ML20235Q730
| ML20235Q730 | |
| Person / Time | |
|---|---|
| Site: | Rancho Seco |
| Issue date: | 09/15/1987 |
| From: | Schaefer R, White L BABCOCK & WILCOX CO. |
| To: | |
| Shared Package | |
| ML20235Q699 | List: |
| References | |
| 51-1167986-01, 51-1167986-1, TAC-64153, NUDOCS 8710070681 | |
| Download: ML20235Q730 (35) | |
Text
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ATTACHMENT 1 Pressurizer Heater Incident Report
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PRESSURIZER HEATER INCIDENT TASK 852 l
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Sacrarento Municipal Utility District
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Box"15830 Sacramento, CA 95852-1830 If.
,BY Babcock & Wilcox Nuclear Power Division j
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Bcx 1260 Lynchburg, VA 24505-0935 1
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EXECUTIVE
SUMMARY
i This report describes k conservative assessment of the expected condition of the Rancho'Seco pressurizer considering the heater j
bundle uncovering event of November 21, 1986.
The objective of this assessment is to evaluate the factors affecting the re-useability of the pressurizer following the sbnormal event.
This assessment was based on an examination of conservative l
transient thermal calculations.
Criteria used to establish the l
f acceptability of the component are ba. sed on 1) engineering judgement considering the original material qualification l
temperatures and stress-relief temperature and 2) fatigue l
calculations according to metal ASME Code, Sectior., III.
Table 6
summarizes the assessment of the maj or pressurizer components.
The maximum base metal at/d cladding temperatures were conservatively determined to be 919F und 1002F respectively.
These temperatures are acceptable based on the established criteria.
Therefore, based on the resulta of the conservative thermal analysis performed, all components axcept the upper and middle heater bundle.1 are acceptable for r9use.
Both of these heater bundles were replaced.
Section 4.0 summarizes the fatigue analysis results, which show that the pressurizer did not suffer any significant fatigue damage as a result of the incident.
Additionally, the calculations in Reference 5 show that the maximum temperature of the heater element at the diaphragm plats is 258F.
This value is significantly less than the no2. mal operating temperature of a hu ter element 'at this location.
Therefore, the electrical connections to the heater elements were not degraded as a result of the heater bundle incident.
In conclusion, the effect of the heater bundle incident on the structural integrity cf the pressurizer it acceptable.
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\\le TABLE OF CONTENTS PAGE 3
9
' EXECUTIVE
SUMMARY
2 TABLE OF CONTENTS 3
i LIST'OF' TABLES 3A LIST OF FIGURES 3A 1.
INTRODUCTION 4
2.
PRESSURIZER TRANSIENT ANALYSIS 5
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2.1 SCOPE 5
2.2 DISCUSSION OF METHODOLOGY 7
'e 2.3 THERMAL ANALYSIS RESULTS 14 3.-
METALLURGICAL EVALUATION
'22 3.1 SCOPE 22 yl.
3.2 RESULTS 22 l
3.3 DISCUSSION 22 1
4.
FATIGUE EVALUATION 25 4.1 PRESSURIZER SHELL 25 4.2 PRESSURIZER HEATER BUNDLE SHELL OPENING 26 5.
CLAD THERMAL EFFECTS 28 6.
REFERENCES 33 x
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LIST OF TABLEE L
TABLE PAGE 1
Heater Bundle Transient 10 2
Assumptions for Determining Heater 11 Activation 3
Conservatism in Assumed Heat Flux Values 14 4
Conservative Assumptions in Pressurizer 16 Shell Thermal Analysis 5
Conservatism Used in Heater Bundle Shell 18 Opening Analysis
'6 Pressurizer Metallurgical Assessment
-24 LIST OF FIGURES FIGURE P_AGI 1
Pressurizer Longitudinal View 6
2 Actual Pressurizer RTD Temperature 8
3 Pressurizer Upper Heater Bundle Heat Flow 9
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Heat Flux vs. Distance from Centerline of 14 Heater Bundle j
5 Pressurizer Shell Regions Analyzed 15 6
Pressurizer Heater Bundle Opening Maximum 20 Temperature j
7 Heater Bundle Opening Thermal contour Plot 21 8
ASME Code Evaluation Limit Curve for 32 Stainless Steel I
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j 1.O INTRODUCTION f
This report describes an assessment of the physical condition of selected pressurizer components and presents recommendations on whether the components can be reused or should be replaced or repaired before repressurization of the pressurizer.
The assessment and recommendations are based on a conservative (i.e.
assumed more severe effects as compared to actual) interpretation of design information and conservative analyses of the transient effects over the period beginning at 3:10 AM and concluding at 9:54 AM on November 21, 1986, combined with summary evaluation of available data from before and after this period.
4
2.0 PRESSURIZER TRANSIENT EVALUATION This evaluation was conducted to provide a reasonable and timely technical basis for the determination of the acceptability of the pressurizer equipment for future operation.
Therefore, the rigorous analytical techniques normally used in the determination of a specific transient and the analysis of pressure boundary components were not used.
The analyses and evaluations performed to assess the metallurgical and functional condition of the pressurizer components (See Figure 1) are based on the data presented in this section.
2.1 Scone The scope of the analysis is limited to providing an assessment of the temperature ~ conditions for the pressurizer components listed in Table 6.
Specifically, the analytical scope is defined as follows:
)
(1)
Establish a
sequence of events on which to base the evaluation.
(2)
Develop a set of compatible analytical parameters consistent with time and cost constraints.
(3)
Develop simplified temperature transient curves.
(4)
Analyze the pressurizer for the temperature transient history during the period from 3:10 AM to 9:54 AM.
(5)
Calculate the revised fatigue usage factors for the pressurizer as a result of the heater bundle uncovery event.
(6)-
Evaluate the potential for clad cracking to have occurred.
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i 2.2 Discussion of MethodolocV Figure 2 shows a plot of the pressurizer RTD probe during the pressurizer heater incident (Time 3:10 AM to 9:54 AM).
Due to a lack of detailed logged data on the actual operating time of the pressurizer heaters, the RTD digital ree.dout was used along with the control room log to determine on/off times for the heaters.
Additionally, the control room log was used to determine the number of heater elements racked out during the incident.
The on-times for the heater transient considered the time required for heatup and cooldown of the heater sheath.
Using this data the transient shown in Figure 3 and tabulated in Table 1 was assembled.
(Note:
Figure 3 and Table 1 list the total heat flow for the upper bundle only).
A detailed discussion of the methodology employed to determine the pressurizer transient is presented in Reference 1.
Table 2 lists the assumptions used in determining the heater bundle element activation time intervals and heat flow from the upper bundle.
These assumptions as a whole provide a very conservative assessment of the actual transient.
The pressurizer heater radiant energy distribution for the shell and upper head surface was formulated.
The formulation assumed radiation of energy from the lateral aurfaces of the heater bundle.
The bundle was treated as a cylindrical heat source and the pressurizer as a black body absorber, a right circular cylinder.
Using this analogy, the calculated net heat flux on the pressurizer shell and head surfaces was calculated.
A computer program was prepared to calculate the required heat flux distribution at any point on the pressurizer inner surface.
A detailed discussion of the method and a listing of the developed computer program is presented in Reference 2.
The pressurizer shell heat flux value calculated represents the effect of the l
upper and middle heater bundles.
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t TABLE 1 PRESSURIZER UPPER HEATER BUNDLE HEAT FLOW DATA TIME ACTIVE HEAT FLOW TIME ACTIVE HEAT FLOW (MIN)
GROUPS (E6 BTU /HR)
(MIN)
GROUPS E6 BTU /HR 0.00 13 1.863 246.00 7
1.003 7.00 13 1.863 249.50 7
1.003 7.00 0
0.000 249.50 0
0.000 87.00 0
0.000 257.00 0
0.000 87.00 13 1.863 257.00 7
1.003 93.50 13 1.863 258.75 7
1.003 93.50 0
0.000 258.75 0
0.000 122.00 0
0.000 269.00 0
0.000 122.00 12 1.720 269.00 7
1.003 126.25 12 1.720 281.00 7
1.003 126.25 0
0.000 281.00 0
0.000 135.00 0
0.000 302.00 0
0.000 135.00 12 1.720 302.00 7
1.003 145.25 12 1.720 308.00 7
1.003 145.25 0
0.000 308.00 0
0.000 153.00 0
0.000 317.00 0
0.000 153.00 12 1.720 317.00 7
1.003 155.00 12 1.720 323.00 7
1.003 155.50 0
0.000 323.00 0
0.000 164.00 0
0.000 330.00 0
0.000 164.00 11 1.576 330.00 7
1.003 167.25 11 1.576 335.75 7
1.003 167.25 0
0.000 335.75 0
0.000 190.00 0
0.000 343.00 0
0.000 190.00 10 1.433 343.00 7
1.003 202.00 10 1.433 345.75 7
1.003 202.00 0
0.000 345.75 0
0.000 208.00 0
0.000 358.00 0
0.000 208.00 9
1.290 358.00 7
1.003 209.75 9
1.290 361.50 7
1.003 209.75 0
0.000 361.50 0
0.000 226.00 0
0.000 370.00 0
0.000 226.00 13 1.863 370.00 7
1.003 233.00 13 1.863 374.00 7
1.003 233.00 9
1.290 374.00 0
0.000 236.75 9
1.290 387.00 0
0.000 236.75 0
0.000 387.00 7
1.003 246.00 0
0.000 390.50 7
1.003 390.50 0
0.000 l
Note: Each group cansists of 3 heater elements or 42 kw.
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TABLE 2 ASSUMPTIONS FOR DETERMINING HEATER ACTIVATIONS (1)
The heater element requires approximately 1.5 minutes to reach 1100F (See figure of Appendix C,
Reference 1,
for predicted thermal response).
It was assumed that at this point the RTD started to digitally register a temperature change.
(2)
The RTD temperature probe records the heater element temperature instantaneously.
(3)
When the heater bundle activation stops, the temperature of the heater element drops.
It was assumed that the RTD probe temperature reading levels off at this point, since a drop in heater radiation can no longer support temperature increases while conduction distributes the metal heat throughout the pressurizer.
(4)
It was assumed that the heaters were deactivated as soon as the RTD temperature probe reading levels off.
(5)
For consistency
- purposes, the heater elements were considered to be on as soon as an RTD temperature increase is recorded.
Similarly, the heater elements were considered to be off as soon as the RTD temperature levels off.
An
- diticM1 1.5 minutes was added to each of these activation time periods to account for the initial heater element heatup delay (Assumption 2).
(6)
The middle heater bundle elements assumed to be exposed were activated during each of the upper bundle activation times.
(7)
The RTD digital readout (see Reference 1) from 8:00 AM to 9:54 AM shows no significant heat input.
However, the total energy input to the thermal analysis considered only the 18 elements of the upper heater bundle racked out (See page 40 of Reference 1) to be off during this time span.
This was considered to be very conservative; since after the incident electrical testing revealed several other elements with high resistance readings and several showing open circuits.
11
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______________________J
Discussion of Methodoloav (Cont'd)
The calculated net heat flux (Figure 4 represents a typical distribution at-the upper. heater bundle centerline) was applied to the computer simulation models of the pressurizer shell and heater-bundle shell opening.
The results are presented in Section 2.3.
Table 3 lists the conservatism inherent in the shell heat flux values used in the thermal analysis.
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TABLE 3 CONSERVATISM IN ASSUMED HEAT FLUX' VALUES (1)
The total heater bundle heat flux is assumed to be concentrated on the outside diameter of the heater bundle.
This assumed distribution concentrates the heat to a
localized area maximizing a local " hot spot" effect on the shell.
(2)
The heat flux value considered the actual location of the middle heater bundle with respect to the upper heater bundle (rotated 40 degrees) and still conservatively considered 25 percent of the elements of the middle heater bundle to be radiating heat on the shell.
This assumption of the percent participation by the middle heater bundle was confirmed with Rancho Seco personnel to be conservative as a result of the inspection performed on the middle bundle (i.e.,
the top heater element was deformed and the next four lower elements were discolored, thus, only a small portion of the middle bundle was ever exposed).
2.3 Thermal Analysis Results Pressurizer Shell The ANSYS finite element computer code was used to perform the thermal analysis of the pressurizer shell (See Reference 3).
A 90 degree section (longitudinal symmetry was assumed) of the pressurizer shell which included one half of the upper heater bundle was modeled (See Figure 5).
The longitudinal dimension included the RTD probe location and accounted for a portion of the water in the bottom head.
Table 4 lists the conservative i
assumptions used in the thermal analysis of the pressurizer shell.
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TABLE 4 CONSERVATIVE ASSUMPTIONS IN PRESSURIZER SHELL THERMAL ANALYSIS
. (1)
The heat flux values used are very conservative.
(2)
The transient times -used in this analysis were a
conservative representation of the actual abnormal event.
Heater "on" times were combined and the'"off" times between them were. eliminated.
-A total of 29.75 minutes of "off" time (soak time) was removed from the analyzed transient.
l (3)
The finite element model boundaries were assumed to: be totally insulated.
(i.a.,
no heat -loss from metal-to' outside environment or back into the gaseous medium - of the pressurizer).
In addition, the heat was confirmed to the material of 102 inch high, 6 inch thick cylinder with only 12 inches of water level.
The heat removing capacity of the 1
actual pressurizer shell is much greater than this.
.(4)
The cladding thickness was not modeled.
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(5)
The total heater bundle heat flux is assumed to be concentrated on the outside diameter of the heater bundle.
This assumed distribution concentrates the heat on localized area maximizing a local " hot spot" effect on the shell.
(6)
A comparison of the temperature at the RTD location to the actual RTD readout indicates that a conservative number'of heater elements are being energized in the thermal analysis.
The maximum localized base metal temperature determined in the pressurizer shell analysis was 700F.
This value is approximately the design temperature (670F) of the pressurizer.
The maximum value was determined to be near the end of the upper bundle in the. horizontal shell section containing the upper heater bundle centerline.
The thermal results at the pressurizer shell computer model at the RTD location indicate that a conservative number of heater elements. is being used in the thermal analysis.
The computer 16
thermal analysis-predicts : significantly higher temperatures at this location than were recorded by the RTD probe during the heater bundle incident.
This conservatism could be reduced in a-future assessment of the pressurizer shell assessment
.The thermal effect ~ of.an individual heater element in direct contact with the shell was evaluated (grown longer due to its increasing _ temperature).
A conservative representation was modeled with the ANSYS finite element computer code.
The.model assumed perfect conduction between the heater tip and the shell with no heat losses to the environment.
The maximum localized j
l-increase in the shall base metal temperature was 144F if the heater element is on 30 minutes.
The shell temperature was assumed to be 600F.
This contact temperature increase is in addition to the shell temperature determined in the Reference 3 analysis.
- However, the contact temperature increase does not.
coincide with the point of maximum shell temperature calculated in Reference 3.
Thus~ no significant effect would have incurred if an individual heater element did contact the shell.
See Reference 6 for details of the evaluation.
Additionally, a calculation was performed to determine the load which could be developed by thermal expansion of the heater bundle elements and if these loads could cause deformation in the pressurizer wall.
The results of the calculation, Reference 8, reveal that the thermal expansion load required to buckle the pressurizer heater element does not produce sufficient force to cause deformation to the pressurizer vessel wall.
]
Uccer Head The results of the pressurizer shell thermal analysis show that no significant temperatures were reached by any portion of the pressurizer upper head.
Thus, no further evaluation of these 17
areas is' required.
See Reference 3 for the thermal distribution in detaili..
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Heater Bundle Shell ODenina The heater bundle shell opening was analyzed. using an axisymmetric thermal model (See, Figure 5).
The finite element computer code ANSYS was used to calculate the-thermal distribution.
Table 5 is a listing of the conservatism used in the analysis of the pressurizer. heater bundle shell opening.
TABLE 5 CONSERVATISM USED IN HEATER BUNDLE SHELL OPENING ANALYSIS (1)
The highest calculated shell heat flux values are in the horizontal plane at the heater bundle center line elevation.
These values were assumed symmetrically over the entire surface of the model.
(2)
The heat flux values used in'- this analysis are the-calculated values shown in Figure - 4 which represent the effect of 1.25 heater bundles.
(3)
A conservative heat flux value was calculated for the unheated portion of the upper heater bundle exposure area inside the opening in the shell.
(4)
No-heat losses' were allowed from the pressurizer model surfaces to the environment.
Therefore, the thermal model was a
very conservative representation of the actual energy being absorbed (especially at the corner node points) by the shell and remaining therein.
The calculated maximum cladding and base metal temperature of the inner corner of the heater bundle opening were calculated to be 1002F and 919F, respectively.
Figure 6 is a plot of the heater f
bundle opening inside corner maximum base metal temperature.
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Figure 7 is a thermal contour plot of the model area'at the time-of. maximum temperature.
The base metal temperature decreases rapidly.in both directions away from the corner node point.
At a
' distance of 1.8 inches the maximum base metal temperature of 919F has decreased to an average temperature of 760F.
This temperature is significantly less than the actual material qualification temperature.
Therefore, based on the following discussion in Section 3.0 and j
the fatigue analysis results presented - in Section 4.0, it is concluded that the pressurizer pressure boundary component parts were not significantly affected by the heater bundle uncovery event.
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PRESSURIZER HEATER BUNDLE 00ENING MAXIMUM TEMPERATURE (Basemetal)
FIGURE 6 I
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FIGURE 7 HEATER BUNDLE OPENING
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fin = tiinimum Temperature Equals 249 F g
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MX=MaximumTemperatureEqugls1002F Clad Temperature = 1002 F Base Metal Temperature = 919 F 21
c 3.0 METALLURGICAL EVALUATION The effects of the
" Higher than Design" temperature on pressurizer materials were assessed based on the calculated temperatures addressed in Section 2.
3.1 Score This evaluation is limited to evaluating temperature effects on the properties of metallic construction materials to assess potential changes in mechanical and fracture toughness properties and corrosion resistance.
The components considered are listed in Table 6.
3.2 Results The metallurgical properties of the pressurizer pressure boundary component materials were not significantly affected by the temperatures to which they were exposed.
Table 6 lists the component materials, and the estimated temperatures to which each component was exposed.
The pressurizer was stress relieved at 1100-1150F for 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> minimum.
The stress relief was performed following weld cladding and structural welding operations.
The postulated temperatures, given in Table 6,
are significantly lower than the stress relief temperature.
For this reason, these j
thermal exposures did not affect the metallurgical condition and the mechanical and fracture toughness properties of the base metal or cladding.
The corrosion and mechanical properties of the heater elements in the upper bundles were significantly affected; however, they have been replaced.
i 3.3 Discussion l
The metallurgical condition of the pressurizer components as presented in Table 6 depends on two primary considerations - (1) l 22 i
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the temperature range and/or' the peak temperature the. material experienced, and (2) the length of time of exposure.
The assessments of material condition was based. on the conservative estimate of the thermal exposure.
Since, the ASME code has no specific guidelines for requalification of pressure vessel that
]
has been subjected to a transient condition : above its design l
temperature, the following methodology was employed.
In general, it was assumed that if the temperature the material experienced t
during the incident was not higher than the fabrication temperatures and that the time of exposure was not excessive,.the
(
material did not suffer metallurgical degradation (e.g.,
pressurizer carbon steel components are stress-relieved at 1100-l 1150F for approximately 5-10 hours followed by furnace cooling at about 15F/h).
I The interior surfaces of the pressurizer near the heater bundle shell opening corner are manually clad with stainless steel..
The first layer is Type 309 and the second layer is Type 308.
The pressurizer stress relief is performed after welding and weld cladding.
The cladding has a microstructure that includes small j
amounts of delta-ferrite that makes it virtually immune to stress corrosion cracking (SCC) following the stress relief heat treatments.
The postulated thermal exposure resulting from the heater bundle. temperature excursion is therefore insignificant with respect to SCC respectively.
Although there is some uncertainty in actual temperature (relative to how much less than the calculated temperature), it is realistic to expect that based on the conservative estimate of probable temperature, the pressurizer pressure boundary component
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parts should not have been degraded as a result of the transient.
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TABLE 6
- J PRESSURIZER MATERIALS,-THERMAL EXPOSURE AND DAMAGE ASSESSMENT Est Incident Probable temp, F/ Time metallurgical i
Commonent Material at teme damace?
Pressurizer SA516, GR-70 700/0.5' Hrs..
No L
Shell Heater Bundle SA508, - CL-1 919(a)/0.5 Hrs.
No Opening Inside Corner Heater Sheath SA213, TY316L' 2000(b)
Yes 1
l Cladding-at Type 309'& 308 1000(c)/0.5: Hrs.-
No Heater Bundle'.
Opening Corner (a)
Localized. value at corner of heater bundle shall opening (see page 15 for discussion)
(b)
Sensitization temperature range is 800-1600F.
(c)
See page 20 for discussion of process.
24
4.O FATIGUE _EyALUATION 4.1 Pressurizer Shell A simplified but conservative method was used to determine the pressurizer shell thermal gradient stresses and associated fatigue usage factors.
The methods used included an assessment of possible localized thermal " hot spot" stresses.
The effect of a rapid rise in pressurizer water level during the event was also considered.
To maximize this effect, the initial temperatures for the shell were taken at the end of the heater actuation interval in which the maximum shell temperature 'was experienced.
For this hypothesized Water-Slap condition, the hot cladding surface elements were instantaneously subjected to a bulk fluid water temperature of 200F.
This grossly simulated the effect of a rising pressurizer water level totally covering the hot shell inside surface.
To account for localized thermal peak stresses at the " hot spot"
- location, the inside surface thermal peak stresses were calculated based on two dimensional rigid constraint equations considering an instantaneous temperature change of the surface.
The condensed version of the equation is:
S = 1.429 (E) (alpha) (Delta T)
The heater bundle uncover event thermal change was based on the maximum shell surface temperature differential experienced during each heater activation interval within the entire heater uncover event transient.
For the hypothesized Water-Slap condition, the maximum temperature change possible during the event was 700F 200F
=
25
500F.
This value used was very conservative since the calculation assumed that the initial temperature through the entire shall thickness was 700F at the initiation of the Water-Slap condition.
- However, since the fatigue effect was not significant in this area of the shell, these stresses were used in lieu of detailed thermal calculations.
The cumulative fatigue usage factor for the event on the pressurizer shell is 0.063.
The maximum cumulative fatigue usage factor in the current stress report, reference 7, for the shell is 0.075.
Therefore, the revised total cumulative fatigue usage factor for the pressurizer shell near the outboard end of the upper heater bundle is 0.14 which is much less than 1.0, the allowable fatigue usage factor.
4.2 Pressurizer Heater Bundle Shell ocenina A conservative method was used to determine the shell thermal gradient peak surface stresses and associated fatigue usage factor.
This method accounted for possible localized " hot spot" stresses at the corner of the heater bundle shell opening.
Six different stress classification lines through the heater l
bundle shell opening area (See Figure 11 of Ref. 4) were analyzed with the ANSYS structural model.
The analysis results are contained in Ref.
4.
The significant tensile stresses generated as a result of the hypothetical Water-Slap condition are almost completely contained within the clad thickness.
These clad peak stresses were calculatad and
- used, very conservatively, in the fatigue calculations for the base metal.
26 j
l
l,,
The fatigue usage factor for the heater bundle shell opening area associated. with this event was 0.10,.
The usage factdr from the pressurizer' stress report. associated with the pressu'izer heate.r r
bundle shall opening area was 0.075.
Therefore, the,. revihed IJ
-cumulative fatigue usage factor for the press',urf. der heater bundle
/
shall opening is 0.10 which is much less than'.l.0, the allowable fatigue usage factor.
In conclusion, even when the conservatively calculated temperatures in References 3 and 4 and discussed in Section 2.0 were used, the calculated pressurizer fatigue ' usage factors due to the heater bundle. uncover / recover event are not significant.
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5.0 CLAD THERMAL EFFECTS This Section provides. an assessment of the potential for conditionsEto have existed that could have precipitated the
/
creation 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.
This
discussion is taken from Appendix C of'Refersnce 4.-
Discussior; i
This discpssion is limited to shell clad areas r, ear the upper heater buridie shell opening.
Due to the fact that the cladding in' this area was raised to the highest temperature of any point f'
on the pressurizer.
i
'l The pressurizer was constructed to the ASME Code,Section III,
{t 1965 Edition with Addenda through Summer 1967.
The clad surface, Las, 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.
l For a crack to initiate in the cladding during operation either; lht.a cyclic. fatigue failure must occur (the actual fatigue usage factor must be greater than 1.0), or
- 2) a brittle fracture must occur.
f3ty,1aue Usece Factor Consideration During normal operation, the stresses in the vicinity of the j
I heater bundle opening in the shell are small.
The existant stresses are due to internal pressure and a steady state thermal l
gradient.
The maximum normal operating stress intensity at the h
heater bundle opening corner (highest stress for this vicinity) j l
is less than 11.0 ksi.
Also, the total number of significant
]
i 28
l l.'. l '.'
It operational cycles is small.
(The pressurizer heater Lundle opening
.'a r e a basically experiences only one significant
'Heatup and Cooldown' for 240 design operational transient cycles).
This transient does not produce high stresses with l
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 during normal operation.
For the heater uncover / recover event at Ranch:.3 Seco, significant thermal gradient effects in the vicinity of the corner radius of the heater bundle opening were calculated.
Surface temperatures were conservatively calculated tc he a maximum of 1002F for'the cladding corner.
These temperatures generated a high compressive stress in the shell.
The only 'sig 2ificant 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 hoc cladding surface elements were instantaneously subjected to.a bulk fluid water temperature of 200F.
This grossly simulated the. etfect of a rising pressurizer water level that total.ly coveru the hot i
surface.
The initial cladding temperature was taken at the end of the heater activation time interval containing the maximum shell temperature.
The corresponding thermal heat transfer j
(film) ccafficients were calculated assuming a natural convection l
I flow regime sinco no significant flow existed in this region.
l During the hypothesized Mater-Slap condition the cladding surface elements will experience high tensile surface stresses (approximately 100 ks.t) in excecs of the clad material yield l
strength.
The cyclic fatigue analysis of the Water-Slap stresses was very conservative.
M the water level actually moved up and 1
29 l
,b 4
~ down during the event, - then the initial shell temperature for each heater activation period would have been much less than those calculated in Reference 4.
Therefore, the corresponding
' final temperatures (and stresses) would have been less.
However, to provide consideration of the possible material damage from the event on the fatigue life of the pressurizer, a fatigue o
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 i
loading cases was 0.1,'page 36, Reference 4.
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 kai).
Therefore, eve.n when the conservatively calculated temperatures j
in the Reference 4 calculation are.used, the calculated usage factor for the heater uncover / recover event is not significant.
I Thermal Crackina Potential Thermally induced cracks are usually the result of high tensile stresses (i.e. above the material yield strength) with an applied J
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.
Figure 6 is representative of the cladding surface temperature fluctuations during the transient.
From the figure it is obvious that there are 19 thermal cycles during the event.
Of these 19 30
4
,,s.
- 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.("b 51ours.
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 uncover / recover event would not induce sufficient hardening to cause cracks.
1 Figure 8,
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 to even discuss exceeding the endurance limit.
In conclusion, based on the fatigue usage factor for the event and the number of possible stress cycles involved, it is evident that the conditions that would be expected to precipitate the creation of cracks in the cladding during the heater uncover / recover event did not exist.
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,,j 6.0. REFERENCES' J. %.
- 2:2 1.
B&W1. Document ~ No. 51-1167607-00, " Heater Element Activation Tinkhbtervals",RanchoSeco.
2.
B&W Document No.
32-1167978-00, "SMUD ' Pressurizer Heater Radiant Energy Distribution", Rancho Seco.
3.
B&W Document No. -
32-1167603-01,
" Pressurizer Shell Temperature Due to Radiant Heating", Rancho Seco.
4.
B&W Document No. 32-1167974-01, " Local Heating Pressurizer Shell Inside Corner", Rancho Seco.
5.
B&W Document No. 32-1167609-00, " Heater Element Temperature at Closure Region", Rancho Seco.
6.
B&W Document No.
32-1167984-00,
" Pressurizer Contact.
Temperature with Heater Element", Rancho Seco.
7.
" Pressurizer Stress Report", B&W NPD Microfilm Roll No. 80-47, Rancho Seco.
8.
B&W Document No. 32-1167616-00, " Pressurizer Heater Element Punch Load", Rancho Seco.
1 I
33 1
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