ML20235Q798
| ML20235Q798 | |
| Person / Time | |
|---|---|
| Site: | Rancho Seco |
| Issue date: | 09/14/1987 |
| From: | Schaefer R, White L BABCOCK & WILCOX CO. |
| To: | |
| Shared Package | |
| ML20235Q699 | List: |
| References | |
| 32-1167974-01, 32-1167974-1, TAC-64153, NUDOCS 8710070703 | |
| Download: ML20235Q798 (45) | |
Text
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ATTACHMENT 2 I
Calculations: . Local Heating of Pressurizer Shell Inside Corner Pressurizer Shell Temperature Due to Radiant Heat l
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.- , Babcock & Wilcox DOCUMENT
SUMMARY
SHEET a McDermott company i
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PREPARED BY: REVIEWED BY:
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REVIEWER INDEPENDENCE ,g, PURPOSE AND
SUMMARY
OF RESULTS: N
Purpose:
Revision 00 on November 21, 1986, while heating up the pressurizer at Rancho Seco Nuclear Station, the water level in the pressurizer dropped below the level of the upper bundle. This resulted in an estimated increase in heater temperature to approximately 2200F.
The radiant heat from these heaters increased the temperature of the shell. For the duration of the event, the heaters were energized and de-energized in a cyclic manner. Also, the total number of heaters energized sometimes varied with the cycles.
The purpose of revision 00 of this calculation package is to conservatively determine the maximum metal temperature resulting from the radiant heat of the exposed heater transient event.
Results: Revision 00 The maximum base metal and cladding temperature occur at the inside corner region of the heater bundle shell opening. They were conservatively calculated to be 919F and 1002F respectively.
Purpose:
Revision 01 The purpose of revision 01 is to determine the fatigue impact of the heater uncovery event on the pressurizer shell opening.
Results: Revision 01 The fatigue usage factor associated with this event was very conservatively calculated to be 0.1. The total revised fatigue usage factor for the pressurizer heater bundle opening is 0.18 which is much lese _than 1.0. the allownbla fa+iqtto il e:a_ qo #= r-t or .
THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT:
CODE / VERSION / REV CODE / VERSION / REV 4A/sys/E/C _ _ _ _
PAGE / OF '-
. v; Babcock &'Wilcox B&W Doc. No. 32-1167974-01 RECORD OF REVISIOl]E REV. NO. DESCRIPTION DAIE O Original' Release 2/87~
'l Complete revision to incorporate 9/87 fatigue analysis of shell sections.
(The microfiche of Revision 00 are applicable to Revision 01.also)
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J. ~~t Babcock & Wilcox- B&W Doc. No. 32-1167974-01 TABLE OF CONTENTS Section Description Eggg 1.0 Purpose 2 2.0 Background 2 3.0 Design Inputs- 3 4.0 Assumptions 3 5.0 Computer Programs 3-6.0 Results 3A 7.0 References 4 8.0 Thermal Calculations 5 Microfiche of Computer Runs 16 Appendices A Thermal Stress Analysis 17 B Fatigue Analysis 27 C Potential for Clad Cracking 36 Total pages = 45 l
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Inserted pages 1A, 1B, 3A & 31A
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Babcock & Wilcox B&W Doc. No. 32-1167974-01
- 1. 2.UEEQSE.
The purpose of this calculation package is to determine the base metal maximum temperature resulting from the radiant heat of the exposed heater transient event.
Revision 01 The purpose of revision 01 is to determine the fatigue impact of the heater uncovery event on the pressurizer shell opening.
- 2. BACKGROUND On November 21, 1986, while heating up the pressurizer at Rancho Seco Nuclear Station, the water level- in the pressurizer dropped below the level of the upper bundle.
This resulted in an estimated increase in heater temperature to approximately 2200F. The radiant heat from these heaters increased the temperature of the shell.
For the duration of the event, the heaters were energized and de-energized in a cyclic manner. Also, the total number of heaters energized sometimes varied with the cycles.
Reference 1 contains the " Time / Heat Output" data points relative to this " heater" transient.
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Babcock & Wilcox B&W Doc. No. 02-1167974-01
- 3. DESIGN INPUTS The definition of time vs. radiant heat output is taken from Reference 1.
The shell material designation and geometry are taken from the original fabrication drawings (Reference 6 and 7).
The radiant heat distribution on the inside surface of shell is taken from Reference 2.
The heater sheath temperature profile adjacent to the shell thickness is taken from Reference 3.
- 4. ASSUMPTIQHE This evaluation conservatively assumes that a maximum equivalent of 1.25 heater bundles were exposed. This l assumption was confirmed to be conservative as a result of the inspection performed on the middle heater bundle (i.e.,
one heater element (top element) was deformed and next lower four elements were discolored, thus, only a small portion of the middle heater bundle was ever exposed).
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- 5. COMPUTER PROGRAMS l i
The general purpose analysis program ANSYS, Reference 5, is l used to perform the thermal analysis, l
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Babcock & Wilcox B&W Doc. No. 32-1167974-01
- 6. PLillLL1, The maximum base metal temperature is conservatively calculated to be 919F and occurs at the inside corner region of the heater bundle shell opening (Node 19, see Figure 5).
.The maximum cladding temperature is conservatively calculated to be 1002F and occurs at.inside corner (Node 1,
.see Figure 6).
The fatigue usage factor associated with this event was conservatively calculated to be 0.1. The design fatigue usage factor in the current Pressurizer Stress Report, Reference 11, associated with this area is 0.075. Therefore the revised total fatigue usage factor for the pressurizer heater bundle opening area is 0.18 which is much less than 1.0, the allowable fatigue usage factor.
Prepared By /- Date 9!/k Reviewed By #d Date N /#7 Page 5/ n'$
Babcock &.Wilcox B&W Doc. No. 32-1167974-01 l
- i. 7. REFERENCES
- 1. 51-1167607-00, " Heater Element Time Intervals", Rancho i
Seco.
- 2. 32-1167978-00, "SMUD Pzr Heater Radiant Heat Energy Distribution", Rancho Seco.
- 3. 32-1167609-00, " Heater Element Temperature at Closure Region", Rancho Seco.
- 4. 32-1167603-01, "Pzr Shell Temperature .Due to Radiant Heating", Rancho Seco.
- 5. B&W Technical Manual NPGD-TM-596, Rev. H, July, 1985, ANSYS, Rev. 4, Swanson Analysis Systems Inc., Houston, PA.
- 6. B&W Drawing 135483E-7, " Pressurizer List of Materials",
Rancho Seco.
- 7. B&W Drawing 135489E-4, " Heater Belt Details", Rancho Seco.
- 8. 51-1155656-00, " Standard Correlation for Natural Convection", No specific contract.
- 9. ASME Boiler and Pressurizer Vessel Code,Section III, 1974 Edition with Addenda through Summer 1975.
- 10. ASME Boiler and Pressure Vessel Code, Section III, Appendices, 1980 Edition with no Addenda (for material properties only).
- 11. " Pressurizer Shell Analysis", Design Report No. 7,
" Pressurizer Stress Report", NPD Microfilm Roll No. 80-47, Rancho Seco.
- 12. " Rationale for a Standard on the Requalification of Nuclear Class 1 Pressure-Boundary Components," EPRI NP-1921, Project 1756-1, Final Report, October 1981.
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- 8. THERMAL CALCULATIONS DISCUSSION This section contains the calculations for determining the shell wall thermal gradients. The associated stresses and fatigue effects of these thermal gradients are calculated in Appendix A.
An axisymmetric model of the pressurizer shell adjacent to the heater bundle opening was constructed. This model is shown in Figure 1. The model includes 12.5" thick base metal forging and the 3/16" nominal thickness stainless steel cladding on the interior surfaces.
The attached microfiche of the computer run contains the nodal locations, element definitions and appropriate material properties from 100 0F to 1500 0F.
The loads are applied to the model in the form of nodal heat flow on a per radian basis.
There are two sources of heat flow into the model: 1) radiant heat from the " active" heated length of the element (heater " active" length begins near inside shell surface and extends essentially to opposite interior surface, i.e.,
approx. shell I.D.) and 2) radiant heat from heater sheath adjacent to shell wall thickness (this section of heater is
" inactive" but has temperature rise due to axial conduction from active section).
4 The heat flux resulting from the " active" and " inactive" sections of the heaters are applied to the appropriate surface nodes. The corner region of the model receives heat J
from both sources, i j
i Prepared By Date 9 7 Reviewed By d Date 87 Page
Babcock & Wilcox B&W Doc. No. 32-1167974-01 The radiant heat from the " active" length of the heater bundle is absorbed by the shell as determined from the distribution of Reference 2. The distribution grid 'in Reference 2 is relative to the bundle centerline.
The evaluation herein uses the Ref. 2 relative distribution to determine " heat absorbed per node" from the upper bundle and, separately, the " heat absorbed per node" from 1/4 of the middle bundle. The two contributing values are then added to obtain the total heat absorbed per node. The resulting values vs. " distance from bundle opening centerline" are shown in Figure 2.
The radiant heat from the " inactive" length of the heater bundle is absorbed by the shell penetration surface. The temperature distribution of the " inactive" length of heater is determined in Reference 3. To approximate the heat absorbed by the penetration surface, the heat output of a 1" long section of " active" bundle (39 heaters) was multiplied by the ratio of:
temperature of 1" length of " inactive" portion R =------------------------------------------------------- .
estimated temperature of the " active" portion (2200 0F)
The resulting values are then multiplied by the ratio of: l
" emitting area" (bundle outside surface) J R = --------------------------------------------- l
" absorbing area" (penetration surface) j i
The resulting values are plotted vs. " distance from inside I corner" in Figure 3. The associated " inactive" heater shcath temperature is also included in Figure 3.
Prepared By Date k'N' 7 j Reviewed By #4d Date 87 Page 4 E
, . i Babcock & Wilcox B&W Doc. No. 32-1167974-01 The heat flux intensity in " Btu /hr/in2 " is determined at each inside surface (shell and penetration) node. It is then multiplied by the " area per radian" attributed to the node (normally, " node radius x 1/2 length of each adjacent element"). It should be noted that the three corner region nodes absorb heat from both the " active" heater length and the " inactive" heat length. The " heat flow / radian" is put into the computer code.
The cyclic operation of the heaters is evaluated by varying the heat flux in accordance with recorded data of Reference
- 1. The cycles are essentially as shown in Figure 4.
The results of the computer run indicate that the maximum base metal temperature occurs at node 19 (corner base metal node). The temperature vs. time plots for this node is L shown in Figure 5.
For information, the temperature vs. time curve for the maximum temperature node for the cladding is shown in Figure 6.
Additionally, for reference only, a temperature vs. time plot for the node at the inside shell surface remote from f '
the opening is provided in Figure 7.
As an aid in visualizing the conduction of the heat throughout the model, Figure 8 provides a temperature i l
contour plot at the transient time of maximum base metal j (and cladding) temperature. .
)
The computer output microfiche of the computer run is located on page 16 of this document. i l
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Prepared By !/- >// '/ # 2' Date Reviewed By [ __ Date 9 8/8 7 Page No
a Babcock & Wilcox B&W Doc. No. 32-1167974-01 APPENDIJl TEERMAL STRESS ANALYSIS l
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4 Babcock & Wilcox B&W Doc. No. 32-1167974-_01 l
Purcose .i
~
This appendix provides the calcu' lated thermal gradient stresses (as a result of the heater bundle uncover / recover effects) in the !
vicinity of the upper. heater sbt.ndle shell opening. The "ANSYS" thermal finite element model of Section 8.0 (page 8) is revised for use in performing thie stress evaluation. The structural I i
finite element model does not include the effects of pressure, the heater bundle diaphragm plate, heater bundle cover or the 4
closure studs (See figure 9) for the following reasons. /
- 1) Only the temperature induced stresses are calculated. The pressure during the event v&x nlmost constant and very low f (approximately 40 psi) and a.9 rouch is not significant. )
- 2) The heater bundle diaphragm plate and' seal weld are not I modeled. The seal weld and diaphragm plate were removed and replaced with the installation of the new upper heater t bundle. The thermal movements of the diaphragm plate do not 1 A
produce any significant , structural loacings on the shell. j These thermal motions primarily affect the intensity of stress in the seal weld itself. Since, the ceal wnld was ;
replaced, calculation of these stresses is not; ,reepired.
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- 3) The thermal motions of the shell surface at the stud end location (as can be seen from the ANSYS stress analysis !
results for the iterations analyzed, see microfiche on page
- 16) are small. As a result, no significant stessses would be developed in the studs from the thermal discontinuity situation. Additionally, since the stud will increase in temperature at a faster rate than the ' cover plate (stud is heated by conduction from the shell) the studs will tend to lose a portion of their original preload. For these reasons the stud stresses are not calculated.
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i Babcock & Wilcox liGW Doc. No. 32-1167974-01
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Babcock & Wilcox B&W Doc. No. 32-1167974-01
- 4) 'The thermal discontinuity stresses at the juncture of the thick heater bundle forging and the thin shell are ~ not calculated. The analysis herein demonstrates that the stresses and fatigue usage factors in the significant areas (areas directly affected by the thermal radiation from the heater _ bundle) are very small. The areas remote from the high temperature regions will be subjected to even smaller affects.. Thus, it is obvious that the heater event had a negligible effect on the thick-to-thin shell juncture.
The following transient times (see Figure 10) were chosen for analysis. These time points represent the range of thermal effects from the heater event transient without accounting for the presence of the fluctuating level of the pressurizer bulk i fluid. By not accounting for the fluid, the maximum metal temperatures are assured. The potential effects of the pressurizer fluid are addressed in the subsequent pages.
Time (hrs) Lead Stee Iteration Microfiche Ref.
O.1167 1,20 20 MCHT 1.450 2,50 31 MMAK 1.583 3,20 45 MMBG 2.421 7,30 101 MMAP 3.767 16,30 216 MMAV 3.883 17,20 230 MMAZ The stress results from the above iterations are tabulated in Appendix B. The computer program input and output are on ]
I microfiche and attached to page 16.
)
l Prepared By - Date /
Reviewed By Date f/A//87 Page [
1 Babcock & Wilcox B&W Doc. No. 32-1167974-01 Water-Slan Evaluation Discussion An evaluation was performed to determine the effect of a rapid l rise in pressurizer water level during the event. To maximize l the effect, the initial temperatures for the heater bundle shell opening region will be taken at the end of the heater actuation interval in which the maximum shell temperature was experienced.
This corresponds to load step 18, iteration 10 (time 3.946 hrs) as shown on Figure 10.
For the hypothesized water-slap condition, the hot cladding surface elements will instantaneously be subjected to a bulk fluid water temperature of 200F. This will grossly simulate the effect of a rising pressurizer water level that is totally covering the hot surface which has been conservatively calculated !
to be approximately 1000F.
Calculations The corresponding thermal heat transfer (film) coefficients (required as "ANSYS" input) will be calculated assuming a natural convection flow regime since no significant flow exists in this region.
The film coefficients are calculated using a standard heat transfer correlation, Reference 8. The formula used for the pressurizer inside vertical surface is a correlation for a surface hotter than the ambient fluid:
h = [ delta T (L) 0. 5 0.
3 4 K7 Kg Kg BTU /HR-FT2 .7 Prepared By b Date ,[ '
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Babcock & Wilcox B&W Doc. No. 32-1167974-01
-Where; delta T = Twall - Tfluid = Twall - 200F L = Height of vertical plate (assumed to be a 1 ft.
1 I
rise in water level)
K7 = 0.0246 K8 = is.from Figure NC3, Reference 8 l
Kg = is from Figure NC4, Reference 8 Tm = (Twall + Tfluid)/2 l
J CLAD INSIDE VERTICAL SURFACE INITIAL TEMPERATURES i
Nede Teme.(F) delta T (F) T, Kq Ko h ,
l
, 1 981 781 590 0.865 5000 1528 j 3 813 613 506 0.81 4500 1168 l 4 691 491 446 0.81 4050 962 l \
5 627 427 414 3850 876 0.82 1 l 6 584 384 392 0.827 3650 803 10 475 275 338 0.86 3050 610 11 458 258 328 0.87 2950 582 15 417 217 309 0.89 2750 518 16 413 213 307 0.89 2750 514 The nodal film coefficients will be converted to surface coefficients typical for input to the ANSYS thermal model.
Surface film coefficient = hs = ((hl + h 2)/2](R m/144)
Where, Rm= (R1 + R2 )/2 hs = BTU /HR-IN 2-F/ RAD h1 = surface node 1 film coefficient h2 = surface node 2 film coefficient R1 = radius to surface node 1 R2 = radius to surface node 2 The surface coefficients are tabulated on the next page.
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Babcock & Wilcox B&W Doc. No. 32-1167974-01 Element Surface lu Rm 1-3 102 10.885 3-4 94 12.75 4-5 94 14.75 5-6 98 16.75 10-11 125 30.25 15-16 175 48.75 The element surfaces 6-7, 7-8, 8-9, 9-10, 11-12, 12-13, 13-14, 14-15 and 16-17 were interpolated from the above data for input to the ANSYS thermal model.
HEATER BUNDLE SHELL OPENING INSIDE SURFACE FIEM COEFFICIENTS The heat transfer film coefficient formula for a horizontal surface (results in larger h value than use of the vertical surface equation) hot side up is:
h = [ delta T]1/3 g46 K6 where; delta T = Tw - Tg = Tw - 200F K46 = 0.17 i K6= (See figure NC 2, Ref. 8) h = 0.17 K6 [ delta T]l/3 j 1
Shell Ocenina Film Coefficients l Node Ti delta T Tm K6 h (F) (F) (F) BTU /HR-I!T2 -
F/ RAD) 1 981 781 591 1010 105.0 35 774 574 487 880 82.6 52 607 407 404 785 65.7 69 514 314 357 730 56.0 86 473 273 337 705 51.6 l 103 451 251 326 685 48.8 120 444 244 322 680 48.0 Prepared By f/ Date f/Y/87
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Reviewed By _ dV Date 4/87 Page M
Babcock'& Wilpox B&W' Doc. No. 32-1167974-01 The average surface film coefficient is:
hs " [(hl + h 2)/2)
Where, hs.= BTU /HR-IN 2-F/ RAD h1 = surface node 1 film coefficient h2 = surface node 2 film coefficient The film coefficients for the surface elements are tabulated in the following table.
Element Surface Film Coefficient 17 94 33 74 49 61 65 54 l n.1 50 97 48 l
These film coefficients were input to the ANSYS thermal model with an initial temperature corresponding to load step 18, !
iteration 10, time 3.946 hours0.0109 days <br />0.263 hours <br />0.00156 weeks <br />3.59953e-4 months <br /> (See figure 10). The ANSYS 1
thermal output for the first 18 load steps is on computer microfiche MCAP (See page 16). The ANSYS thermal output for the l
water-slap condition is on computer microfiche MIIH (See page j 16). l The first three thermal iterations from the computer output were f i
input to the structural analysis model. These iterations contain the maximum thermal changes in the cladding temperature as a result of the hypothesized pressurizer water level increase. The iterations are tablulated as:
1 J
l TIME (HRS) ITf2ATIQE MICROFICHE REF.
3.9477 1C MIJC {
3.9493 2C MLCP 3.9510 3C MDEV j l
1 Prepared By Date f/ 87
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Reviewed By N Date 9/M #7 Page $ l 1
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. I Babcock & Wilcox B&W Doc. No. 32-1167974-01 The stress results from these iterations represent the maximum j cooling effect on the cladding and are tabulated in Appendix B.
l The computer input and output are on microfiche and attached to page 16.
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Babcock & Wilcox B&W Doc. No. 32-1167974-01 1
APPENDIX B FATIGUE ANALYSIS Prepared By Date 7/V 87 Reviewed By Id Date [ /4/8/ Page.I /
Babcoc' & Wilcox B&W Doc. No. 32-1167974-01 Pressur?.zer Fatierue Analysis B.1 Introduction A conservative method will be used to determine the shell thermal gradient peak surface stresses and associated fatigue usage factor. This method will account for possible localized " hot spot" stresses at the corner of the heater bundle shell opening.
B.2 Assumptions
- 1. The through-wall thermal gradient will be analyzed assuming that the gradient is uniform around the circumference of the pressurizer heater bundle shell opening.
- 2. To account for localized peak stresses at the inside corner opening, the actual cladding nodal thermal stresses (from node 1, element 2) will be used in the fatigue evaluation of the base metal. l
- 3. The ASME Code methodology for fatigue analysis will be )
followed even though the calculated temperature of the base metal inside corner node is greater than 700F, (the upper limit for the ASME Code specified fatigue curves). Reference 12, page 4-15 gives a suggested guideline for evaluating ferritic materials when the material is subjected to temperatures above 700F for a short period of time. These guidelines will be used in i fatigue analysis for the base metal.
I Prepared By / Date I* # 7 Reviewed By MM Date 9 / Page E
'Habcock & Wilcox B&W Doc. No. 32-1167974-01 B.3 Discussion i The ASME Code,Section III does not give allowable stress values for SA 516 GR 70 above a temperature of 700F. However, yield and tensile strength values are given in Section III. for higher temperatures. Reference 12, page 4-15 suggests that the l allowable stress intensity, Sm (to be used in evaluating primary stresses which occur at temperatures higher than the 700F value permitted by Section III for ferritic materials) be calculated as the lower of: l l
l 1/3 of the ultimate tensile strength, or 2/3 of the yield strength at temperature j using the properties tabulated in Tables I-3.1 and I-2.1, respectively, of Section III when the time at a given temperature is less than that tabulated in the following: )
Metal Maximum Time at
{
Temperature (OF) Metal Temperature (hoursi 750 10000 800 1000 850 100 900 10 950 1 A review of the calculated thermal results shows that the above limits for time at temperature were not exceeded.
The maximum base metal temperature calculated in Section 8.0 is 919F. Therefore, the allowable stress intensity for SA 516 GR 70 l at 919F may be calculated as follows:
Prepared By Date 2//F
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Babcock & Wilcox B&W Doc. No. 32-3167974-01 Table T-3.1 Table I-2.1 T = 950F Su = 46.20 ksi Sy = 23.20 ksi T = 919F Su = 49.68 ksi Sy = 23.74 ksi T = 900F Su = 52.00 ksi Sy = 24.10 ksi Thus, Sm equals the lesser of 1/3 Su or 2/3 Sy 1/3 Su = 1/3 (49.68) = 16.6 ksi 2/3 Sy = 2/3 (23.74) = 15.8 ksi Use Sm = 15.8 ksi 0 920F From Figure NB-3222-1, Note (1), Ref. 9, the value for S m may be taken as the average value of S m at the highest and lowest temperatures of the metal during the transient.
Therefore; Sm 0 920F = 15.8 ksi Sm 0 100F = 23.3 kai Sm= (15.8 + 23.3)/2 = 19.55 ksi 3Sm = 3(19.55) = 58.65 ksi Prepared By [ --- Date k N87 Reviewed By &# Date f Af/ #7 Page 8 O
. -. l Babcock & Wilcox B&W Doc. No. 32-1167974-01 Primary elus Secondary Stress Intensity i
I Figure 11 shows the stress lines through the heater bundle shell l opening area that were analyzed with the ANSYS structural model.
Only.the base metal elements are used since the design code, Ref.
9, does not allow strength of be attributed to cladding. The results of all six stress lines are on the computer microfiche listing. However, only lines 1, 2 & 3 are tabulated since they ,
l are the significant areas of interest.
An additional stress intensity is tabulated from the Pressurizer Stress Report, Ref. 11, to account for stresses due to normal Heatup and Cooldown. The addition of this stress will determine the new maximum ' total primary plus secondary stress intensity range for this area of the pressurizer. The stresses from Segment number 23, page B-9-24 of Ref. 11 are tabulated as typical of the heater bundle opening area.
Stress Classification Line 1 Primary plus Secondary Stress Intensity Range S12 S23 S31 Microfiche Iter (ksii (ksii (ksii Reference 230 6.3 57.2 -63.5 MMAZ 216 2.4 24.6 -27.0 MMAV 101 6.0 55.1 -61.1 MMAP 45 4.6 40.7 -45.2 MMBG 31 0.2 2.6 - 2.8 MMAK 20 3 ., 9 33.8 -37.7 MCHT 1C 5.6 50.6 -56.4 MIJC 2C 4.4 42.8 -47.3 MLCP 3C 3.3 35.9 -39.3 MDEV S23 0.9 3.3 6.5 Ref, 11 Range 6.3 57.2 70.0 Prepared By /i Date N[ '
Reviewed By M Date f /87 Page $/
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Babcock & Wilcox B&W Doc. No. 32-1167974-01 Etress Classification Line 2 }
Primary plus Secondary Stress Intensity Range i
S12 S23 S31 Microfiche Iter (ksii (ksii (ksi) Reference l
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230 5.9 59.3 -65.2 MMAZ !
216 1.8 26.0 -27.8 MMAV i
101 5.8 56.7 -62.5 MMAP j 45 4.6 41.6 -46.3 MMBG l 31 0.2 2.8 - 2.9 MMAK 20 4.0 34.5 -38.5 MCHT 1C 4.9 52.6 -57.5 MIJC 2C 3.2 44.6 -47.8 MLCP 3C 1.8 37.4 -39.3 MDEV S23 0.9 3.3 6.5 Ref. 11 Range 5.9 59.3 71.7 Stress Classification Line 3 Primary plus Secondary Stress Intensity Range S12 S23 S31 Microfiche Iter (ksi) (ksi) (ksi) Referenga 230 6.2 48.1 -54.3 MMAZ 216 3.0 24.1 -27.0 MMAV 101 6.4 45.5 -51.8 MMAP 45 4.3 31.9 -36.2 MMBG 31 0.3 2.6 - 2.8 MMAK 20 3.5 25.8 -29.3 MCHT 1C 6.8 43.4 -50.2 MIJC 2C 5.6 40.2 -45.7 MLCP 3C 4.6 37.1 -41.7 MDEV S23 0.9 3.3 6.5 Ref. 11 Range 6.8 48.1 60.8 Prepared By [/ - Date 0 // ~'/ #
Reviewed By 8M Date f4 / Page
Babcock & Wilcox B&W Doc. No. 32-1167974-01 The maximum primary plus secondary stress intensity range considered the initial. effects of the Water-Slap stress condition. The long term cooling would not have a significant effect on the maximum range of stress since 'the cooling only i serves to bring the entire thickness to an isothermal condition.
I Also, before any long - term cooling could be accomplished the heaters were again reactivated. The maximum stress range occurs ,
for. stress line 2:
t S= 71.7 ksi > 3 Sm = 58.65 ksi The maximum range of primary plus secondary stress may exceed the allowable stress if the requirements of NB-3228.3 (Simplified Elastic Plastic Analysis) are satisfied. These requirements are satisfied in the fatigue analysis section of this report.
Therefore, it is concluded that this primary plus secondary stress intensity range is acceptable.
PEAK STRESS CALCULATION The significant tensile stresses generated as a result of the hypothetical Water-Slap condition are almost completely contained within the clad thickness. Therefore, the clad peak stresses calculated will be used in the fatigue calculations for the base
)
metal.
)i The highest thermal surface stresses occur at the inside corner l of the heater bundle shell opening. The stress lines analyzed )
pass through the element 2 surface. Thus, the finite element model node 1 peak stress from element 2 will be used in the fatigue analysis.
Iteration 230 was recalculated (microfiche reference MBEI) by rerunning the structural model with the cladding elements active l
i Prepared By //- ' Date bY# T I Reviewed By Date f 4 '/ Page J3
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e Babcock & Wilcox B&W Doc. No. 32-1167974-01 (designated as iteration 230C) . This allowed the calculation of i actual stresses on the clad surface. The following is a tabulation of the clad peak stresses from the significant iterations analyzed (Iteration 230 was the only iteration of significance with regard to peak thermal stresses from the heater radiation effects). The principal stresses Si, S2 and S3 were calculated from the element 2, node 1 stresses for iterations 1C, 2C, 3C and 230C. The principal stresses and stress intensities are tabulated in the following table.
PEAK STRESS TABULATIOR Iter S1 S2 S3 S12 S23 S31 1C -1.7 115.4 47.0 -117.1 68.4 48.7 2C -1.7 100.2 54.5 -101.9 45.7 56.2 i 3C -1.0 79.2 50.0 -80.2 29.2 51.0 230C -60.7 15.5 -180.9 -76.2 196.4 -120.2 Range 117.1 196.4 176.2 NOTE: All stresses are in ksi.
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Babcock & Wilcox B&W Doc. No. 32-1167974-01 FATIGUE USAGE FACTOR CALCULATION l The fatigue usage factor will be . calculated according to NB-3228.3 of Ref. 9. The fatigue strength reduction factor K e equals:
Ke = 1.0 +-[(1.0 - n)/n(m - 1.0)][(S n/3Sm) - 1.0]
where; n = 0.2 NB-3228.3, Ref. 9 m = 3.0 NB-3228.3, Ref. 9 Sn = 71.7 ksi page 33 3Sm = 58.65 ksi page 33 Ke = 1.445 The alternating stress intensity equals; Salt = 0.5 (K )e (Sp)(Ecurve/Eused) where; Sp = 196.4'ksi page 34 Ecurve = 30 E6 Figure I-9.1, Ref. 9-Eused = (E70F + Egigy)/2
= (29.9E6 + 24.2E6)/2 psi Ref. 10
= 27.05E6 psi Salt = 0. 5 (1. 4 4 5) (19 6. 4 ) (3 0. 0E6/27. 05E6)
= 157.4 ksi i
From Figure I-9.1 of Ref. 9, the allowed number of fatigue cycles is 190. The required number of fatigue cycles is 19 as can be ;
seen from the cyclic paterns in Figure 6. Therefore, the fatigue usage factor associated with this event is:
)
U = 19/190 = 0.1 The usage factor from the pressurizer stress report associated I with the pressurizer heater bundle shell opening area is 0.075.
Therefore, the revised cumulative fatigue usage factor for the pressurizer heater bundle shell opening is 0.18 which is much less than 1.0, the allowable fatigue usage factor.
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Babcock & Wilcox B&W Doc. No. 32-1167974 1 APPENDIX C i Potential for Clad Crackina j
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]SaLbcock & Wilcox B&W Doc. No. 32-1167974-01 1 Purrose The purpose of this appendix is to provide an assessment of the potential for conditions to exist that could precipitate the ]
creation of, and subsequent effects of, cracks in the cladding on the inner surface of the pressurizer as a result of the Rancho )
Seco pressurizer heater bundle uncover / recover event of November 21, 1986.
Discussion i
This discussion is limited to shell clad areas near the upper i'
heater bundle shell opening. Due to the fact that the cladding in this area was raised to the highest temperature of any point on the pressurizer.
The pressurizer was constructed to the ASME Code,Section III, 1965 Edition with Addenda through Summer 1967. The clad surface, as manufactured, received and passed an ultrasonic test (UT) for bonding and a dye penetrant test (PT) for surface defects. Thus, the original surface was free of detectable cracks.
For a crack to initiate in the cladding during operation either;
- 1) a cyclic fatigue failure must occur (the actual fatigue usage factor must be greater than 1.0), or
- 2) a brittle fracture must occur.
Faticue Usace Factor Consideration l During normal operation, the stresses in the vicinity of the heater bundle opening in the shell are small. The existant stresses are due to internal pressure and a steady state thermal gradient. From Reference 2, the maximum normal operating stress intensity at the heater bundle opening corner (highest stress for Prepared By b Date ((" M'7 Reviewed By M Date M/ 87 Page
Babcock & Wilcox B&W Doc. No. 32-1167974-01 this vicinity) is less than 11.0 ksi. Also, the total number of significant operational cycles is small. (The pressurizer heater bundle opening area basically experiences only one significant operational transient --
'Heatup and Cooldown' for 240 cycles).
This transient does not produce high stresses with respect to fatigue. The vessel has been analyzed to the ASME Code requirements and the fatigue usage factor due to normal operation is very small (0.075 for 40 year operation) . Thus, the -Rancho ;
Seco pressurizer is not expected to develop any clad cracking ,
1 during normal operation.
For the heater bundle uncover / recover event at Rancho Seco, significant thermal gradient effects in the vicinity of the i corner radius of the heater bundle opening were calculated.
Surface temperatures were conservatively calculated to be a maximum of 1002F for the cladding corner. These temperatures generated a high compressive stress in the shell. The only significant tensile stresses would occur when the relatively cool bulk fluid pressurizer level hypothetically raised to again cover the ' hot' cladding. Thus, providing rapid cooling. )
For this hypothesized Water-Slap condition, the hot cladding l surface elements will instantaneously be subjected to a bulk i fluid water temperature of 200F. This will grossly simulate the effect of a rising pressurizer water level that is totally l covering the hot surface. The initial cladding temperature is taken at the end of the heater activation time interval !
I containing the maximum shell temperature. From the analysis '
presented in Appendix A, this corresponds to load step 18, ,
iteration 10 (time 3.946 hours0.0109 days <br />0.263 hours <br />0.00156 weeks <br />3.59953e-4 months <br />). The corresponding thermal heat I transfer (film) coefficients (required as "ANSYS" input) were I calculated assuming a natural convection flow regime since no l significant flow exists in this region. I 1
i Prepared By // Date
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\ . l Babcock & Wilcox B&W Doc. No. 32-1167974-01 1 I
During the hypothesized Water-Slap condition the cladding surface j i
elements will experience high tensile surface stresses ]
(approximately 100 ksi) in excess of the clad material yield strength. The cyclic fatigue analysis of the Water-Slap stresses is very conservative. If the water level actually moved up and
]
down during the event, then the initial shell temperature for each heater activation period would have been much less than those calculated in Appendix A. Therefore, the corresponding i final temperatures (and stresses) would have been less.
I However, to provide consideration of the possible material damage from the event on the fatigue life of the pressurizer, a fatigue strength reduction factor was calculated taking into account these Water-Slap stresses. This factor takes into account the elastically calculated stresses (beyond the yield strength) along with their cyclic effects in the fatigue usage factor calculation. The fatigue usage factor for the event considering all thermal loading cases was 0.1, page 35, Also, it should be noted that while the surface cladding tensile stresses are above yield, the base metal stresses are very much below the material yield stress (i.e., maximum tensile base metal stress is less than 10 ksi).
Therefore, even when the conservatively calculated temperatures in this calculation are used, the calculated usage factor for the heater bundle uncover / recover event is not significant.
Iharmal Crackinc Potential Thermally induced cracks are usually the result of high tensile stresses (i.e. above the material yield strength) with an applied large number of stress cycles. Since stainless steel is a tough material at higher temperatures, a significant number of stress cycles would normally be required to produce a crack.
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Babcock & Wilcox B&W Doc. No. 32-1167974-01 Figure 6 is representative of the cladding surface temperature during the transient. From the figure it is obvious that there are 19 thermal cycles during the event. Of these 19 cycles, only seven have significant thermal ranges. The significant cycle times correspond to 0.0, 1.5, 2.0, 2.3, 3.2, 3.8 and 4.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br />.
Due to plastic deformation, these subsequent stress cycles would not induce stresses much above the yield strength of the material. Also, due to the inherent ductility of the cladding material, these few stress cycles experienced during the heater bundle uncover / recover event would not induce sufficient hardening to cause cracks.
Figure 12, is the ASME Code endurance limit curve for stainless steel material. As can be seen from the figure, the endurance limit is given as a function of the number of stress cycles experienced. The number of stress cycles seen in this evaluation is for too small even to discuss any possible exceedence of the l endurance limit.
In conclusion, based on the fatigue usage factor for the event and che number of possible stress cycles involved, it evident that the potential for creation of cracks in the cladding during the heater bundle uncover / recover event does not exist.
Prepared By [ -
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