
33 SRV8LDWN ZER0 LOAD
-15157.
-1907.
17064.
30803.
18811.
19242.
8880.
5570.
3310.
w
TABLE B-5-18-2 PEAK INTENSITY RANGES F0R SECTION 7-0 RANGE STRESS STRESS THERMAL PEAK MECHANICAL PEAK MECH-ADDER PEAK NO STATE 1 STATE 2 RL L*H H-R R-L LH H-R R-L L-H H-R E
1 SOPCRLHU ZEROLOAD 16324.
1167. -17491.
16996.
15438.
1795.
3572.
3087.
485.
]
2 SOPCRLCD SOPCRLHU
-22162.
-2507.
24669.
8830.
7863.
1058.
3572.
3087.
485.
o 3 SOPCRLHU SHUTDNSW 23550.
-3690. -19860.
8830.
7863.
1058.
3572.
3087.
485.
S 4 LOFP-NUO SOPCRLHU
-12609.
-4524.
17133.
4818.
4305.
629.
3572.
3087.
485.
as 5 SOPCRLHU EXCESSHU 13353.
3852. -17205.
4604.
4087.
631.
3572.
3087.
485.
6 START-UP SDPCRLHU
-14467.
-2845.
17312.
- 4604, 4087.
631.
3572.
3087.
485.
7 LOFWHCDN SOPCRLHU
-18798.
1092.
17706.
7930.
7032.
1098.
3572.
3087.
485.
O 8 NORMOP2 SOPCRLHU
-17007.
-524.
17531.
3885.
35%.
376.
3572.
3087.
485.
E 9 SCRAMHUP SOPCRLHU
-13501.
-3718.
17219.
148.
146.
3.
3572.
3087.
485.
7 10 SHUTDOWN SOPCRLHU
-25160.
6901.
18259.
8707.
8006.
824.
3572.
3087.
485.
=
11 LOFWHHUP SOPCRLHU
-15729.
-1704.
17433.
39.
37.
3.
3572.
3087.
485.
U T
12 LOFP-CD0 NORMOP2
-16990.
15493.
1497.
14413.
13013.
1717.
3572.
3087.
485.
A
+'
m 13 SOPCRLHU FWiEMPRD 17418.
432. -17850.
39.
37.
3.
3572.
3087.
485.
5 14 LOFP-CD0 SCRAMMUP
-204%.
18687.
1809.
10676.
9563.
1344.
3572.
3087.
485.
15 NORMOP1 SOPCRLHU
-17007.
-524.
17531.
39.
37.
3.
3572.
3087.
485.
16 PFWHBCDN SOPCRLHU
-17030.
-523.
17553.
27.
25.
2.
3572.
3087.
485.
17 LOFP-CD0 LOFWHHUP
-18268.
16673.
1595.
10567.
9453.
1343.
3572.
3087.
485.
o 18 LOFP-CD0 FUTEMPRD
-16579.
15401.
1178.
10567.
9453.
1343.
3572.
3087.
485.
- o' 19 LOFP-CD0 NORMOP1
-16990.
15493.
1497.
10567.
9453.
1343.
3572.
3087.
485.
20 LOFP-CD0 PFWHBCDN
-16%7.
15492.
1475.
10501.
9392.
- 1339, 3572.
3087.
485.
U 21 SOPCRLHU EXCESSCD 30462. -11742. -18720.
8707.
8006.
'824.
3572.
3087.
485.
y 22 LOFP-CD0 SOPCRLHU
-33997.
14 % 9.
19028.
10528.
9417.
1341.
3572.
3087.
485.
o 23 LOFP-CD0 SOPCRLCD
-11835.
17476.
-5641.
3935.
3628.
- 446, 3572.
3087.
485.
8 24 LOFP-CD0 LOFP-NUG
-21388.
19493.
1895.
6024.
5404.
735.
3572.
3087.
485.
?*
25 SCRAMCDN ZER0 LOAD
-9897.
9036.
861.
15391.
13939.
1688.
3572.
3087.
485.
26 LOFP-CD0 EXCESSHU
-20644.
18821.
1823.
6317.
5695.
738.
3572.
3087.
485.
27 SOPCRLHU SRV9LDWN 25160.
-6901.
-18259.
2776.
2573.
203.
3572.
3087.
485.
28 LOFP-CD0 START-UP
-19530.
17814.
1716.
6317.
5695.
738.
3572.
3087.
485.
29 NORMOP2 EXCESSCD 13455. -12266.
-1189.
12592.
11602.
1200.
3572.
3087.
485.
30 SCRAMHUP EXCESSCD 16961. -15460.
-1501.
8855.
8151.
827.
3572.
3087.
485.
31 SCRAMCDN SOPCRLCD
-4059.
10376.
-6317.
7515.
6656.
952.
3572.
3087.
485.
f Appendix B GE-NE-523-61-0593 DRF 137-0010-6 6.
REFERENCES B-1.
" Peach Bottom Atomic Power Station, Units 2 & 3 Reactor Vessel Thermal Cycle Roanalysis", Memorandum Reference No. 1-20-2820-213, January, 1992 B-2.
Recirculation Inlet Nozzle and Safe End Certified Stress Report, Peach Bottom 2, 23A4274 Revision 2, General Electric Company B-3.
Reactor Vessel - Recirculation Inlet Safe End Design Specification, Peach Bottom 2, 23A4158, Revision 3, General Electric Company.
B-4.
Reactor Vessel - Recirculation Inlet Safe End Design Specification, Peach Bottom 3, 23A5719, Revision 0, General Electric Company.
B-5.
ASME Boiler and Pressure Vessel Code,Section III, Nuclear Power Plant Components; 1989 Edition with Addenda to and including Winter 1990, American Society of Mechanical Engineers, New York, NY.
B-6.
Recirculation Inlet Nozzle and Safe End Certified Stress Report, Peach Bottom 3, DC23A5720 Revision 0, General Electric Company B-7.
G.J. DeSalvo and J.A. Swanson, "ANSYS User's Manual", 2nd Edition, March 1975, Swanson Analysis Systems Inc., Houston, PA.
B-8.
"BWR Plant Materials Properties Handbook", January, 1982, General Electric Company B-9.
Certified Design Document for Peach Bottom II Reactor, Stress Report VPF No. 1896-142-1, Volumes 1-6, The Bobcock & Wilcox company, December, 1970.
B-10.
ASME Boiler and Pressure Vessel Code,Section III, Nuclear Power Plant Components; 1980 Edition with Addenda to and including Summer 1981. American Society of Mechanical Engineers, l
New York, NY.
l f
B-11.
Safe End Drawing 137C8434, Revision 0, General Electric Company l
(
l B-53 i
l 1
I i
APPENDIX C Peach Bottom II and III Evaluation of Power Rerate Including Modified Cycles i
l l
l
i l
l APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 l
l Table of Contents 1.0 Introduction C-2 2.0 Power Rerate Conditions C-2 2.1 Design Condition Changes C-2 l
2.2 Operating Condition Changes C-2 3.0 Power Rerate Stress Analysis for lionbolting Materials C-5 3.1 ASME Code Stress Analysis C-5 l
3.2 Design Changes C-6 3.3 Normal Conditions C-6 l
3.4 Emergency and Faulted Conditions C-11 1
4.0 Power Rerate Stress Analysis for Bolting Materials C-11 4.1 Design Conditions C-11 4.2 Normal, Upset, and Emergency Conditions C-12 4.3 Faulted Conditions C-14 5.0 Component Analysis C-14 5.1 Selection Criteria for Power Rerate Stress Analysis C-14 5.2 Closure Bolts C-16 l
5.3 Feedwater Nozzle C-20 5.4 Support Skirt C-27 5.5 Refueling Containment Skirt C-31 5.6 Recirculation Inlet Nozzle C-35 6.0 Conclusions C-40 7.0 References C-42 C-1
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6
1.0 INTRODUCTION
This appendix documents the ASME Boiler and Pressure Vessel Code,Section III, analysis of limiting reactor pressure vessel (RPV) components for the Peach Bottom Nuclear Power Plant Units 2 and 3 (Peach Bottom 2 and 3). This analysis is based on proposed increases j
in core thermal power levels (power rerate) which will change some original design basis operating parameters such as coolant pressures, temperatures and nozzle flow rates. These changes in pressures, temperatures and flows will, in general, increase the original stress values in the RPV components. In addition, the expected number of cycles was modified based on actual cycles expected at Peach Bottom 2 and 3. This analysis constitutes the stress report reconciliation for validating the use of existing RPV components for the power rerate conditions.
Unless otherwise stated, any section number stated refers to those in this appendix.
l 2.0 POWER RERATE CONTITIONS 2.1 Design Condition Changes As noted in Paragraph 4.3 of the Design Specification for the Peach Bottom Power Rerate program (Reference C-1), the power rerate design requirements are unchanged from the original design requirements specified in the reactor pressure vessel purchase documents (References C-2 and C-3).
2.2 Operating Condition Changes l
t As noted in Paragraph 4.4.1 of Reference C-1, the changes to the Reactor Cycles document i
l (Reference C-4) are as follows:
l In Region A, the operating pressure shall be increased from 1000 to 1038 psig (1015 e
to 1053 psia).
In Regions A, B and C, the operating temperature shall be increased from 546*F to 551*F when specified.
C-2
i
)
J n
}
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 i
In the " scram" transient events, increase the following identified Region A pressures e
by 38 psig (38 psia). Increase the corresponding temperatures as shown in Table j
j C-2-1:
~
Table C-2-1 POWER RERATE CHANGES FOR THE SCRAM TRANSIENT Original Power Rerate Original Power Rerate J
Pressure Pressure Temp.
Temp.
(psig/ psia)
(psig/ psia)
(*F)
(*F) 1180 / 1195 1218 / 1233 573 573 875 / 890 913 / 928 530 536 1125 / 1140 1163 / 1178 561 565 1000 / 1015 1038 / 1053 546 551 665 / 680 703 / 718 500 506 930 / 945 968 / 983 538 542 l
i In Regions B and C, the identified operating temperatures shall be increased as follows: 522 to 527*F, 538 to 542*F, 500 to 506*F, 546 to 551*F, and 512 to 517*F.
As noted in Paragraph 4.4.2 of Reference C-1, the changes to the Reactor Vessel Nozzle Thermal Cycles document (Reference C-5) are as follows:
On sheet 1 (Recirculation Outlet), the 100% rated flow per nozzle increases from e
43,500 to 49,700 gpm.
On sheet 2 (Recirculation Inlet), the 100% rated flow per nozzle increases from 8700 o
to 9940 gpm.
C-3
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 On sheet 3 (Steam Outlet), the 100% rated flow / nozzle increases from 3.35 x 106o t
e 6
3.81 x 10 lb/hr.
o On sheets 4 and 5 (Feedwater), the feedwater inlet temperature increases from 376 F to 387 F.
On sheets 4 and 5 (Feedwater), the 100% rated flow / nozzle increases from 5550 to o
5890 gpm.
On all sheets, the vessel bulk temperature, when specified, shall be increased from o
546*F to 551*F.
On all sheets, the vessel pressure shall be increased from 1000 psig to 1038 psig o
(1015 to 1053 psia) when specified.
As noted in Paragraph 4.4.3 of Reference C-1, the changes in the Reactor Vessel Purchased Part Drawing (Reference C-3) are as follows:
e On sheet 7, operating pipe reactions shall be maximum at 551*F in Note 7. The references to 546 in this note shall be revised to 551*F.
e On sheet 7, in the ""*" note in Table 4, revise "500* to 550*F" to "500 F to 551*F".
As noted in Paragraph 4.4.4 of Reference C-1, the change in the Reactor Vessel Purchase Specification (Reference C-2) is as follows:
In Paragraph 7.5.1, the operating shroud support pressure differentials of 34.0 and e
21.85 psid shall be increased to 35.68 and 26.32 psid.
l l
C-4 l
l
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 As noted in Paragraph 4.4.6 of Reference C-1, the changes in the supplemental design specifications are as follows:
In 23A5722, the 100% rated flow per nozzle increases from 43,500 gpm to 49,700 e
gpm.
j In 23A4158, the 100% rated flow per nozzle increases from 8700 gpm to 9940 gpm.
e In 23A5719, the 100% rated flow per nozzle increases from 8700 gpm to 9940 gpm.
e In 23A6656 (Reference C-6), the 100% rated flow per nozzle increases from 5100 e
gpm to 5890 gpm, and the reference in Figure 1 to 546*F shall be revised to 551*F.
In 23A6090, references to operating conditions of 1000 psig (1015 psia) and 546*F e
shall be revised to 1038 psig (1053 psia) and 551*F.
As noted in Paragraphs 4.4.5 and 4.4.7 of Reference C-1, all other operating requirements are unchanged from those specified in Reference C-2 and the supplemental specifications.
3.0 POWER RERATE STRESS ANALYSIS FOR NONBOLTING MATERIALS 3.1 ASME Code Stress Analysis The power rerate stress analysis uses the guidelines and procedures of the ASME Boiler and Pressure Vessel Code,Section III (Code). For the component under consideration, the 1965 Code (Reference C-7), which is the Code of construction, shall be the governing Code. However, if a component underwent a design modification, the governing Code shall be the Code used in the stress analysis of the modified component. Also, if the original analysis used a later version of the code, that version is used.
C-5
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 3.2 Design Conditions Since there are no changes in the design conditions due to power rerate, the design-based stresses (general primary membrane, primary membrane plus primary bending) remam 1
unchanged, and the Code requirements of Paragraphs N-414.1 through N-414.3 of Reference C-7 and Section NB-3221 of Reference C-8 are still met for all RPV components analyzed.
3.3 Normal and Upset Conditions 3.3.1 Effects of Channes in Pressure. Temperature and Nozzle Flow Rates In general, changes in normal operation pressures, temperatures and nozzle flow rates will increase the primary plus secondary (P+Q) stresses and the primary plus secondary plus peak (P+Q+F) stresses at a particular location on the RPV component.
The stress components [3 normal (a a. a6) and 3 shear (fr2, 79. Ter)] of the P+Q r
z stresses and the P+Q+F stresses consist of pressure stress com.ponents, thermal cycling stress components, and mechanical stress components due to seismic loads or reaction loads from attached piping and/or thermal sleeves.
The magnitude of the normal stress due to pressure is directly proportional to the coolant pressure, and the magnitude of the normal stress due to thermal cycling is proportional to the temperature change during a thermal transient (final transient temperature minus initial transient temperature).
In the case of nozzles, increases in coolant flow through the nozzle will increase the forced convection heat transfer coefficients on the inside (fluid side) surface of the nozzle. These increases in heat transfer coefficients will change the temperature distribution through the nozzle, thus changing the thermal stresses in the nozzle slightly. In general, small changes in the heat transfer coefficient on the nozzle inside surface have a negligible effect on the temperature distribution through the nozzle. However, for a large change in the heat transfer coefficient, the change in the normal stress is a function of the change in heat t
l transfer coefficients resulting from an increased flow rate (Reference C-9).
C-6
1 APPENh.' C GE-NE-523-61-0593 DRF 137-0010-6 3.3.2 Power Rerate Scaline Techniave A technique was developed to conservatively scale up the original stresses to account for pressure and temperature increases due to power rerate.
In many pressure vessel calculations, the shear stress components are zero and, thus, the principal stress directions coincide with the normal stress directions. Therefore, the magnitude of the princiV. 'ess due to pressure is directly proportional to the coolant pressure, and the magnitue e principal stress due to thermal cycling is proportional to the temperature change during a thermal transient. Since there are usually no changes in the mechanical stresses due to power rerate, the new (power rerate) value for principal stress is:
atotal, new
= apressure,old * (b._; sold)
+ athermal,old * (ATnew/ATold)
+ amechanical or:
atotal, new
= apressure,old
- s Ap
+ athermal,old * (SCF)T
+ amechanical where (SCF)p = pressure scaling factor = (Pnew/ old).
P (SCF)T = thermal scaling factor = (ATnew/ATold).
AT
= (final transient temperature) - (initial transient temperature)
If the flow rate increase resulting from the power rerate conditions is large, an additional scaling factor (SCF)p must be zpplied. The (SCF)p is required because an increase in the flow rate will increase the heat transfer and, therefore, the thermal stresses. For this case, C-7
.n e
i APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 1
the (SCF)p is determined as a function of heat transfer coefficients as described in Reference C-9. The overall thermal scaling factor becomes:
(SCF)T = (ATnew/ATold) * (SCF)p Most stress reports do not explicitly report values for pressure stresses, thermal cycling stresses and mechanical stresses separately. Therefore, it is not possible to calculate power i
rerate principal stresses by scaling the original pressure stress by (SCF)p and the original thermal cycling stress by (SCF)T and combining them with the original mechanical stress.
i Thus, a conservative scaling technique was developed where the original principal stress components are scaled up by the larger of (SCF)p and (SCF)T. This is conservative because (1) the larger scaling factor (SCF) is used to scale up the stresses, and (2) the j
mechanical stress is scaled up as well. In addition, if the scaling factor is less than unity, the power rerate principal stresses are not scaled down. In that case, a scale factor of 1.0 is used (i.e., the original values are retained).
To further simplify the scaling technique, the larger scaling factor (SCF) can be applied directly to the original stress intensity values instead of to the original principal stresses.
Stress intensity (or " stress difference") is determined by taking the algebraic difference between any pair of principal stresses. The following example illustrates why the scale factor can be applied to the stress intensities directly:
S12,new = al,new - a2,new
= (al,old
- SCF) - (a2,old
- SCF)
= (al,old - a2,old)
- SCF
=S12,old
- SCF.
3.3.3 ASME Code Stress Limits for Normal and Uoset Conditions According to Paragraph N-414.4 of Reference C-7 and Sections NB-3222 and NB-3223 of Reference C-8, structural adequacy is met if the maximum primary plus secondary (P+Q) stress intensity range (Sn) at a location on the component is less than 3 Sm f the material.
o If the 3 Sm limit is not met, then plastic behavior is assumed and the simplified C-8
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 elastic-plastic analysis of Paragraph NB-3228.3 of' Reference C-8 can be used to determine structural adequacy.
For those components that do not meet the requirements of Paragraph N-415.1 of Reference C-7 or Paragraph NB-3222.4(d) of Reference C-8, a fatigue evaluation must be made to assure that the component does not fail by material fatigue. For adequacy, the cumulative fatigue usage factor must be less than 1.0.
3.3.4 Procedure for Calculatine Power Rerate P+O Stress Intensity Rance The following general procedure is used for calculating the power rerate P+Q stress intensity range (Sn,new) for the limiting location on the RPV component of interest. This power rerate value will then be compared with the ASME Code stress limit described in Section 3.3.3 of this report:
(1) Determine the pressure scaling factors [(SCF)p] and the thermal scaling factors
[(SCF)T] for all stress cycles originally evaluated using the appropriate power rerate operating condition changes. If there is a large flow rate increase, include the flow rate scaling factor [(SCF)p] in the thermal scaling factor. For each stress cycle, select the larger of the thermal and scaling factors. This value is SCF.
(2) Multiply the P+Q stress intensity values from the original governing stress report by the SCF for each stress cycle to obtain the power rerate P+Q stress intensity.
The original governing stress report is the most recent stress report listed in Paragraph 2.1 of Reference C-1.
(3) Determine the maximum value of the range through which the power rerate stress intensities (calculated in Step 2) fluctuate over time. This value is Sn,new. Note that there are two stress conditions associated with the value of Sn-(4) Determine the allowable P+Q stress intensity range (3 Sm,new) evaluated at the maximum temperature of the two limiting stress cycles of Step 3. However, according to Note 1 of Figure N-414 of Reference C-7, the averace value of 3 Sm for the highest and lowest temperatures of the metal during the transient can be C-9 1
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 used in the analysis if it can be shown that the secondary stress is due to thermal loads and not mechanical loads.
(5) Compare Sn,new with 3 Sm,new-If Sn,new < 3 Sm,new, the ASME Code stress limit is met. If Sn,new > 3 Sm,new, then the guidelines of Paragraph NB-3228.3 of Reference C-8 must be followed, using power rerate values where applicable.
3.3.5 Procedure for Power Rerate Faticue Evaluation The following general procedure is used for calculating the power rerate cumulative fatigue usage factor (Unew) for the limiting location on the RPV component of interest. This power rerate value will then be compared with the ASME Code stress limit described in Section 3.3.3 of this report.
(1) Multiply the original P+Q+F stress intensity values of the original governing stress report by the scale factor for each stress cycle to obtain the power rerate P+Q+F.
(2) For each of the limiting stress cycle pairs used in the fatigue analysis of the original governing stress report, determine the absolute value of the difference of the power rerate P+Q+F stress intensities (calculated in Step 1). This value is Sp + Q + F,new-(3) Determine the power rerate alternating stress intensity, Salt,new, for each of the original limiting stress cycle pairs as follows:
Salt,new
= (1/2)
- Ke,new * (Ec/Ea)
- S +Q+F,new P
where Ke,new
= simplified clastic-plastic factor
= 1.0, Sn,new < 3 Sm,new
= 1 + (1-n)/(n(m-1)) * (Sn,new/(3Sm,new) - 1),
3Sm,new < Sn,new < 3mSm, nee
= 1/n, Sn,new > 3mSm,new C-10
r APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 t
(Ec/Ea)
= elastic modulus correction factor t
I (both factors are described in Paragraphs NB-3228.3 and NB-3222.4 of Reference C-8).
(4)
Use Salt,new as the value of the ordinate when entering the applicable design fatigue curve in Reference C-7 or Reference C-8 to find the corresponding I
allowable number of cycles (N,new) for each of the limiting stress cycle pairs.
i (5)
Calculate the power rerate incremental fatigue usage factor (U,new =
i
/N,new) for each of the limiting stress cycle pairs, where ni s the lesser of the i
ni i
actual number of design cycles for each pair. The lesser number is used because the value of the P+Q+F stress intensity range for the limiting stress cycle pair is
)
only experienced by the component over the lesser number of cyc'.es.
(6)
Calculate the power rerate cumulative fatigue usage factor (Unew = EU,new).
i If Unew < 1.0, the ASME Code stress limit is met.
3.4 Emergency and Faulted Conditions 1
In general, the stresses due to emergency and faulted conditions are based on loads such as peak pressure, which remain unchanged. Therefore, Section NB-3224 and NB-3225 Code requirements (Reference C-8) are still met for all RPV components analyzed.
4.0 POWER RERATE STRESS ANALYSIS FOR BOLTING MATERIAIS 4.1 Design Conditions Since there are no changes in the design conditions due to power rerate, the bolt design stresses remain unchanged and the ASME Code requirements of Section N-416 of
)
i Reference C-7 and Section NB-3231 of Reference C-8 are still met.
4 C-11
--g-y w--
y
+r e
aw -
- w---ey w i
,,og-w-y
,-e--w,w y
v-wwryiy-my-wy
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 4.2 Normal, Upset, and Emergency Conditions 4.2.1 Effects of Chances in Pressure and Temperature In general, changes in normal operation pressures and temperatures will increase the bolt service stresses, both averaged across the bolt cross section and at the periphery of the bolt cross section, and increase the peak bolt stresses.
4.2.2 ASME Code Stress Limits for Normal. Uoset and Emercency Conditions According to Paragraph N-416.1 of Reference C-7 and Paragraph NB-3232.1 of Reference C-8, structural adequacy is met if the maximum value of service stress, averaged across the bolt cross section and neglecting stress concentrations, is less than 2 Sm of the bolting material.
According to Paragraph N-416.1 of Reference C-7 and Paragraph NB-3232.2 of Reference C-8, structural adequacy is met if the maximum value of service stress at the periphery of the bolt cross section is less than 3 Sm f the bolting material.
o For those components that do not meet the requirements of Paragraph N-415.1 of Reference C-7 or Paragraph NB-3222.4(d) of Reference C-8, a fatigue evaluation must be performed to assure that the bolts do not fail bv material fatigue. For adequacy, the cumulative fatigue usage factor must be less than 1.0.
4.2.3 Procedure for Calculatine Power Rerate Service Stresses The general procedure for calculating the power rerate service stresses for the limiting location on the bolt and evaluating them against the ASME Code stress limits described in Section 4.2.2 of this report is as follows:
(1)
Determine the pressure scaling factors [(SCF)p] and the thermal scaling factors
[(SCF)T] for all stress cycles originally analyzed using the appropriate power rerate operating condition changes. For each stress cycle, determine the larger of the two scaling factors. This value is SCF.
C-12
APPENDDC C GE-NE-523-61-0593 DRF 137-0010-6 (2)
Multiply SCF for each stress cycle by the original service stress values, both averaged across the bolt cross section and at the bolt periphery, of the original governing stress report.
(3)
Determine the allowable service stress values, 2 Sm.new for the service stress averaged across the bolt cross section and 3 Sm,new for the service stress at the bolt periphery, evaluated at the maximum temperature of the limiting stress cycles.
I i
(4)
Compare the power rerate service stresses to the allowable values. If the allowable values are not exceeded, the ASME Code stress limits are met.
4.2.4 Procedure for Power Rerate Fatigue Evaluation The general procedure for calculating the power rerate cumulative fatigue usage factor (Unew) for the limiting location on the bolt as compared to the ASME Code limit described in Section 4.2.2, is as follows:
(1)
Multiply the original peak bolt stress values of the original governing stress report by the corresponding SCF for each stress cycle.
(2)
For each of the limiting stress cycle pairs used in the fatigue analysis of the original governing stress report, determine the absolute value of the difference of the power rerate peak bolt stresses (calculated in Step 1). This value is S eak,new-j l
p (3)
Determine the power rerate alternating stress intensity (Salt,new) for each of the original limiting stress cycle pairs as follows:
Salt,new
= (1/2) * (Ec/Ea)
- S eak,new, p
where (Ec/Ea)
= elastic modulus correction factor (See Section 3.3.5).
C-13
l l
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 (4)
Use Salt,new as the value of the ordinate when entering the design fatigue curve (Figure I-9.4) of Reference C-8 to find the corresponding allowable number of cycles (N,new) for each of the limiting stress cycle pairs.
i (5)
Calculate the power rerate incremental fatigue usage factor (U,new -
i
/N,new) for each of the limiting stress cycle pairs, where ni s the lesser of the i
i ni actual number of cycles for each pair. The lesser number is used because the value of the peak bolt stress range for the limiting stress cycle pair is only experienced by the bolt over the lesser number of cycles.
(6)
Calculate the power rerate cumulative fatigue usage factor (Unew = EU,new).
i If Unew < 1.0, the ASME Code stress limit is met.
4.3 Faulted Conditions In general, the stresses due to faulted conditions are based on loads such as peak pressure, which remain unchanged. Therefore, Section NB-3225 Code requirements (Reference C-8) are still met for all RPV components analyzed.
5.0 COMPONENT ANALYSIS 5.1 Selection Criteria for Power Rerate Stress Analysis The RPV components analyzed in this appendix are shown in Table C-5-1 alcag with the original fatigue usage factors. Further discussion on the selection of these components is given in Section 3.0 of the main report.
In addition to the requirements of Reference C-1, the number of cycles for each event over the 40 year life of the plant was modified based on plant operation experience. Numbers of cycles and fatigue usage, are taken from the main body of this report. Additional input is taken from the original vessel design reports (References C-10 and C-11), or GE stress j
reports for modified components.
l l
i 1
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 l
Table C-5-1
SUMMARY
OF ORIGINAL FATIGUE USAGE FACTORS RPV Original Cum ative (I) -
Component Fatigue Usage Factor l
I Closure Bolts 1.09 Feedwater Nozzle 0.795 Support Skirt 0.896 Refueling Containment Skirt 0.583 Recirculation Inlet Nozzle 0.511 l
Notes:
(1)
Fatigue values are taken from the main body of this report.
l l
I
{
C-15 i
i I
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.2 Closure Bolts l
5.2.1 Results of Oricinal Analysis The location of highest original cumulative fatigue usage in the closure assembly is on the bolts. The original fatigue usage factor calculated in Reference C-10 is 0.762 at the bottom l
outside surface of the bolt.
l The maximum original value of the service stress averaged across the bolt cross section is 47.1 ksi, and the maximum original value of the service stress at the bolt periphery is 98.9 ksi at the bottom inside surface of the bolt.
5.2.2 Power Rerate Service Stress Limit Check The service stress limits of Section 4.2.2 of this report must be met after applying the power rerate operating conditions to the original analysis.
The original values of the maximum service stresses affected by power rerate operating conditions are scaled up by the appropriate scale factors (SCFs). The combined stress cycles used in the original analysis are retained in this analysis.
The results of the power rerate maximum membrane stress analysis are shown in Table C-5-2. It can be seen that the power rerate value is below the allowable limit of 2 Sm-The original Sm values from Report #3 of Reference C-10 are used. The largest overall (bounding) value of the service stress averaged across the bolt cross section is 48.9 ksi due to the composite Startup stress cycle.
l l
C-16
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 l
i Table C-5-2 i
POWER RERATE MAXIMUM SERVICE STRESS RESULTS l
FOR THE CLOSURE BOLTS Maximum Maximum Membrane Membrane
+ Bending Stress Stress Original Stress Report Values:(1) 47.1 98.9 i
(x Power Rerate Scaling Factors): 1.038 1.038 Power Rerate Values:
48.9 102.7 Allowable:
2Sm =73.4 3Sm = 110.1 Note: (1)
Values taken from p. B-16-5 of Reference C-10, Report 3. All stresses in units of ksi.
i i
i C-17 l
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 The results of the power rerate maximum membrane plus bending stress analysis are shown in Table C-5-2. The largest overall value of the service stress at the bolt periphery is 102.7 ksi (due to the Shutdown Flooding stress cycle), and meets the allowable of 110.I ksi.
5.2.3 Power Rerate Fatieue Usace Check The fatigue usage limit of Section 4.2.2 of this report must be met after applying the power rerate operating conditions to the original analysis.
The original values of the peak bolt stresses are scaled up by the appropriate scale factors.
These power rerate peak bolt stresses are used in the power rerate fatigue evaluation.
The results of the power rerate fatigue evaluation can be found in Table C-5-3. The original analysis used the high strength bolting design fatigue curve (Figure N-415), which was introduced in the Winter 1967 Addenda to the 1965 ASME Code (Reference C-12).
Also, the allowable number of cycles is determined conservatively from the 3.0 Sm curve.
These methods are used in the current analysis as well. Number of cycles are taken from Section 4.4 of the main body of this report.
The results of Table C-5-3 show that the power rerate cumulative fatigue usage factor is 1.09, which is above the allowable limit of 1.0.
5.2.4 Discussion of Power Rerate Results The fatigue usage of the bolts is greater than 1.0. Two options are available to address the high usage. The first would be to examine the closure stud analysis to remove any conservatisms. The second option is to inspect the studs per ASME Code requirements and GE RICSIL 055 Rev. I recommendations and replace the studs if indications are found.
C-18
i APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-3 POWER RERATE FATIGUE RESULTS FOR THE CLOSURE BOLTS A
Component: Closure Region Bolts Location:
Bottom Outside (Ec/Es)=
1.000 (1)
Old New (2)
Stress Cycle Peak Scale Peak Allow Actual incremental Pair Stress Factor Stress 5,r 5,a N
n Ui Tension 302 1.0 302 302 151 340 120 0.353 untension 0.0 1.0 0.0 Hydrotest 302 1.0 302 302 151 340 6
0.018 Zero Load Design Pressure Insignificant i
Pressure Test Hestup
-19 1.038
-19.7 321.7 160.9 300 216 0.72 Cooldown 302 1.0 302
)
total "
Notes:
(1) As used in the original stress report fatigue evaluation (Reference C-10).
(2) Allowable number of cycles determined by the 3.0 Se curve of Figure N-416 (Reference C 12).
I i
i C-19 l
t
=
P APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.3 Feedwater Nozzle i
5.3.1 Results of Original Analysis P
The location of highest original cumulative fatigue usage on the feedwater nozzle is on the outside of the safe end near the upstream end of the thermal sleeve (Element 126). Note that the feedwater nozzle was modified to accept a new' thermal sleeve after the original RPV vendor stress report was issued. Therefore, the governing stress report is Reference C-13 for rapid cycling and Reference C-14 for system cycling. The fatigue usage factor calculated in Section 4.1 of the main body of this report is 0.795, and is almost entirely due to system cycling. Power rerate is expected to yield no increase in rapid cycling. The governing ASME Code is Reference C-8.
5.3.2 Power Rerate P+O Stress Intensity Limit Check l
i The primary plus secondary stress intensity limit of Section 3.3.3 of this report must be i
met after applying the power rerate operating conditions to the original analysis.
l Examination of Reference C-14 reveals that it does not provide P+Q stress intensity values for each stress cycle. It only provides thermal and mechanical stress components.
Therefore, the simplified procedure of Section 3.3.2 for directly scaling the stress intensity values cannot be used. The scaling factors must be applied at the stress component level instead. Then, values of the power rerate P+Q stress intensity range for certain stress cycle pairs, the same as those used in the fatigue analysis, can be determined after first I
calculating the power rerate principal stresses using the procedure of Paragraph NB-3216.2 of Reference C-8.
The power rerate scaling factors were applied to the P+Q stress components; specifically, j
the temperature scale factors were applied to thermal stresses, and the pressure scale factors were applied to mechanical and pressure stresses. Then, P+Q stress intensity range values were determined (Table C-5-4) using Sm values evaluated at power rerate conditions.
l C-20 1
Tabte C-5-4 POWER RERATE P+Q STRESS INTENSITY RESULTS FOR THE FEEDWATER N0ZZLE Casponent: Feedwater Nozzle Material: SA-105 Cl 11 Carbon Steet Location: Element 126 m
Stress Cycle sigma sigma sigma teu detta dette detta detta prin prin prin 51 SI SI Range Allow Pair RR ZZ TT RZ sig RR sig ZZ sig TT teu RZ sig 1 sig 2 sig 3 12 23 31 Sn 35m LFP.14 0 -75722 -41985 922 DSNHYDRO 0
5323 9574
-262 0 -81045 -51559 1184 17 -81062 -51559 81080 -29503 -51576 81080 55080 0
h.
DSNHYDRO O
5323 9574
-262 TRco.14 0 -67811 -35453 956 0 73134 45027
-1218 73154
-20 45027 73175 -45047 -28127 73175 55080 ZER0 LOAD 0
0 0
0 Q
TRCD.14 0 -67811 -35453 956 0 67811 35453
-956 67824
-13 35453 67838 -35466 -32371 67838 55740 y
t.b.
W ZEROLO40 0
0 0
0 MSB.14 0 -66973 -34811 940 0 66973 34811
-940 66986
-13 34811 66999 -34824 -32175 66999 55740
-6 TR.14 0
-4069 13329 1218 HSs.14 0 -66973 -34811 940 0 62904 48140 278 62905.
-1 48140 62906 -48141 -14765 62906 55740 u
LFP9 0 -10444
% 34 1353 U
)
h mss.14 0 -66973 -34811 940 0 56529 44445 413 56532
-3 44445 56535 -44448 -12087' 56535 55080 FH864.5 0 -15558 3856 1034 C
MSB.14 0 -66973 -34811 940 0-51415 38667 94 51415
-0 38667 51415 -38667 -12748 51415 55740 4
s N00PM 0 -18021 2216 1019 HSB.14 0 -66973 -34811 940 0 48952 37027 79 48952 37027 48952 -37027 -11925 48952 55740 g) m N00PM 0 27023 13064
-1297 50102.73 0 -14556 -17270
-713 0 41579 30334
-584 41587
-8 30334 41595 -30342 -11253 41595 62340 Note:
(1) All stress values and stress intensity values are in units of psi.
a
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 l
The events during which stresses were calculated are abbreviated as shown below:
l Event Abbreviation i
Turbine Roll Cooldown at 0.14 Minutes TRCD.14 l
Turbine Roll Warmup at 0.14 Minutes TR.14 j
Feedwater Heater Bypass at 64.5 Minutes FHB64.5 i
Loss of Feedwater Pumps at 0.14 Minutes LFP.14 f
I.oss of Feedwater Pumps at 9 Minutes LFP9 Normal Operation, Minimum Conditions NOOPM h
Hot Standby at 0.14 Minutes HSB.14 Shutdown at 102.73 Minutes SD102.73 l
Design Hydrotest DSNHYDRO l
l-The results of Table C-5-4 show that the maximum P+Q stress intensity range, Sn, which l
is equal to 81.0 ksi (due to the LFP.14 - DSNHYDRO stress cycle pair), exceeds the l
allowable 3 Sm limit (for that pair) of 55.1 ksi. Thus, the simplified elastic-plastic analysis of Paragraph NB-3228.3 of Reference C-8 must be performed.
l 5.3.2.1 Removal of Thermal Bending Stresses i
j According to Paragraph NB-3228.3(a) of Reference C-8, the P+Q stress intensity range, j
after removing thermal bending stresses, shall be less than 3 Sm. Power rerate P+Q stresses are recalculated after removing the thermal bending stresses from the stress components. These values are then used to determine the P+Q (minus thermal bending) l stress intensity range for the desired stress cycle pairs.
l The results of Table C-5-5 show that the maximum P+Q (minus thermal bending) stress i
intensity range is 29.1 ksi (due to the ZEROLOAD - LFP.14 stress cycle pair), which is I
less than the allowable 3 Sm limit (for that pair) of 55.1 ksi. Thus, the requirement of 1
j Paragraph NB-3228.3(a) is met.
i C-22 i
i 1
2
~.
m
._m-
... ~ _. - _... _. <
I i
Table C-5-5 POWER RERATE P+0 STRESS INTENSITY RESULTS FOR THE TEEDWATER N0ZZLE l
EXCLUDING THERMAL BENDING STRESSES-t Component: Feedwater Nottle Materlat: SA-105 Cl 11 Carbon Steel Location:
Element 126 y
T
- C
~ !
i (T1 Stress Cycle sigma slyne sigma teu detta dette detta dette prin prin prin SI SI St Range Attow Z
)
i Pair RR ZZ TT RZ sig RR sig ZZ sig TT teu RZ sig 1 sig 2 sig 3 12 23 31 Sn 3Sm U_
l X
O LFP.14 0 28818 20210
-1394 DSNNYDRO O
8152 10256
-408 0 20666 9954
-966 20713
-47 9954 20760 -10001 -10759 20760 55080 atil h'
DSNNYDRO O
8152 10256
-408 TRCD.14 0 28818 19790
-1360 0 -20666
-9534 952 44 -20710
-9534 20754 -11176
-9578 20754 55080 a
y O
y ZER0 LOAD 0
0 0
0 w
TRCD.14 0 28818 19790
-1360 0 -28818 -19790 1360 64 -28882 -19790 28946
-9092 -19854 28946 55740 W
MSB.14 0 28118 19734
-1376 0 -28118 -19734 1376 67 -28185 -19734 28252
-8451 -19801 28252 55740 ZER0 LOAD 0
0 0
0 c
ta TR.14 0 28117 11010
-1098 HSB.14 0 28118 19734
-1376 0
-1
-8724 278 278
-279
-8724 556 8445
-9002 9002 55740 g
LFP9 0 28117 12409
-963 HSs.14 0 28118 19734
-1376 0
-1
-7325 413 413
-414
-7325 826 6911
-7738 7738 55080
-wY 8
i Note:
(1) All stress values and stress intensity values are in units of psi.
9 i
e i
1 i
a
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.3.2.2 Calculation of Ke and the Cumulative Fatigue Usage Factor Per Paragraphs NB-3228.3(b) and NB-3228.3(c) of Reference C-8, a fatigue evaluation is performed taking the elastic-plastic factor (Ke) into account. A conservative method for calculating the power rerate value of Salt was developed using the original Salt values and scale factors. Since Salt,old = (1/2)
- Ke,old * (Ec/Ea)
- S +Q+F,old P
it is assumed that:
Salt,new = (1/2)
- Ke,new * [(Ec/Ea)
- S +Q+F,old]
- SCF P
By combining the above two equations, we have:
Salt,new = Salt,old
- SCF * (Ke,new / Ke,old)
It should be noted that this method requires the determination of the original Sn values (for calculating Ke,old). Original Sn values for certain stress cycle pairs are determined from the original stress components from Reference C-14. These values are then used in the fatigue evaluation of Table C-5-6. The stress cycle pairs TT126 - HSB.14 and SVB1 -
NOOPM were neglected since they contribute very little to fatigue usage.
The results of Table C-5-6 show that the power rerate fatigue usage factor is 0.894, which is below the allowable limit of 1.0.
1 C-24
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-6 POWER RERATE FATIGUE RESULTS FOR THE FEEDWATER NOZZLE Location:
Safe end, Element 126 Material:
SA-105 Cl 11 Carbon Steel m=
3.0 n=
0.2 i
(2)
(1)
New (1)
New Stress Old Old Old Old New New New Mew Scale Old New Allow Actual increm Cycle Pair Sn 3Se 3mSe Ke
$n 3Sa 3n6m Ke Factor S att S att N
n Ui LFP.14 DSNNYDRO 80.0 54.6 163.8 1.929 81.1 55.1 165.2 1.944 1.038 116.3 121.7 382 60 0.157 DSNHYDRO TRCD.14 72.1 54.6 163.8 1.641 73.2 55.1 165.2 1.657 1.038
%.4 101.1 622 70 0.113 ZER0 LOAD TRCD.14 66.8 54.6 163.8 1.445 67.8 55.7 167.2 1.434 1.038 72.2 74.3 1392 186 0.134 ZER0 LOAD HSB.14 65.9 54.6 163.8 1.415 67.0 55.7 167.2 1.404 1.038 69.8 71.9 1522 200 0.131 TR.14 HSB.14 61.4 54.6 163.8 1.251 62.9 55.7 167.2 1.257 1.038 56.5 59.0 2684 454 0.169 LFP9 HSS.14 55.2 54.6 163.8 1.021 56.5 55.i
'33.2 1.053 1.038 41.7 44.6 6215 60 0.010 FHB64.5 HSB.14 50.0 54.6 163.8 1.000 51.4 55.7 167.2 1.000 1.115 36.3 40.5 8273 286 0.035 N00PM HSB.14 47.7 54.6 163.8 1.000 49.0 55.7 167.2 1.000 1.038 34.7 36.0 12012 1600 0.133 N00PM S0102.73 41.0 54.6 163.8 1.000 41.6 62.3 187.0 1.000 1.038 31.5 32.7 16754 216 0.013 U tot = 0.894 Notes:
(1) From p. 274 of Reference C 14 (2) Determined using fatigue curve for carbon steels (Fig. I-9.1 of Reference C 8).
(3) All stress intensities are in units of ksi.
C-25 l
l c,--
I APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.3.2.3 Thermal Stress Ratchet Check l
l Per Paragraph NB-3228.3(d) of Reference C-8, the thermal stress ratchet check of l
Paragraph NB-3222.5 of Reference C-8 must be performed. The analysis shows that compliance is met, as follows:
1 Maximum power rerate general membrane stress due to pressure:
06M,old = Pold
- R / t, l
where Pold = original max. normal operating vessel pressure (from p. 80 of Reference C-14)
Therefore, 06M,new = 36M,old * (Pnew / old)
P
= 7985 psi * (1038 psig /1000 psig)
= 8288 psi new = 36M,new / (1.5 Sm,new) x
= 8288 psi / (1.5
- 18360 psi)
= 0.301 where Sm,new is evaluated at 565'F.
For 0 < x < 0.5, y' = 1/x and y' = Sn / (1.5 Sm)
Sn = 1.5 Sm / x (max, allowable value for Sn)
Sn = 1.5
- 18360 psi / 0.301
- 91.51 ksi l
l The largest cyclic range of thermal stress (mechanical stress conservatively included) occurring at Element 126 is 81.1 ksi. Therefore, since 81.1 ksi < 91.51 ksi, no thermal.
stress ratchet effect will be experienced at that location.
C-26
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.3.2.4 Maximum Temperature and Strength Ratio Checks According to Paragraph NB-3228.3(e) of Reference C-8, the temperature used in the analysis should not exceed those of the table of Paragraph NB-3228.3. The maximum temperature in the table is 700*F for carbon steel, and the maximum analysis temperature is 565*F for any stress cycle for the nozzle safe end. Since 565 F < 700*F, this requirement is met.
According to Paragraph NB-3228.3(f), the ratio of the material's minimum specified yield strength (S,ms) to the minimum specified ultimate strength (Su,ms) shall be less than y
0.80. From Table I-1.1 of Appendix I of Reference C-8, S,ms = 36.0 ksi for SA-105 Cl y
II carbon steel and Su,ms = 70.0 ksi, which gives us a ratio of 0.51. Since 0.51 < 0.80, this requirement is met.
5.3.3 Discussion of Power Rerate Results
)
The power rerate analysis for the feedwater nozzle safe end shows that the Code stress limits are met and that the structural integrity of the feedwater nozzle is acceptable for the power rerate conditions and modified cycles.
j 5.4 Support Skirt 5.4.1 Results of Orieinal Analysis The location of highest original cumulative fatigue usage on the support skirt is at the inside surface of juncture 3 (node 260). The original fatigue usage factor calculated in Section 4.2 of the main body of this report is 0.896.
5.4.2 Power Rerate P+O Stress Intensity Limit Check The primary plus secondary (P+Q) stress intensity limit of Section 3.3.3 of this report must be met after applying the power rerate operating conditions to the original analysis.
C-27
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 The P+Q stress intensity values from Appendix A of this report were scaled up by the j
scale factors for each cycle. The results listed in Table C-5-7 show that the maximum power rerate P+Q stress intensity range (Sn) is 99.4 ksi, which exceeds the allowable of 3 Sm = 80.1 ksi. Thus, the simplified elastic-plastic analysis of Reference C-8 must be i
l performed.
5.4.2.1 Removal of Thermal Bendine Stresses According to Paragraph NB-3228.3(a) of Reference C-8, the P+Q stress intensity range, after removing thermal bending stresses, shall be less than 3 Sm. Power rerate P+Q stresses are recalculated after removing the thermal bending stresses from the stress components. These values are then used to determine the P+Q (minus thermal bending) stress intensity range for the desired stress cycle pairs.
l The results of Table C-5-7 show that the maximum P+Q (minus thermal bending) stress intensity range is 75.8 ksi, which is less than the allowable 3 Sm limit of 80.1 ksi. Thus, the requirement of Paragraph NB-3228.3(a) is met.
5.4.2.2 Calculation of K; and the Cumulative Fatieue Usage Factor Per Paragraph N-417.6 of Reference C-7, a fatigue evaluation is to be performed taking the elastic-plastic factor (Ke) into account.
The original values of the P+Q+F stress intensities are scaled up by the scale factors.
These power rerate P+Q+F stress intensity values are used in the power rerate fatigue evaluation shown in Table C-5-8. The incremental fatigue usage factor due to the additional seismic analysis is added.
Table C-5-8 shows the results of the power rerate cumulative fatigue evaluation. The calculated cumulative fatigue usage factor exceeds 1 if power rerate and the modified cycles is considered for 40 years. The number of allowable years at the power uprate condition with the modified cycles can be determined by restricting the fatigue usage to C-28
Table C-5-7 POWER RERATE P*Q STRESS INTENSITY RESULTS FOR THE SUPPORT SKIRT Location:
Node 260 Stress Pair sigma signa sigma tau delta delta detta dette prin prin prin 51 St SI Range Allow w/ th bend XX YY ZZ XY sig XX sig YY sig ZZ teu XY sig 1 sig 2 sig 3 12 23 31 Sn 35m HU/C0(6,1)
-10162 -53017 847 2957 T
HU/CD(26,1) 4903 26116
-1187
-1332 -15065 -79133 2034 4289 -14 779 -79419 2034 64640 -81453 16813 81453 80100 T
HU/CD(26,1) 4903 26116
-1187
-1332 LOFP(41,1)
-10466 -54642 952 3058 15369 80758
-2139 4390 81051 15076
-2139 65976 17215 -83190 83190 80100
-X HU/CD(6,1)
-10162 -53017 847 2957 O
EX HU/CD(26,1) 6016 31755
-3649
-1753 -16178 -84772 44 %
4710 -15856 -85094 4496 69238 -89590 20352 89590 80100 Q
LOFP(16,1) 6179 33495
-3570
-1547 7
LOFP(41,1)
-10466 -54642 952 3058 16645 88137
-4522
-4605 88432 16350
-4522 72083 20872 -92954 92954 80100 Z
(T1 LOFP(16,1) 6179 33495
-3570
-1547 6
Q EX HU/CD(6,1) -11308 -58851 3211 3384 17487 92346
-6781
-4931 92669 17164
-6781 75506 23945 -99450 99450 80100 y
u C
3 w/o th bend 6
tA0 HU/CD(6,1)
-8573
-5213 16652 2957 HU/CD(26,1) 4021 391
-6712
-1332 -12594
-5604 23364 4289
-3566 -14632 23364 11065 -37996 26930 37996 80100 U
PC HU/CD(26,1) 4021 391
-6712
-1332 T
LOFP(41,1)
-8820
-5241 17279 3058 12841 5632 -23991
-4390 14917 3556 -23991 11360 27547 -38908 38908 80100 C
4 HU/CD(6,1)
-8573
-5213 16652 2957 g
EX HU/CD(26,1) 4926 501 -14046
-1753 -13499
-5714 30698 4710
-3496 -15717 30698 12221 -46415 34194 46415 80100
_9 LOFP(16,1) 4896
-2492 -15602
-1547 m
LOTP(41,1)
-8820
-5241 17279 3058 13716 2749 -32881
-4605 15393 1072 -32881 14321 33953 -48274 48274 80100 LOFP(16,1) 4896
-2492 -15602
-1547 EX HU/CD(6,1) -11308 -58851 3211 3384 16204 56359 -18813
-4931 56956 15607 -18813 41348 34420 -75769 75769 80100 Note:
(1) All stress values and stress intensity values are in units of psi.
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-8 POWER RERATE FATIGUE RESULTS FOR TIIE SUPPORT SKIRT Location:
Node 260 Material:
SA-533B Cl.1 Low Alloy Steet l
m=
2.0 n=
0.2 (Ec/Ea) 1.071 (1) l (4)
Old New (3)
(4)
Stress Old New Range Range allow act increm.
Cycle Pair Sn Sn 35m 3mse Ke 5,p+q+f S,p+q+f S, alt u
n UI HU/CD(6,1)
HU/CD(26,1) 79.9 81.5 80.0 160.0 1.07 109.8 112.0 116.8 389 246 0.632 HU/CD(26,1)
LOFP(41,1) 81.2 83.2 80.0 160.0 1.16 111.6 114.3 128.9 309 10 0.032 HU/CD(6,1)
EX HU/CD(26,1) 87.8 89.6 80.0 160.0 1.48 100.0 101.9 160.3 184 10 0.054 i
LOFP(16,1) l LOFP(41,1) 91.1 93.0 80.0 160.0 1.65 124.0 126.5 203.6 102 20 0.197 i
i i
LOFP(16,1)
EX HU/CD(6,1) 97.9 99.5 80.0 160.0 1.97 132.7 134.8 260.1 57 10 0.175 l
l l
Power Rerate 40 Year usage = 1.090 l
i Rerate in 1993 Years at terate Power =
21 Usage = 0.572 Years at Current Power =
19 usage = 0.426 j
Total Usage = 0.998 Notes:
(1) As used in the original fatigue evaluation (Appendix A).
(2) All stress intensity values are in units of ksi.
(3) Determined by using figure N 415(A) of Reference C-7, f atigue curve for low alloy steels (4) Taken from Section 4.2 of the main body of this report.
C-30
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 below 1.0. The results of this calculation show that the number of allowable years with the power rerate conditions is 21.
5.4.2.3 Thermal Stress Ratchet Check Since there is no membrane stress in the skirt due to pressure, the thermal stress ratchet check of Paragraph N-417.3 of Reference C-7 need not be performed.
5.4.3 Discussion of Power Rerate Results The power rerate analysis for the support skirt shows that Code stress limits are met and the structural integrity of the support skirt is acceptable for 21 years of operation with power rerate and modified cycles.
5.5 Refueling Containment Skirt 5.5.1 Results of Original Analysis The location of highest original cumulative fatigue usage on the refueling containment skirt is at the outside surface of juncture 1. The fatigue usage factor calculated in Section 4.3 of the main body of this report is 0.583.
5.5.2 Power Rerate P+O Stress Intensity Limit Check The primary plus secondary (P+Q) stress intensity limit of Section 3.3.3 of this report must be met after applying the power rerate operating conditions to the original analysis.
The P+Q stress intensity values from the original stress report (Table C-5-9) were scaled up by the scale factors for each cycle. The results listed in Table C-5-9 show that the maximum power rerate P+Q stress intensity range (Sn) is 90.0 ksi, which exceeds the i
allowable of 2 Sy = 88.0 ksi. Thus, the simplified elastic-plastic analysis of Reference C-15 must be performed. The method described in Paragraph NB-3228.3 of Reference C-8 cannot be completed, as there is insufficient information. Therefore, the C-31 1
{
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-9 POWER RERATE P+Q STRESS INTENSITY RESULTS FOR THE REFUELING CONTAINMENT SKIRT Stresses P
P+Q Original Stress Report Values: (1) 7.0 86.7 (x Power Rerate Scaling Factors):
1.038 1.038 Power Rerate Values:
7.3 90.0 Allowable:
40.0 88.0 1.5 Sm 2 Sy Notes: (1) Values taken from pp. B-11-1 & B-ll-21 of Reference C-10, Report #9.
All stresses in units of ksi.
I i
C-32 1
l APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 elastic-plastic analysis method used in the original sitess report for the support skirt (Reference C-10, Report #8) is used.
5.5.2.1 Calculation of Ke and the Cumulative Fatigue Usage Factor Per Paragraph N-417.6 of Reference C-7, a fatigue evaluation is to be performed taking the elastic-plastic factor (Ke) into account.
The original values of the P+Q+F stress intensities are scaled up by the scale factors.
These power rerate P+Q+F stress intensity values are used in the power rerate fatigue evaluation shown in Table C-5-10. The incremental fatigue usage factor due to the additional seismic analysis is added.
Table C-5-10 shows the results of the power rerate cumulative fatigue evaluation. The calculated cumulative fatigue usage factor is 0.777, which is below the allowable value of 1.0.
5.5.2.2 Thermal Stress Ratchet Check Since there is no membrane stress in the skirt due to pressure under normal operation, the thermal stress ratchet check of Paragraph N-417.3 of Reference C-7 need not be performed.
5.5.3 Discussion of Power Rerate Results The power rerate analysis for the refueling containment skirt shows that Code stress limits are met and the structural integrity of the refueling containment skirt is acceptable for the power rerate conditions.
C-33
1 i
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 i
i 1
i Table C-5-10 a
l POWER RERATE FATIGUE RESULTS FOR THE REFUELING CONTAINMENT SKIRT 2
l i
i Location:
J meture 1, outside surface noterial: SA 302 Grade B i
a=
2.0 n=
0.2 (Ec/Es)= 1.000 (1) i i
(4)
(3)
(4)
I stress Old New allow act increm.
)
Cycle Pair
$n SCF Sn 35m 3mse sn/3Se Ke
$, alt N
n UI
/
a j
j All Events 78.8 1.038 81.8 80.0 160.0 1.02 1.09 100.3 600 466 0.777 I
l U tot = 0.777 i
1
)
i Notes:
(1) As used in the original stress report fatigue evaluation (Reference C 10, Report W ).
]
(2) Att stress intensity values are in units of ksi.
J i
(3) Deterwined by using Figure N-415(A) of Reference C 7, f atigue curve for low alloy steets.
2 (4) From p. B-12-1 of Reference C-10, Report M.
e r
?
4 i
C-34
l l
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 5.6 Recirculation Inlet Nozzle 5.6.1 Results of Oricinal Analysis The location of highest original cumulative fatigue usage on the recirculation inlet nozzle is on the inside surface of Section 7 of the nozzle safe end. Note that the nozzle was modified to accept a new wie end after the original RPV vendor stress report was issued.
l Therefore, tSe governing stress report is Reference C-16. The revised analysis for modified cycles is performed in Appendix B of this report. The fatigue usage factor calculated in Appendix B is 0.511. The ASME Code used in Appendix B is Reference C-17.
5.6.2 Power Rerate P+O Stress Intensity Limit Check The primary plus secondary stress intensity limit of Paragraph 3.3.3 of this report must be met after applying the power rerate operating conditions to the original analysis.
I The original stress report P+Q stress intensity values (Table C-5-13) were scaled up by the SCF values. The results of Table C-5-13 show that the maximum power rerate P+Q stress intensity range (Sn) is 70.6 ksi, which exceeds the allowable of 51.75 ksi. It should be noted that the allowable used is the 3 Sm value for the design temperature of the nozzle, which is conservative.
4 5.6.2.1 Removal of Thermal Bending Stresses Since the maximum power rerate P+Q stress intensity range (Sn) exceeds the Code allowable, an clastic-plastic analysis must be performed. According to Paragraph NB-3228.3(a) of Reference C-17, the P+Q stress intensity range, after removing thermal bending stresses, shall be less than 3 Sm. The original stress repon P+Q stress intensities are recalculated (scaled up) after removing the thermal bending stresses from the stress components.
C-35
(
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-13 POWER RERATE P+Q STRESS INTENSITY RESULTS FOR THE RECIRCULATION INLET NOZZLE Location: Section 7 Stress Intensity P+Q P+ Q-Qb Original Stress Report Values: (1)
Inside:
68.5 40.2 Outside:
68.4 39.9 (x Power Rerate Scaling Factors):
1.032 1.032 Power Rerate Values:
Inside:
70.6 41.5 Outside:
70.6 41.1 Allowable (3Sm):
51.75 51.75 Note: (1) Values taken from Section 4.5 of the main body of this report. All stresses in units of ksi.
t l
l C-36 l
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-13 results show that the maximum P+Q (minus thermal bending) stress intensity range is 41.5 ksi, which is less than the allowable 3 Sm limit of 51.75 ksi. Thus, the j
requirement of Paragraph NB-3228.3(a) is met.
5.6.3 Power Rerate Fatigue Usane Check The fatigue usage limit of Paragraph 3.3.3 of this report must be met after applying the power rerate operating conditions to the original analysis.
Power rerate P+Q stress intensity values were used to calculate the power rerate cumulative fatigue usage factor in Table C-5-14 using the methods in Reference C-17.
l Table C-5-14 results show that the power rerate cumulative fatigue usage factor is 0.549, which is below the allowable value of 1.0.
I 5.6.4 Thermal Stress Ratchet Check i
i Per Paragraph NB-3228.3(d) of Reference C-17, the thermal stress ratchet check of l
l Paragraph N3-3222.5 of Reference C-17 must be performed. The analysis shows that compliance is met, as follows:
Maximum power rerate general membrane stress due to pressure:
(PM)6 = 6.77*1.032 = 6.99 ksi (Section 4-I)
(from p. 63 of Reference C-16) l i
x = (P )0 /S M
y l
= 6.99 ksi /19.08 ksi
= 0.366 where S is taken from p. 63 of Reference C-16.
y C-37 l
l APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 Table C-5-14 POWER RERATE FATIGUE RESULTS l'
FOR THE RECIRCULATION INLET NOZZLE Location: Section 7, outside surface Material:
SA-182 Grade F316 Stainless Steel a=
1.7 (Ec/Ea) 1.000 (1) n=
0.3 (2)
Old New (3)
(2)
Stress Old New Range Range Attow Act Incram.
Cycle Pair Sn SCF Sn 3Se 3mSe Ke S,f S,f S,att N
n UI SOPCRLMU ZEROLOAD 113.6 1.013 115.0 51.75 88.0 3.33 43.1 43.7 264.5 97 40 0.413 LOFP CD0 NORMOP2 68.4 1.032 70.6 51.75 88.0 2.21 58.0 59.9 144.4 541 10 0.018 LOFP CD0 SCRAMHUP 68.2 1.032 70.4 51.75 88.0 2.20 57.7 59.6 143.0 558 40 0.072 SCRAMCDM ZER0 LOAD 58.3 1.000 58.3 51.75 88.0 1.42 48.4 48.4 76.0 5003 180 0.036 SCRAMMUP EXCESSCD 55.7 1.000 55.7 51.75 88.0 1.25 48.4 48.4 65.1 9315 10 0.001 SRVBLDWN ZER0 LOAD 54.8 1.000 54.8 51.75 88.0 1.20 44.9 44.9 59.8 13138 40 0.003 SCRAMMUP ZER0 LOAD 49.0 1.000 49.0 51.75 88.0 1.00 39.8 39.8 44.4 63139 130 0.002 NORM 0P1 i
ZER0 LOAD 45.0 1.000 45.0 51.75 88.0 1.00 36.1 36.1 40.6 103224 130 0.001 l
SHUTDOWN LOFWHHUP 42.5 1.032 43.8 51.75 88.0 1.00 36.2 37.3 40.6 103205 80 0.001 SHUTDOWN FWTEMPRD 42.1 1.000 42.1 51.75 88.0 1.00 34.4 34.4 38.2 142339 136 0.001 U=
0.549 Notes:
(1) As used in the original stress report fatigue evaluation.
l (2) Att stress intensity values are in snits of ksi. Taken f rom Appendix B.
(3) Determined from the tabulated fatigue curve values for stainless steels.
C-38
i H
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 For 0 < x < 0.5, y' = 1/x and y' = Sn/Sy Sn = S / x (max. allowable value for Sn) y Sn = 19.08 ksi / 0.366
= 52.1 ksi The largest cyclic range of thermal stress occurring at the location of interest is 43.22*l.032 = 44.6 ksi (from page 63 of Reference C-16). Therefore, since 44.6 ksi <
3 52.1 ksi, no thermal ratchet effect will be experienced at that location.
5.6.5 Maximum Temocrature and Strength Ratio Checks According to Paragraph NB-3228.3(e), the temperature used in the analysis should not exceed those of the table of Paragraph NB-3228.3. The maximum temperature from this table is 800*F for austenitic stainless steel, and the maximum analysis temperature is 552*F for any stress cycle for the recirculation inlet nozzle. Since 552*F < 800 F, this requirement is met.
i According to Paragraph NB-3228.3(f), the ratio of the material's minimum specified yield i
strength (S,ms) to the minimum specified ultimate strength (Su,ms) shall be less than y
0.80. From page 10 of Reference C-16, S,ms = 30.0 ksi for SA-182 Grade F316 i
y stainless steel and Su,ms = 70.0 ksi, which gives a ratio of 0.43. Since 0.43 < 0.80, this requirement is met.
5.6.6 Discussion of Power Rerate Results The power rerate analysis for the recirculation inlet nozzle shows that the Code stress limits are met and that the structural integrity of the recirculation inlet nozzle is acceptable for the power rerate conditions.
)
C-39 J
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6
6.0 CONCLUSION
S l
The results of the power rerate analysis are summarized in Table C-6-1. The results show that structural integrity is maintained for the feedwater nozzle, refueling containment skirt, l
and recirculation inlet nozzle for the increased coolant pressures and temperatures. The l
fatigue usage of the support skirt is below one for 21 years of power rerate and modified cycles. The closure studs do not meet the fatigue limits, and therefore must be either inspected or reevaluated.
C-40
1 i
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 1
1 i
Table C-6-1 i
SUMMARY
OF POWER RERATE STRESS AND FATIGUE USAGE RESULTS i
Orig.
Power Power Orig.
Power Power Limiting Fatigue Rerate Rerste P+Q Rerate Rerate Conponent Location usage usage Allowable SI (ksi)
P+0 $1 3Sm Closure Bolts Bottom Outside 1.09 1.09 1.0 47.1 (2) 48.9 (2) 73.4 98.9 (3) 102.7 (3) 110.1 Feedwater Notzte Element 126 0.795 0.894 1.0 80.0 81.1 55.1 l
28.1 (1) 29.0 (1) -
Support Skirt Juncture 3 0.896 0.998 (5) 1.0 103.9 107.8 80.1 Inside surface 56.4 (1) 58.6 (1)
Refueling Juncture 3 0.583 0.777 1.0 86.7 90.0 88.0 (4)
Containment Skirt inside surface l
Recire Inlet Section 7 0.511 0.549 1.0 68.5 70.6 51.75 Nozzle 40.2 (1) 41.5 (1) t i
Notes:
(1) Excluding thermal bending stresses.
l l
l (2) Maximun service stress averaged across bolt cross section.
i l
(3) Maximun service stress at periphery of bolt.
I (4) The P+Q limit was satisfied by performing the elastic-plastic analysis from Reference C 10.
(5) Allowable ruter of years limited to 21 at power rerste conditions.
l C-41
l I
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6
7.0 REFERENCES
4 l
C-1)
GE Document 25A5341 Rev. O, " Reactor Vessel -- Power Rerate, Certified Design Specification," GE-NE, San Jose, CA, November 1992.
4 C-2)
GE Document 21 All11, Rev. 9, " Reactor Pressure Vessel, Purchase Specification," APED, San Jose, CA, May 1970.
I C-3)
GE Drawing 886D499, Rev.12, " Reactor Vessel, Purchase Part Drawing,"
APED, San Jose, CA, February 1968.
l C-4)
GE Drawing 729E762, Rev. O, " Reactor Thermal Cycles," APED, San Jose, CA, May 1967.
2 C-5)
GE Drawing 135B9990, Rev.1, " Nozzle Thermal Cycles," APED, San Jose, 4
CA, May 1967.
C-6)
GE Document 22A6656, Rev. O, " Reactor Vessel (System Cycling), Design l
Specification," NEED, San Jose, CA, March 1979.
C-7)
American Society of Mechanical Engineers, " Rules for Construction of Nuclear Vessels," ASME Boiler and Pressure Vessel Code,Section III,1965 Edition with Addenda to and including Winter 1965.
j C-8)
American Society of Mechanical Engineers, " Rules for Construction of Nuclear i
Vessels," ASME Boiler and Pressure Vessel Code,Section III,1974 Edition l
with Addenda to and including Summer 1976.
i C-9)
D.R. McNeill and J.E. Brock, " Charts for Transient Temperatures in Pipes,"
Heating / Piping / Air Conditioning, November 1971.
C-10)
Babcock & Wilcox Company, " Certified Design Report for Peach Bottom 2 Reactor," Mount Vernon, IN, GE VPF 1896-142-1, December 1970.
4 C-11)
Babcock & Wilcox Company, " Final Design Document for Peach Bottom 3,"
j Mount Vernon, IN, GE VPF 1896-148-2, October 1971.
a C-12)
American Society of Mechanical Engineers, " Rules for Construction of Nuclear l
Vessels," ASME Boiler and Pressure Vessel Code,Section III,1965 Edition with Addenda to and including Winter 1967.
C-13)
GE Document 22A6648, Rev. O, "Feedwater Nozzle Stress Report (Rapid Cycling), Peach Bottom 2 & 3," San Jose, CA, September 1979.
C-14)
GE Document 22A6647, Rev. O, "Feedwater Nozzle Stress Report, Peach j
Bottom 2 & 3," San Jose, CA, August 1979.
4 1
i C-42 i
4 4
,l
I f
APPENDIX C GE-NE-523-61-0593 DRF 137-0010-6 I
l C-15)
S.W. Tagart, Jr., " Plastic Fatigue Analysis of Pressure Components," Joint Conference with the Pressure Vessels and Piping Division in Dallas, TX, September 22-25,1968, ASME Paper 68-PVP-3, May 1968.
j C-16)
GE Document 23A4274, Rev. 2, " Recirculation Inlet Nozzle and Safe End i
Certified Stress Report," GE-NEBO, San Jose, CA, August 1984.
J 1
C-17)
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