ML20204G276
| ML20204G276 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 04/13/1983 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| Shared Package | |
| ML20204G258 | List: |
| References | |
| 77-1141514, 77-1141514-00, NUDOCS 8305020404 | |
| Download: ML20204G276 (61) | |
Text
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BABC0CK & WILC0X 177 FUEL ASSEMBLY OWNER'S GROUP SAFE END TASK FORCE REPORT ON GENERIC INVESTIGATION OF HPI/MU N0ZZLE COMPONENT CRACKING B&W Document Number: 77-1141514-00 Prepared for Consumers Power Company
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The Babcock & Wilcox Company Utility Power Generation Division Lynchburg, Virginia 8305020404 830413 PDR ADOCK 05000329 E
TABLE OF CONTENTS PAGE
- 1. 0 EXECUTIVE
SUMMARY
1
- 2. 0 INTRODUCTION 2
2.1 Background
2
- 2. 2 Scope 5
- 2. 3 Results 6
- 2. 4 Organization 7
- 3. 0 COMPILATON OF FACTS 8
3.1 Failure Analyses 8
3.2 Matrix of Facts 11
- 4. 0 REVIEW 0F INDUSTRY EXPERIENCE 14
- 5. 0 CRYSTAL RIVER-3 INSTRUMENTED N0ZZLE DATA EVALUATION 17
- 6. 0 ANALYTICAL INVESTIGATION OF EXISTING DESIGN 19
- 7. 0 POSSIBLE ROOT AND CONTRIBUTORY CAUSES 21
- 8. 0 PROBABLE FAILURE SCENARIO 25
- 9. 0 MODIFIED THERMAL SLEEVE DESIGN 27 9.1 Conceptual Designs 27 f
- 9. 2 Design Improvements 28 9
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TABLE OF CONTENTS (cont'd)
PAGE 10.0 MAKEUP SYSTEM OPERATING CONDITIONS 30 3
11.0 AUGMENTED' INSERVICE INSPECTION PLAN 31 12.0 JUSTIFICATION OF LONG TERM OPERATION 33
13.0 CONCLUSION
S 35 14.0 RECOMMENDATIONS 36
15.0 REFERENCES
38 4
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LIST OF FIGURES Title 1.
Typical Elevation View of Reactor Coolant System Arrangement Showing Location of HPI Nozzle 2.
Typical Plan View of Reactor Coolant System Arrangement Showing Location of HPI Nozzle 3
Typical HPI and HPI/MU Nozzle 4
Typical Layout of HPI and HPI/MU Line 5.
Safe-End Task Force Action Plan 6.
Instrumentation Arrangement ac Crystal River-3 7.
" Goodness of Roll" Results 8.
Hard Rolled HPI/MU Nozzle Concept 9.
Integral HPI/MU Nozzle Concept 10.
Flanged HPI/MU Nozzle Concept
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- 1. 0 EXECUTIVE
SUMMARY
The purpose of this report is to summarize the Safe-End Task Force's involvement in the high pressure injection / makeup (HPI/MU) nozzle cracking problems which affected Crystal River-3, Oconee-3, Oconee-2, Arkansas Nuclear One-1, and Rancho Seco.
Formed by the Babcock & Wilcox (B&W) 177 Fuel Assembly Owner's Group, the Task Force has identified the root cause of the failures, recommended modifications to eliminate future failures, and identified studies to support these modifications on a long term basis.
Site inspections conducted in February-April 1982 indicated that both the HPI only nozzles and the double-duty HPI/MU nozzles were affected.
- Loose, out-of-place, and cracked thermal sleeves were observed in 6 of the HPI only nozzles, while 4 of the double-duty nozzles also contained cracked safe-ends.
Failure analyses indicated that the cracks were initiated on the inside diameter and were propagated by thermal fatigue.
The cracked safe-end at Crystal River also contained mechanically initiated outside diameter cracking which appeared to be unrelated.
Previous inspections at two plants (Davis Besse-1 and Three Mile Island-2) under construction revealed that one of the Davis Besse sleeves was loose.
All four sleeves were subsequently re-rolled at Davis Besse (hard rolled, instead of contact expanded as originally specified).
Recent inspections at Midland have also shown that gaps may be present between the thermal sleeve and safe-end in the contact expanded joint.
These findings along with stress analysis have implicated insufficient contact expansion of the thermal sleeves as the most probable root cause of the failures.
With this in mind, B&W has recommended modifications to the design, operation and inspection of the HPI/MU nozzles.
A hard rolled thermal sleeve design has been developed which helps prevent thermal shock to the
,1 nozzle assembly and helps reduce flow induced vibrations more effectively.
An increase in minimum continuous makeup flow has been suggested to help prevent thermal stratification in the MU line and more effectively cool the safe-end.
An inservice inspection (ISI) plan has also been developed to f
provide a means of early problem detection,
2.0 INTRODUCTION
On January 24, 1982, normal monitoring of the Crystal River-3 reactor coolant system (RCS) indicated an unexplained loss of coolant.
After an orderly plant shutdown, the double duty high pressure injection / makeup (HPI/MU) nozzle check valve-43 was identified as the source.
The valve, the valve to the safe-end weld, the safe-end, and the thermal sleeve were cracked as a result of thermal and/or mechanical fatigue.
Inspections at other Babcock & Wilcox (B&W) operating plants indicated similar types of cracking, but to a lesser extent.
As a result, the Safe-End Task Force (SETF) was formed to compile the pertinent facts and to determine a most probable root cause for the failures.
Since the failures were apparently generic in nature, the following report was compiled describing the Task Force's investigation.
Specifically, the relevant facts and most probable failure scenario are presented, as well as recommended modifications to the thermal sleeve design, makeup system operating conditons and inservice inspection (ISI) plan.
2.1 Background
On the 145,177 and 205 fuel assembly (FA) plants, four HPI/MU nozzles (one per cold leg) are used to:
(1) provide a coolant source for emergency core cooling, and (2) supply normal makeup (purification flow) to the primary system (see Figures 1 and 2).
In general, one or two of the nozzles are used for both HPI and MU, while the remaining nozzles are used for HPI alone.
The incorporation of a thermal sleeve into a nozzle assembly is a common practice in the nuclear industry (See Figure 3).
The function of the thermal sleeve is to provide a thermal barrier between the cold HPI/MU fluid and the hot high pressure injection nozzle.
This helps prevent thermal shock and fatigue of the nozzle.
The purpose of the safe-end is to make the field weld east sr (pipe to safe-end) by allowing similar metals to be welded.
..le dissimilar metal weld between the safe-end and the nozzle can then be made under controlled conditions in the vendor's shop.
The use of the safe-end also eliminates the need to do any post-weld heat treating in the field. _ _ _ _
While monitoring the Crystal River-3 RCS for unidentified leakage, a notable increase was observed on January 24, 1982.
On January 25, a further increase in leakage was observed and the unit was subsequently placed in Hot Standby on January 28.
The check valve (MUV-43) to safe-end weld on the double duty HPI/MU nozzle contained a thru-wall circumferential crack which caused the leak.
Following removal of the valve, visual inspection of the safe-end and thermal sleeve revealed that both components were cracked and worn (see Figure 3).
Inspection of the other three HPI nozzles indicated that no cracking or wear was present, and no sleeve movement had occurred.
Following the incident at Crystal River-3, letters were issued to each of the B&W 177 FA utilities informing them of the discoveries at Crystal River-3.
Inspections were performed at all 177 FA plants to determine whether the problem was site-specific, or generic in nature.
Oconee-1 was shutdown for refueling when Duke Power received B&W's correspondence.
Consequently, Oconee-1 was the first unit to be inspected in detail.
Radiographic tests (RT) and ultrasonic tests (UT) of the four suspect nozzles indicated that no abnormal conditions were present in any of the nozzles.
These findings suggested that the problem may be site-specific to Crystal River-3.
Oconee-3 was also shutdown at that time for a Once-Through Steam Generator (OTSG)tubeleak.
Radiography of one of the makeup nozzles (A2) showed that the thermal sleeve was displaced about 5/8 inch upstream from its normal location.
The radiographic test also revealed that a gap was present between the outside diameter (00) of the thermal sleeve and the inside diameter (ID) of the safe-end in the contact expanded region.
The weld buttons in the safe-end, which 5
prevent upstream motion of the thermal sleeve, had been worn away (see Figure 3).
Weld buttons in the nozzle throat, which prevent f
downstream motion of the thermal sleeve, were still present, but were worn.
A UT of the nozzle also revealed that cracking was present.
Given these indications, the HPI/MU piping and warming line were cut from the safe-end and a dye penetrant test'(PT) of the safe-end and associated hardware was conducted (see Figure 4). The safe-end, thermal sk we, spool piece and warming line were cracked.
Subsequent RT's of the remaining nozzles revealed that the other makeup nozzle (A1) and one of the HPI nozzles (B2) were not damaged and the thermal sleeves were in position.
However, the other HPI nozzle (B1) had a
.030 inch gap between the thermal sleeve OD and the safe-end ID as indicated by the RT.
With the cracking problem substantiated at Oconee-3, Duke quickly inspected their Oconee-2 unit. 'Three of the Oconee-2 nozzles contained anomalies:
(1) the makeup nozzle (A2) had a cracked safe-end and a loose thermal sleeve, (2) the HPI nozzle (B1) had a 1/32 inch gap between the thermal sleeve and safe-end as indicated by the RT, and (3) the HPI nozzle (B2) had a tight thermal sleeve which contained a circumferential crack in the roll expanded region.
a Inspections at four other operating plants were also conducted.
Tiee thermal sleeves at Davis Besse-1 and Three Mile Island-1 (TMI-1) were in position and tight.
No cracking was observed and the weld buttons were not worn.
However, inspections at Arkansas Nuclear One-1 (ANO-1) and Rancho Seco indicated that abnormal conditions were present at these sites.
At AN0-1, three problems were discovered:
(1) one HPI nozzle ( A1) had a loose sleeve, (2) one HPI nozzle ( A2) had a tight I
sleeve with a partial gap indicated by radiography between the sleeve j
and safe-end, and (3) the HPI/MU nozzle (B2) had a tight sleeve which contained a circumferential crack in the roll expanded region (similar to the Oconee-2(B2) failure).
At Rancho Seco, two problems were discovered:
(1) the HPI nozzle (A1) had a loose sleeve, and (2) the HPI/MU nozzle (A2) had a cracked safe-end and a missing thermal sleeve.
9 Inspections at two plants under construction, Midland and North Anna, l
were also conducted to determine the conditions present prior to l
initial plant startup.
Radiographs of the two Midland units indicated that a number of the nozzles may have gaps between the thermal sleeve and safe-end.
Supplemental visual inspections revealed that all 8 sleeves were tight and in place.
However, one of the HPI thermal sleeves on Unit 2 was conspicuously skewed relative to the safe-end center line.
Visual inspections at North Anna revealed that one sleeve had a partial gap in the rolled region, but the sleeve was tight and in place.
The length of the rolled region was also observed to vary between 1 1/2 and 2 inches at North Anna.
In addition, the TMI-2 and Davis Besse-1 nozzles were inspected in 1971 while the plants were under construction.
At TMI-2, all 4 HPI/MU nozzles were inspected and no defects were observed.
However, at Davis Besse-1, one of the sleeves was found to be loose and all 4 sleeves were subsequently re-rolled (hard rolled, instead of contact expanded).
These findings indicate that loose sleeves, or sleeves with gaps between the thermal sleeve and safe-end, may have becn present in other plants prior to initial plant startup.
- 2. 2 Scope Given this background information, the Task Force chose to approach the problem from a generic standpoint (see Figure 5 for the Task Force Action Plan).
To do this, a root cause(s) must first be identified, and then a generic solution could be recommended.
To determine the robt cause(s), the following tasks were performed:
1.
reviewed manufacturing data 2.
compiled and compared site specific facts and inspection results 3.
evaluated metallurgical examinations Y
4.
reviewed industry experience 5.
evaluated data from the instrumented Crystal River-3 HPI/MU l
nozzle 6.
evaluated the existing design analytically 7.
postulated possible failure scenarios 8.
determined a most probable root cause(s)._
m Having determined a most probable root cause(s), a solution was developed which addressed:
1 1.
modified thermal sleeve design for the damaged nozzles 2.
makeup system operating conditions 3.
augmented inservice inspection plan Finally, B&W also proposed studies to demonstrate the adequacy of the recommended fix on a long term basis.
- 2. 3 Resul ts Results of the investigation indicate the following facts:
1.
The thermal sleeve manufacturing installation procedure called for a contact roll of the thermal sleeve, not a hard roll.
2.
Varying degrees of contact expansion rolls could be performed even for the same plant.
3.
Gaps between the thermal sleeve and safe-end have been found in plants under construction.
4.
All cracked safe-ends were associated with loose thermal sleeves.
However, not all loose thermal sleeves had safe-ends that were cracked.
5.
All cracked safe-ends were associated with the makeup nozzle.
6.
A makeup nozzle may be subject to random and continuous makeup flow oscillations.
7 7.
The cracks found were ID initiated (Crystal River-3 OD crack initiation appeared to be unrelated).
[
8.
The cracks were propagated by thermal fatigue.
I 9.
Where controlled hard rolling of the thermal sleeve was accomplished, no failures have occurred.
10.
Oconee-1, which has the most operating experience, contained no abnormal conditions when recently inspected.
Oconee-1 is the only plant which uses a double thermal sleeve design.
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- 2. 4 Organization This report has been organized to address three primary questions:
1 l
1.
How did the Task Force determine the root cause of the problem?
2.
What modifications (design, operation, inspection) were made to correct the problem?
3.
What was done to justify these modifications?
Specifically, sections 3 through 8 describe what was done to determine a most probable root cause, sections 9 through 11 describe what modifications were suggested, and section 12 supplies the justification for these modifications.
In addition, sections 13 and 14 summarize the conclusions and recommendations of this investigation.
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- 3. 0 COMPILATION OF FACTS Following the incidents at Crystal River-3 'and Oconee, the Safe-End Task-Force requested that B&W compile a list of facts concerning the HPI/MU nozzle cracking problem, such that possible correlations between plants could be identified.
To accomplish this task, B&W reviewed the manufacturing records and the site specific failure analysis reports, and then developed a matrix of facts.
3.1 Failure Analyses Failure analyses were performed on four of the units (Crystal River-3, Oconee-3, Oconee-2, and Arkansas Nuclear One-1).
These studies were conducted to determine the most probable method of crack initiation and propagation.
The results are as follows:
Crystal River-3/ Florida Power Corporation While the repair efforts were being completed on the Crystal River-3 unit, the cracked safe-end and thernal sleeve of the HPI/MU nozzle (A1) were sent to B&W's Lynchburg Research Center (LRC), and the cracked valve and section of pipe near MUV-43 were sent to Battelle Columbus Laboratories for failure analysis.
The results of the LRC study indicated that both the sleeve and the safe-end most likely failed by thermal fatigue.
Cracking initiated on the ID of both components and was transgranular.
The thermal sleeve cracking was confined to the roll expansion area only.
The safe-end was cracked in the valve end down to the seating area of the thermal sleeve.
Extensive wear was found on the safe-end ID and the thermal sleeve OD in the region of roll expansion of the sleeve into the safe-end.
From this and other surface damage, it was concluded that the sleeve had become unseated and was probably rotating due to flow forces.
Evidence to confirm or refute whether the sleeve had been roll expanded on installation was not conclusive.[1] l
Battelle's inspection of the pipe section revealed that separate circumferential cracks from the inside diameter (ID) and the outside diameter (00) on half of the pipe section were present, as well as multiple longitudinal cracks.
The circumferential crack on the ID was associated with a machine tool mark, while the crack on the OD was 6ssociated with the valve to weld bead discontinuity.
Fractographic evidence suggested that fatigue was responsible for both the ID and OD circumferential cracks.
Metallography showed that the cracks were transgranular.
The ID cracks were believed to have initiated by thermal fatigue caused by (1) turbulent mixing of hot and cold water during makeup system additions, and/or by (2) periodic chilling of hot i
metal during makeup system additions.
Crack propagation probably occurred by combined thermal and mechanical loading of the system.
The OD crack is believed to have initiated and propagated by mechanical loading of the system.[2]
Oconee-3/ Duke Power The LRC examined the safe-end, thermal sleeve, spool piece, and warming line of the damaged Oconee-3 makeup nozzle ( A2) (See Figures -3 and 4).
Component failures were due to thermal fatigue as with Crystal River; however, the cracking was not as deep or as widespread.
The cracking was transgranular and confined to three regions:
1 1.
the roll expanded end of the thermal sleeve 2.
the safe-end ID from the upstream edge of the thermal sleeve seat to the spool piece weld 3.
the spool piece from the safe-end to about 2 inches upstream of the warming line tee I
In addition, evidence of wear was found on the thermal sleeve OD and the safe-end ID in the area of the contact expansion seat.
As with the Crystal River components, this suggests that the thermal sleeve had become unseated and was rotating / vibrating due to flow forces.[3]
Oconee-2/ Duke Power B&W's LRC also peformed the metallurgical examination of the Oconee-2 HPI nozzle (B2) thermal sleeve.
This sleeve contained a visually observable crack extending approximately 270' around the circumference located about 1 1/2 inches from the roll expanded end of the sleeve.
This large crack was transgranular and at one location was shown to be propagating from ID to 0D.
A small axial branch of this crack contained some fatigue striations, but the bulk of the fracture surface could not be f aterpreted due to heavy oxidation and damage incurred during removal.
Metallographic examination also revealed shallow (<3 mils) transgranular cracking on the OD near the large crack.
This sleeve did not contain a large amount of wear compared to the Oconee-3 and Crystal River sleeves; however, the downstream collar contained a peened surface along a 180* arc (See Figure 3).
In general, the basic failure mode appeared to be transgranular fatigue as occurred in the Crystal River and Oconee-3 thermal sleeves, but the arrangement of the cracking pattern and differences in surface damage suggested that the stress state required to create this failure was either different, or more dominant than in the previous failures.[4]
t l
l Arkansas Nuclear One-1/ Arkansas Power & Light The Lynchburg Research Center also performed a metallurgical examination of the ANO-1 HPl/MU nozzle (B2) thermal sleeve.
The sleeve contained a visible crack extending approximately 270' around the circumference located about 1 1/2 inches from the roll expanded end of the sleeve.
The crack was transgranular, had propagated by fatigue, and followed a machining mark.
No axial cracking was present.
The collar end of the sleeve showed damage to the collar itself approximately 180' around the circumference.
Below this damaged area, approximately 90* apart, two gouged out areas were also present.
The failure mode of this sleeve appeared to be similar in nature to that suggested for the Oconee-2 thermal sleeve.[5]
- 3. 2 Matrix of Facts While the failure analysis studies were being conducted, a site specific matrix of facts was compiled.
Five major areas were addressed:
(1) system characterization, (2) component characterization, (3) operating conditions, (4) unit operation, and (5) inspection results.
Within each specific area, the following items were included:
1.
System Characterization e loop designation e nozzle type (HPI/MU) e pipe layout e pump characterization
- rotation (CW/CCW)
- distance from pump discharge l
- number of impeller vanes l
- number of diffuser vanes e makeup recirculation control
2.
Component Characterization e thermal sleeve geometry e safe-end geometry e thermal sleeve / safe-end interface e material e sleeve expansion procedure 3.
Operating Conditions e minimum bypass flow e total makeup flow e total HPI flow e minimum RC pressure to provide net positive suction head (NPSH) e borated water storage tank (BWST) temperature 4.
Unit Operation e full power years e reactor trips e estimated HPI actuations 5.
Inspection Results e gaps between thermal sleeve 00 and safe-end ID e thermal sleeve axial location a weld button integrity / geometry e thermal sleeve cracking e safe-end cracking Table 1 contains the matrix of facts compiled by B&W.
Examination of this table suggests that two possible correlations may exist between HPI/MU nozzle failures and sites.
1 _ _ _.
First, neither of the units operating with RC pumps which contain 7 impeller vanes (0conee-1 and TMI-1) have ever shown any indications of loosening or cracking of the thermal sleeves.
On the other hand, 5 out of 6 units operating with RC pumps which contain 5 impeller vanes have shown indications of loosening or cracking of the thermal sleeves.
This implies thit the dynamics of the pressure field generated by the RC pumps may lead to flow induced vibration damage.
However, these observations may simply reflect design differences among the plants (0conee-1 uses a double thermal sleeve and TMI-1 uses an Inconel safe-end).
Second, either operating unit which has undergone post-installation inspection or modification (0conee-1 and Davis Besse-1) has not shown any indications of loosening or cracking when recently inspected.
At Oconee-1, a single thermal sleeve was originally installed which extended into the cold leg flowstream approximately 21/8 inches less than the sleeves used at the other plants.
A number of boiling water reactors (BWR) employing a similar design experienced cracking problems.
Consequently, a second longer sleeve was re-rolled inside of the original sleeve.
Aside from increasing the overall length of the sleeve assembly, the rolling of the second sleeve may have also resulted in the re-rolling of the original sleeve.
The second sleeve also had an interlocking flange which contained 4 axial notches in the flanged region.
Weld buttons were placed within these notches to provide additional anti-rotation protection.
At Davis Besse-1, an inspection of the HPI/MU nozzles was performed in 1977 prior to operation.
One sleeve was found to be loose and all four sleeves were subsequently re-rolled.
Consequently, the post-installation modifications and inspections have at least mitigated the problem, and may have completely eliminated the problem.
L -
- 4. 0 REVIEW 0F INDUSTRY EXPERIENCE A literature review of recent nuclear industry experience in cracking problems was performed by B&W.
Five events of interest were identified:
Babcock & Wilcox PWR, Indian Point Thermal Sleeve Failure,1970 [6]
While plugging tubes at the Indian Point-1 facility, fragments of the makeup line thermal sleeve were discovered in the primary side of the steam generator water box.
Apparently, the sleeve had failed as a result of thermal fatigue in the sleeve to makeup line welded area.
The thermal stresses resulted from the flow and temperature gradients associated with normal plant makeup system operations.
The problem was eliminated by (1) using a thermal sleeve assembly made from a solid forging, (2) projecting the thermal sleeve into the RC cold leg an additional 1/R inch to induce better mixing, and (3) increasing the minimum makeup flow to 5000 lb/hr.
GE - BWR, Feedwater Nozzle /Sparger Cracking, 1974-1980 [7]
From 1974 through 1980, 22 of 23 BWR's inspected had experienced some degree of cracking in their primary system feedwater nozzles.
The failures occurred due to thermal fatigue with crack initiation caused by turbulent mixing (high-cycle) and crack propagation caused by intermittent feedwater flow (low-cycle) during startup, shutdown, and hot standby.
The " loose sleeve design" was identified as the root cause which allowed bypass flow within the annulus between the sleeve and the nozzle.
A tight fitting thermal sleeve to restrict bypass flow was used as an interim fix and a triple thermal sleeve design was recommended as a permanent fix.
l )
y GE - BWR, Control Rod Drive Return Line Nozzle Cracking,1975 [8]
In 1975, 12 BWR's were inspected and found to have cracking in the control rod drive return lines (CRDRL) and the reactor vessel beneath the nozzles.
As with the BWR feedwater problem, the failures were attributed to thermal fatigue cracking due to turbulent mixing and intermittent cold water flow.
The problem was eliminated by plugging the nozzlea and rerouting the CRDRL.
Westinghouse - PWR, Steam Generator Feedwater Line Cracking, 1979 [8-10]
In 1979 cracking was discovered in the steam generator feedwater lines of 5 operating PWR systems.
The cracking was attributed to thermal fatigue due
.to flow stratification in the feedwater lines.
Corrosion fatigue was subsequently declared to be the root cause.
Westinghouse - PWR, Loss of Thermal Sleeves in Reactor Coolant System Piping at Certain Westinghouse PWR Power Plants, 1982 [14]
In 1982, 2 Westinghouse PWR's were inspected and found to have missing thermal sleeves in their safety injection (SI) nozzles.
Radiography and ultrasonic examinations confirmed that the 10-inch thermal sleeves were missing from all four SI nozzles at the Trojan nuclear plant.
Supplemental inspections of the sleeves in the pressurizer surge line, and normal and alternate charging lines revealed that cracking was present in some of the retaining welds.
l At Duke Power's McGuire-1 reactor, radiography and underwater camera inspection revealed that the thermal sleeve in one of the four SI accumulator piping nozzles to RCS cold leg piping was missing.
Radiography l
confirmed that the other three SI sleeves and the pressurizer surge line sleeve were in place.
Westinghouse recommended that (1) the loose parts I
monitoring system be fully operational, and (2) a non-destructive i
examination be performed to assess the thermal sleeve conditions of the affected systems at the next extended plant outage.
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in summary, the following observations can be made:
1.
Crack initiation was due to high-cycle thermal fatigue caused by turbulent mixing.
2.
Crack propagation was due to low-cycle tharmal fatigue caused by intermittent flow of cold water.
3.
Tests conducted by Hu et.al. [9] have shown that for loose fitting thermal sleeves, leakage flow (up or down stream) may occur within the annulus between the sleeve and. nozzle.
4.
Cracking occurs in high stress areas, i.e., counter bore transition, weld discontinuities, nozzles blend radius, etc.
5.
All failed components were subjected to a stratified flow caused by low flow rates.
I
- 5. 0 CRYSTAL RIVER-3 INSTRUMENTED HPI/MU N0ZZLE DATA EVALUATION Following the cracking incident at Crystal River-3, metallurgical examinations of the thermal sleeve, safe-end and spool piece were conducted by the LRC and Battelle as previously discussed.
Results of these studies indicated that the cracking was attributable to thermal fatigue.
Given this information, qualitative modifications were made to minimize the thermal st asses within the nozzle assembly.
Subsequent to this effort, the thermal sleeve was replaced with a modified design, the safe-end was replaced, and the HPI/MU check valve was replaced and relocated approximately 5 inches upstream from its original location.
To verify the structural integrity of the modified HPI/MU nozzle design (see Figure 8) and gain insight into the failures, B&W recommended that the makeup nozzle assembly ( A1) be instrumented.
Information was required regarding the thermal stresses and vibrational environment associated with normal plant heatup, hot standby, and power operation.
To provide this information,12 thermocouples, 4 welded strain gauges, 4 bonded strain gauges and 2 accelerometers were installed at three axial planes (A, B, and C) as shown in Figure 6.
Evaluation of the data obtained from the instrumented nozzle indicated that:
1.
The external temperature of the safe-end (plane B) remains at or near the makeup water temperature, while the thick portion of the nozzle (plane A) tends to follow the RC cold leg temperature.
2.
Circumferential temperature gradients were small indicating that no i
significant " hot spots" or flow stratification was occurring.
3.
Several continuous makeup flow rates were tested (1.6, 5.0,15.0, and 130.0 gallons per minute).
In all cases, the safe-end metal temperature did not change, while the nozzle metal temperature changed by a maximum of 20 F. -_____-
4.
During heatup, the makeup flow cycled approximately every three minutes.
The resultant stresses were small.
5.
Makeup flow induced vibrations could be detected with the supplemental instrumentation and tended to increase as makeup flow increased.
The resultant stresses were small.
6.
Noz71e/ safe-end stresses due to thermal expansion are smaller than design values.
7.
High stresses were recorded while a pipe hanger was being set.
This was an isolated occurrence and had no significant influence on the other test results.
For further details, the reader is referred to B&W document 77-1134571-00, " Evaluation of Crystal River-3 HPI/MU Nozzle Testing".
[11]
{
- 6. 0 _ ANALYTICAL INVESTIGATION OF EXISTING DESIGN The previous discussion revealed that the thermal stresses in the modified HPI/MU nozzle at Crystal River-3 were within design values.
However, no data was obtained for the old nozzle design.
Consequently, B&W developed a program to analyze the original (existing) design.
The purpose of the analytical study was two-fold: (1) to determine the relationship between wall thinning of the HPI thermal sleeve during roll expansion and residual stresses at the thermal sleeve to safe-end interface, and (2) to determine if the rolled joint becomes loose during steady-state plant operation, or 'during the most severe transient (HPI event).
1 To determine the thermal sleeve thinning to thermal sleeve / safe-end interfacial residual stress relationship, a finite element model was constructed for a radial sector of the assembly in the contact expanded region (See Figure 3).
Assuming that a generalized plane strain condition exists within this region and that end effects are negligible, a simple axisymmetric, non-linear, inelastic analysis was performed using the ANSYS Code. [12] Results of this finite element analysis follow; however, these results have not been verified and should be used for information only.
The relationship between thermal sleeve wall thinning and sleeve / safe-end interfacial stress is shown in Figure 7.
For wall reductions in the 2-10%
range, the resulting interfacial residual stress lies in the 4000-4200 psi range.
The residual stress varies in a non-linear fashion which suggests that above a certain degree of wall thinning, probably greater than 5%, the beneficial effects of increased wall thinning are negligible.
The loosening of the rolled joint during steady-state and most severe transient operation was investigated analytically by imposing appropriate thru-wall temperature variations on the model used to determine interfacial
(
residual stress.
The temperature distributions were determined assuming one-dimensional heat transfer.
The results show that no gap forms between the sleeve and safe-end during steady-state operation.
However, the -
results indicate that during an HPI event (most severe transient), the thermal sleeve contraction relative to the safe-end causes a small gap to form between the sleeve and safe-end for a short period of time.
l f._ ____.--._.- _ ____ ____ -- _.. _ _______.__ _ __ _ __ _.._ _ __ _ _ _. _.-._ ___... _-,_,.-.. _. __-_.__-._.. _ _
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- 7. 0 POSSIBLE ROOT AND CONTRIBUTORY CAUSES Following the discovery of cracking at Crystal River-3, an effort was made to identify possible root and contributory causes.
The following causes were hypothesized:
1.
Makeup flow conditions maintained outside of design limits - this includes either a low MU temperature, or an incorrect bypass flow rate.
In particular, the bypass flow rate may have been set at ambient conditions instead of at operating conditions, or may not have been properly maintained.
2.
Excessive cycling of the check valve due to improper valve performance 3.
Flow stratification in the MU line due to minimal MU flow 4.
Thermal stratification and recirculation in the MU line due to minimal flow 5.
Cold working of the thermal sleeve due to roll expansion 6.
Stress corrosion cracking of the thermal sleeve due to excessive roll expansion 7.
Convective heating of the safe-end due to an air gap in the insulation 8.
External loading of the attached piping due to thermal transients 9.
Sympathetic vibration of the thermal sleeve due to dynamic pressure field generated by the RC pumps. _. -
u 10.
Flow induced vibrations due to cross-flow in the RC cold leg pipe 11.
Annular flow between the thermal sleeve 00 and the safe-end ID due to insufficient rolling of the thermal sleeves As additional information was obtained from the failure analysis studies and the site inspections, the validity of these causes could be suitably evaluated.
It must also be pointed out that this list was compiled after Crystal River-3; therefore, some of the causes identified are site specific to Crystal River-3 and, thus, do not-apply to all of the sites.
Of the 11 postulated causes, the first 4 pertain to the makeup system exclusively. A quick inspection of the matrix of facts, Table 1, reveals that both HPI and MU nozzles were affected.
Consequently, any cause(s) which pertain to the MV nozzles alone can only be contributory at best.
With this in mind, the validity of each cause was evaluated as follows:
1.
Makeup flow control problems -due to improper maintenance of minimum bypass flow may have occurred at all of the sites.
Plant data i
obtained during heatup and cooldown revealed that makeup flow rates were often unknown to the operators.
As such, minimum continuous flow rates may not have been properly maintained which could lead to thermal fatigue of the nozzle components.
However, since all of the plants experienced similar flow control problems and only 5 of the operating plants contained anomalies, makeup flow control was probably not the root cause.
2.
Excessive cycling of the MU check valve may have contributed to the
]
failure at Crystal River-3, but this was probably an isolated occurrence.
4.
3.
Flow stratification in the MU line due to minimal MU flow may have occurred at all of the plants since the same design value (1-3 gpm) was used inclusively.
However, the results from the instrumented Crystal River-3 nozzle indicated that no significant circumferential temperature gradients were present, even at the lowest flow rate tested (1.6 gpm).
From these findings, it can be inferred that the makeup flow was probably not stratified.
4.
Low flow velocities in the MU line could also lead to thermal stratification and recirculation zones in the thermal sleeve.
However, since the MU line is predominantly filled with MU flow, the thermal shock to the sleeve should not be too extensive (compared to the flow stratification described in 3).
As a result, this can be disregarded as a probable cause.
5.
Cold working of the thermal sleeve was not responsible for crack initiation or growth according to the failure analysis reports discussed in section 3.2.
Consequently, this cannot be considered a probable cause.
6.
Also, stress corrosion cracking due to roll expansion was not observed in the failure analysis studies.
Consequently, this too cannot be considered a probable cause.
7.
Convective heating of the safe-end via an air gap in the insulation may have contributed to the failure at Crystal River-3; however, since some of the plants are uninsulated, this can be disregarded as a probable cause.
l 8.
Excessive loading of the attached piping due to thermal transients may occur at all of the plants.
To ascertain the ex;ent of the thermal transient loading, a structural analysis was performed for the Crystal River-3 piping arrangement.
The results indicated that all stresses were well within the allowable design constraints.
l Therefore, this cause can be disregarded.
j i
l l
l _ _ - _ _ _ _
9.
Sympathetic vibration of the thermal sleeve induced by the motion of the impeller vanes past the discharge port of the RC pumps may have occurred at all of the plants.
The matrix of facts, Table 1, indicates that 5 of 6 plants using RC pumps with 5 impeller vanes have shown loosening or damage of the thermal sleeves.
In contrast, both plants which use RC pumps with 7 impeller vanes have not shown any signs of failure.
The results from the instrumented nozzle at Crystal River-3 indicated that the flow induced vibrations (FIV), as measured by strain gauges and accelerometers, were minimal.
From these findings, it can be inferred that (1) the modifications made at Crystal River-3 have either substantially reduced or el'iminated the FIV problem, and/or (2) the FIV problem is a typical high-cycle fatigue problem which takes a finite amount of time to loosen the rolled joint.
Loosening of the joint would allow mixing of hot RC cold leg water and cold MU water in the annular region between the thermal sleeve and safe-endr This, in turn, would lead to thermal fatigue of the thermal sleeve and safe-end as described in the failure analysis reports.
Consequently, FIV due to the RC pumps may have contributed to the failures.
10.
Similarily, FIV due to cross-flow in the RC cold leg.nay have loosened the rolled joints.
However, all of the plants experienced this form of FIV and were not affected.
Therefore, this is probably not a root cause.
11.
The thermal sleeves could have been rolled to varying degrees (loose and/or with gaps between the thermal sleeve and safe-end) when originally installed.
This would allow mixing of the hot RC cold leg flow and the cold HPI/MU flow in the annular region between the thermal sleeve 0D and the safe-end ID.
This phenomenon, in turn, would thermally shock the nozzle components and eventually lead to crack initiation and propagation.
l -
- 8. 0 PROBABLE FAILURE SCENARIO With the foregoing discussion in mind, the Safe End Task Force developed a probable failure scenario based on hypothesis 11 of Section 7.0.
"The most likely scenario for failure is that the thermal sleeve is loose after construction or a minimum contact expansion roll becomes loose during operation due to mechanical vibration and/or thermal cycling of the contact expansion joint.
This looseness causes wear of the OD of the thermal sleeve and the ID of the safe-end.
This wear in the rolled area allows a larger gap to form between the thermal sleeve and safe-end.
Hot reactor coolant flows around the sleeve through this gap.
The hot coolant randomly impacts the safe-end and thermal sleeve area because of random motions of the sleeve.
The cooler makeup flow cools these heated areas when random motion shuts off the annular flow or makeup flow is increased.
This random alternating heating and cooling eventually causes thermal fatigue cracking of the safe-end.
This cracking may be aggravated by heating and cooling caused by significant cycling of makeup flow."[13]
Facts to support this hypothesis are as follows:
e Inspections conducted at Davis Besse, Midland and North Anna have shown that loose sleeves, or sleeves with gaps between the thermal sleeve and safe-end were present in plants under construction.
In addition, the North Anna inspection indicated that the length of the rolled area varied from nozzle to nozzle between 1-1/2 and 2 inches, e The thermal sleeve contact expansion process, as defined in the original installation procedure, is ambiguous.
e Since the sleeves were rerolled (hard rolled to 3% wall thinning) at Davis Besse-1 in 1977, no additional problems have been observed.
wm--
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e When the modified thermal sleeve was meticulously rolled into the HPI/MU nozzle at Crystal River-3, no abnormal conditions were observed.
e When the failure analyses were performed (see section 3.2), thermal fatigue was identified as the mechanism of crack propagation.
1 t
f t
r
9,0 MODIFIED THERMAL SLEEVE DESIGN The previous sections of this report have been dedicated to determining the root cause of the HPI/MU nozzle cracking problem.
The next three sections address the modifications made to alleviate the problem.
Specifically, these modifications affect the design, operation, and inspection of the HPI/MU nozzles.
9.1 Conceptual Designs In the aftermath of the Crystal River incident, the effectiveness of the contact rolled thermal sleeve design was re-evaluated.
Three alternative concepts for shielding the HPI nozzle from cold injection water were developed.
Each concept uses a stainless steel thermal sleeve which is secured into the nozzle and projects into the RC cold leg piping.
The approaches are as follows:
Hard Rolled Thermal Sleeve Concept A hard rolled thermal sleeve design was developed (see Figure 8),
which requires a hard roll of the upstream end of the thermal sleeve, instead of a contact roll.
Since the same concept was used in the original design, the hard rolled concept should be easy to implement.
However, the problem of loosening of the rolled joints may still exist.
Integral Thermal Sleeve Concept An integral thermal sleeve concept was developed which incorporates the thermal sleeve and the safe-end into a single component (see Figure 9).
This design eliminates the possibility of the sleeve loosening and also eliminates the concern about annular flow between._.
the thermal sleeve and the safe-end.
However, disadvantages of this concept include:
(1) increased pressure drop due to reduced thermal sleeve ID, (2) fabrication problems, (3) welding problems, (4) excessive cost, and (5) an inability to meet fatigue design requirements as specified in code B31.7,1968 draft.
Flanged Thermal Sleeve Concept B8W's flanged thermal sleeve concept is shown in Figure 10.
The flanged connections allow easy access to the thermal sleeves for inspection and replacement.
The concept also provides a positive seal against water flow in the annular region.
The disadvantages of this concept, on the other hand, include:
(1) re-routing of piping, (2) thermal shock to the gasket, and (3) reliability of the gasket.
B&W angineers concluded that the hard rolled thermal sleeve concept represented the optimum choice from a cost, licensing, and leakage standpoint.
- 9. 2 Design Improvements The redesigned hard rolled thermal sleeve (See Figure 8) was developed with some notable improvements:
1.
Bell shaped upstream end on the thermal sleeve - This should prevent movement of the sleeve towards the RC cold leg piping.
2.
Increased length and width of the upstream end of the thermal sleeve - This feature provides more roll surface contact area l
and more metal to be cold worked during the rolling process.
l l
l I,
I t
3.
Hard roll of the thermal sleeve shoulder - The original thermal sleeve was only contact rolled.
The increased compression and subsequent deformation of the thermal sleeve material should provide a more secure bond with the safe-end.
Also, the additional wall thinning should mitigate sleeve to safe-end separation during HPI events.
4.
Contact roll at the thermal sleeve collar - The effects of possible flow induced vibration will be reduced with the sleeve surface in contact with the nozzle ID.
5.
Axially notched upstream end of the thermal sleeve - The 4 notches allow the placement of weld beads to provide additional anti-rotation protection.
In summary, the thermal sleeve has been redesigned to eliminate the causes which contributed to the failures at Crystal River, Oconee, ANO, and Rancho Seco.
10.0 MAKEUP SYSTEM OPERATINr CONDITIONS l
Aside from the redesign of the thermal sleeve, modifications to the makeup l
system operating conditions were also suggested following the Crystal River incident.
The original design specification called for a minimum continuous makeup flow of 1-3 gpm.
It was believed that at this limited flow rate, flow and thermal stratification could occur in the makeup line which may lead to thermal fatigue of the nozzle assembly.
Similar flow conditions at 5 Westinghouse PWR's [8-10] in 1979 lead to cracking of the steam generator feedwater lines.
Consequently, a minimum bypass flow of 15 gpm was suggested to eliminate' or at least mitigate this potential problem.
As additional information was obtained, the recommended 15 gpm minimum makeup flow rate was re-evaluated.
The results from the instrumented Crystal River-3 nozzle indicated that the new design achieved all design requirements even at the lowest flow rate tested (1.6 gpm).
The safe-end remained cool, while the outer surface of the nozzle varied by at most 20 F.
The circumferential temperature gradients were small indicating that no significant " hot spots" or flow stratification was occurring.
Also, as the makeup flow rate was increased to a maximum of 130 gpm, the nozzle thermal stresses tended to decrease.
In light of these findings, a minimum continuous makeup flow of 1-3 gpm (as originally specified) should adequately maintain all design parameters within analyzed limits and prevent thermal stratification.
However, it must also be pointed out that increasing continuous makeup flow may decrease the nozzls thermal stresses.
1
11.0 AUGMENTED INSERVICE INSPECTION PLAN Along with the thermal sleeve redesign and the MU system operating changes, an augmented inservice inspection (ISI) plan was also developed.
An ISI provides a means of early problem detection, such that repairs can be effected before extensive damage occurs.
Prior to Crystal River, no HPI/MU nozzle assembly inspection was required.
B&W and the Safe-End Task Force developed an augmented ISI for the 177 FA Owner's Group.
Specifically, the plan calls for:
l Makeup Nozzles 1.
Unrepaired Nozzles
- RT during the next five refueling outages to ensure that the thermal sleeve is in the proper location and no gap exists between the thermal sleeve and safe end.
Ensure RT is comparable with
" baseline" first RT taken.
Perform RT every fifth refueling outage thereafter.
- UT the safe end and some length of adjacent pipe / valve during the next five refueling outages to ensure no cracking.
Perform UT every fifth refueling outage thereafter.
2.
Repaired Nozzles (New Sleeve Design)
- RT during the first refueling outage to ensure that the thermal sleeve is in the proper location and no gap has formed.
- UT safe end, cold leg ID nozzle knuckle transitian, and adjacent piping / valve during the first refueling nutage to ensure no cracking exists.
- RT and UT again at third and fifth refueling outages after repair and every fifth refueling outage thereafter.
3.
Repaired Nozzles (with re-rolling)
- RT during the next five refueling outages to ensure that the thermal sleeve is in the proper location and no gap exists between the thermal sleeve and safe end.
Ensure RT is comparable with
" baseline" first RT taken.
Perform RT every fifth refueling outage thereafter.
High Pressure Injection Nozzles 1.
Unrepaired
- RT during the next five refueling outages to ensure that the thermal sleeve is in the proper location and no gap exists.
Ensure RT is comparable with " baseline" first RT taken.
Perform RT every fifth refueling outage thereafter.
2.
Repaired (New Sleeve Design)
- RT during first refueling outage to ensure that the thermal sleeve is in the proper location and no gap has formed.
RT during third and fifth refueling outages and every fifth refueling outage thereafter.
- UT the ID nozzle / cold leg transition knuckle area during the first refueling outage to assure that no cracking is present.
UT during third and fifth refueling outages thereafter.
3.
Repaired (with re-rolling)
- RT during the next five refueling outages and every fifth refueling 7
outage thereafter to ensure a gap does not form. [
12.0 JUSTIFICATION OF LONG TERM OPERATION l
Finally, having described the modifications (design, operation, inspection) made to correct the problem, we must now consider the steps taken to support these changes.
Specifically, continued operation on a long term basis will be justified both analytically and by inspection of nozzles in service (see section 11 for details).
After the repair efforts were completed at the damaged sites, the NRC staff required that the new design be proven safe for operation in the near term.
In response to this request, B&W provided certified field change authorizations (FCA) to the utilities.
These FCA's were predicated on simple, yet conservative stress analysis, worst case operational histories, and the consideration of continued nozzle usage through the next fuel cycle only.
As such, these studies were only valid in the short term.
In order to justify long term use, B&W recommended a more extensive stress analysis.
The stress information required for more detailed evaluation of makeup and HPI nozzle design changes can be obtained most accurately through the use of the finite element method of structural analysis.
This analysis technique will determine, in detail, the stresses in the critical areas and will provide the means to assess the impact of unanticipated operating transients on the makeup and HPI nozzles.
Such an analytical capaoility will be invaluable at some later date if, for example, an HPI nozzle that had a loose thermal sleeve was subjected to more HPI flow cycles than can presently be shown to be acceptable using conservative techniques.
In addition, evaluation of thermal sleeve / safe-end interface stresses may be required, at a later date, for unanticipated makeup nozzle flow transients.
Inservice inspection (ISI) detected flaws could also be less conservatively evaluated if the new detailed stress profiles were available for use in determining the number of cycles for thru-wall crack propagation. -
B&W's modified nozzle design is currently being used for both the double-duty HPI/MU nozzles and the HPI only nozzles.
However, design differences in service conditions between the two nozzle functions lead to radically different stress distributions.
For the HPI/MU nozzle with continuous 95 F makeup flow, injection of HPI water at 40*F (design temperature) is normally not considered to be a severe transient.
The highest stresses for this nozzle are at the point where the HPI/MU pipe penetrates the RC pipe (nozzle " knuckle" region) and are due to the steady axial temperature gradient between the relatively cool safe-end and the hot RC pipe.
On the other hand, the insulated HPI only nozzle is kept hot through heat conduction from the RC pipe under conditions of no HPI flow.
When HPI is actuated, the sudden flow of 40*F water (design conditions) causes severe thermal stresses at the thin walled portion of the upstream end of the safe-end.
Contributing to the stresses in this region are a severe radial temperature gradient and a local axial temperature gradient.
Although the HPI/MU and HPI only nozzles see different service conditions and experience different stress distributions, a single finite element model will suffice for both nozzle functions.
The only exception will be substructured regions where a refined mesh is required to investigate highly stressed locations (e.g., near the wide collar for the makeup nozzle and in the safe-end for the HPI nozzle).
Ultimately, the stress analysis using this model will quantify the usable lifetime of the modified design.
13.0 CONCLUSION
S Based upon the information presented, the following conclusions can be drawn:
1.
Variations in contact expansion of the thermal sleeves is the most probable root cause of the failures.
2.
Continued operation in the short term is acceptable with the modified design.
3.
If continued inspections show that the sleeves are properly in place, it is not expected that the sleeves will loosen during plant operation prior to subsequent inspections, w
l 14.0 RECOMMENDATIONS As a result of the Safe End Task Force's investigation into the HPI/MU nozzle component failures, the fol',owing recommendations are made:
1.
In terms of future repairs, it is recommended that:
Nozzles with Original Design Thermal Sleeves Reroll the upstream end of the thermal sleeve when inspections indicate that a gap exists.
A 5.0% wall reduction is suggested to achieve an adequate interfacial residual stress and avoid stress corrosion cracking of the thermal sleeve.
Nozzles with Modified Design Thermal Sleeve Repair and/or replace the damaged components if inspections reveal that abnormal conditions are present.
In either case, the affected utility should also verify that the components attached to the safe-end meet the design constraints used in the stress analysis.
2.
In order to ensure proper HPI/MU system operation, it is recommended that:
- A continuous makeup flow via bypass of the Pressurizer Level Control Valve should be maintained.
- A known amount of bypass flow which is greater than 1.5 gpm should be maintained and checked frequently (increased flows of up to about 1
10-15 gpm may be preferable depending upon plant configuration and operating practices). -.
- There should be a consistent set of procedures to initiate continuous bypass flow o RCS temperature o RCS pressure o Bypass flow rate o Frequency of adjustment and calibration
- The makeup tank temperature should be maintained within the proper control band as determined by other plant parameters.
- In the event that future anomalies are discovered, proper logging of HPI initiations will be invaluable.
This procedure should include:
I o Nozzles used o Temperature of BWST o Temperature of cold leg before and after HPI initiation o Pressure o Flow rate o Duration of HPI flow 3.
An augmented inservice inspection plan as stated in Section 11.0 should be implemented.
4.
A detailed stress analysis of a nozzle with a modified thermal sleeve design should be performed to justify long term operation.
4 ____
REFERENCES l
1.
Lynchburg Research Center, " Examination of Normal Makeup Nozzle Thermal Sleeve and Safe-End", No. 77-1131543-00.
2.
Sykes, J. F., Hoch, G. H., Beavers, J. A. and Berry, W. E. of Battelle Columbus Laboratories, " Failure Analysis of Safe-End to Check Valve Pipe Weld in the Crystal River-3 Makeup Water System", March 1,1982.
3.
D. L. Baty of Lynchburg Research Center, " Examination of Oconee-3 HPI Thermal Sleeve, Safe-End", LR:82:5463-08:01, March 26,1982.
4.
D. L. Baty of Lynchburg Research Center, " Examination of HPI Thermal Sleeve from Oconee-2", RDD:83:5492-08:01, June 2, 1982.
5.
M. R. Dietrich of Lynchburg Research Center, " Examination of Thermal Sleeve from AN0", RDD:83:5032-08:01, June 28, 1982.
6.
A. Flynn, et al., " Thermal Sleeve Failure and Repairs - Indian Point #1 Nuclear Unit (285 MW)", Nuclear Technology, Vol. 25, January 1975.
7.
R. Snaider, "BWR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking", NUREG-0619, November 1980.
8.
R. W. Klecker, " Thermal Fatigue Cracking in Pressurized Water Reactor Feedwater Lines", Pipe Crack Study Group Supporting Document, June 18, 1980.
9.
M. H. Hu, J. L. Houtman and D. H. White, " Flow Model Test for the Investigation of Feedwater Line Cracking for PWR Steam Generators", ASME Conference, Denver, Colorado, June 21-25, 1981.
- 10. W. H. Bamford, A. Thurman and M. Mahlab, " Fatigue Crack Growth in Pressurized Water Reactor Feedwater Lines", ASME Conference, Denver, Colorado, June 21-25, 1981.
- 11. " Evaluation of Crystal River-3 HPI/MU Nozzle Testing", B&W Document ID 77-1134571-00, May 1982.
- 12. "ANSYS Engineering Analysis System - User's Manual", Swanson Analysis Systems, Inc.,1978 Edition, Revision 3, (Update 67L).
- 13. " Evaluation of Preliminary Cause of HPI/MU Nozzle Thermal Sleeve and Safe-End Failures", Minutes from the Safe End Task Force Meeting, Bethesda, Maryland, May 7, 1982.
8
- 14. Loss of Thermal Sleeves in Reactor Coolant System Piping at Certain Westinghouse PWR Power Plants, Nuclear Regulatory Commission Office of Inspection and Enforcement, IN 82-30, Washington, D.C. (1982), pp 1-2.
9 -
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TABLE 1 MATRIX OF FACTS PAGE 1 DIAMETRICAL PLANT RV PIPE CUST.
N0ZILE INSPECTION TH. SLEEVE N0ZZLE ID GAP BETWEEN THERMAL SAFE END
TY?E RESULTS COLLAR 00 IN COLLAR TH.SL. COL.8 SLFEVE ID (See Note 2)
AREA N0Z. (MIL) 10/0D OCOWEE 1 a.X Al MU/!PI OK XY A2 MU/'IPI YZ B2 iPI ZW B1
- 1 OCONEE 2' WX B44 B1 iPI 6
1.762 XY B41 B2 HPI C'
2.031 1.763 YZ B40 A2 MU/iPI A
1.763 ZW B46 Al MUT.iPI OK 1.763 OC0:4EE 3 WX B44 81
{PI B
2.003 2.015 12 1.500/1.754 1.762 XV (See Note 4)
B2
. lPI OK YZ 840 A2 MU/ilP!
A 1.992 2.003 11 1.500/1.752 1.762 ZW (See Note 4)
Al MJ/4PI OK TM! I WX 844
. tPI OK Af
$41 t/@I YI 340
- !!'1 24 LA6
(* 1 1M1 2 WX XY YZ 7W CR 3 UA B44 A2 lIPI OK 1.993 2.004 11 1.498/1.754 1.763 XY B41 Al MU/itPI A
1.994 2.004 10 1.498/1.752 1.764 YZ 840 S1
- 1PI OK 1.992 S.003 11 1.497/1.753 1.762 IW 846 B2 IIPI OK 2.003 2.013 10 1.502/1.754 1.763 ANO 1 WX B44 C
- iPI OK 1.991 2.002 11 1.500/t.754 1.762 XY B41 0
MU/11PI C
1.989 2.002 13 1.499/1.754 1.762 YZ B4U 3
!!PI B
1.982 1.994 12 1.500/1.754 1.762 ZW C41 A
!!PI B
1.993 2.003 10 1.499/1/754 1.764 RANCH 3 WX E44 0
llPI OK 1.989 2.000 11 7/7 1.762 SECO XY 346 C
flPI OK 1.992 2.003 11 1.500/1.754 1.762 YZ 84J A
MU/hP1 A
1.981 1.992 11 1.500/1.753 1.761 ZW B41 B
HP!
B 1.990 2.003 13 1.498/1.754 1.764 MIDLAND 1 WX 844 A
HPI l.993 2.005 12 1.762 Xf B41 B
MU/HPI 1.993 2.006 13 1.762 YZ B40 C
HPI 1.990 2.002 12 1.762 TW B41 0
HPI 2.008 2.020 12 1.762 MIDLAND 2
'd X 844 C
HP1 1.998 2.010 12 X'
B41 D
HPI 1.993 2.006 13 G
B40 A
MU/HPI 1.997 2.010 13 IW B41 B
HPI 1.994 2.006 12 DAVIS NX S56 A2 iPI OK 2.004 2.016 12 1.500/1.753 1.762 BESSE 1 XY 061 Al IPI 2.003 2.015 12 1.498/1.754 (See
?1 BS7 81 MUn P1 1.985 1.991 12 Note 5)
IW B44 E2 iPI 2.003 2.018 15 1.500/1.750
TABLE 1 MATRIX OF FACTS PAGE 2 EXPANSION IWO.
SOURCE REFERENCE DOC.
PLANT RV PIPE CUST.
>Mor htcbaus SITE COLD LEG ASS'Y IDENT.
TYPE HT. NO. ANO HT. NO. AND LOC-DATE TOOL NO.
REFERENCE IDENTIFIED BY MAT'L. SPEC.
MAT'L. SPEC.
DRAWINGS PIPE SER.NO.
OCONEE 1 WX Al MU/HP!
XY A2 MU/ EPI YZ B2 HPI ZW B1 HPI CCONEE 2 WX B44 B1 HPI
-A336F8M 43116-A336F8M SITE (See (See 146614E-5 B44-204-50-1 XY B41 82 H?!
Mote 3)
Note 3) 146629E-7 B41-204-50-1 YZ B40 A2 MU/HPI B40-204-50-1 ZW B46 Al MU/HPI B46-204-50-1 OCONEE 3 WX B44 B1 FPI 05477-A336F M 65047-A336F M MTV 150141E-7 B44-209-50-1 XY (See Note 4)
B2 FPI 150156E-7 YZ B40 A2 MU/HPI 11-18-71 7573-1 B40-209-50-1 ZW (See Note 4)
Al MU/HP1 TMI 1 WX B44 FPI
-$8 166 SITE (See (See 131956E 7 XY B41 MU/HPI Note 3)
Note 3) 160493E-0 YZ B40 Foi 131960E-9 ZW B46 FPI TMI 2 WX 141578E-9 XY 141576E-13 YZ ZW CR 3 WX 844 A2 FPI 05477-A336F8M 810906-A336FM MTV 9-7-71 7573-1 141599E-5 B44-207-50-1 XY 841 Al MU/FPI 141597E-5 B41-207-50-1 YZ B40 B1 PPI 9-8-71 7573-1 B40-207-50-1 ZW B46 B2 FPI 9-11-71 7573-1 B46-207-50-1 ANO 1 WX B44 C
FPI 05477-A336F8M 811236-A336F8M MTV 3-7-72 7573-1 131998E-4 B44-208-50-1 XY B41 D
MU/HPI 3-15-72 1319%E-6 B41-208-50-2 YZ 840 B
FPI 81564-11-12-71 B40-208-50-1 ZW B41 A
HP!
811236-12-1-71 B41-208-50-1 RANCHO WX 844 D
FPI 05477-A336FM 129186-A336F8M MTV 1-8-72 7573-1 143491E-7 B44-2011-50-1 SECO XY B46 C
FPI 12-30-71 143509E-8 B46-2011-50-5 YZ 840 A
MU/FPI 12-30-71 B40-2011-50-1 ZW B41 B
FPI 1-6-72 B41-2011-50-1 MIDLAND 1 WX B44 A
HPI 818442-A336F8M 43116-A336F8M MTV 9-20-74 7573-1 150176E-6 B44-2012-50-1 XY B41 B
MU/HPI 12-9-74 150191E-1 B41-2012 50-1 YZ 840 C
HPI 10-16-74 B40-2C12 1 ZW B41 D
HPI 9-27-74 B41-2012-50-2 '
MIDLAND 2 WX 844 C
HP!
121294-A336F8M 817962-A336F8F MTV 10-15-7!
7573-1 150206E-4 844-2013-50-1 XY B41 D
HPI 29006 9-28-75 150221E-2 B41-2013-50-1 YZ B40 A
MU/HPI 817962-10-16-7!
B40-2013-50-1 ZW B41 B
HPI 43116-9-23-75 B41-2013-50-2 DAVIS WX B56 A2 PPI 05477-A336F8M S11584-A336F8M MTV & 6-27-72 7673-1 152027E-4 B56-2014-50-1 BESSE 1 XY B61 Al FPI SITE 7-6-72 152042E-4 B61-2014-50-1 (See YZ 859 B1 MU/PPI (See 6-16-72 B59-2014-50-1 Note 5)
ZW B44 B2 PPI 48417 Note 7-3-72 B44-2014-50-1 51
t TAILE 1 MATRIX OF FACTS PAGE 3 (a)
NO. OF RG NO. OF RG PLANT RV PIPE CUST.
N0ZZLE PLMP ROTATION FLOW LENGTH COLD LEG GEOM.
2/2 RC FLOW PtMP PtNP SITE COLD LEG ASS *Y IDENT.
TYPE FROM RC PIMP
& N0ZZLE DATE (5 of 131.3 IMPELLER DIFFUSER ORIENTATION x 10 lbm/hr)
VANES VANES OCONEE 1 WX Al MU/HPI 2 CCW/ LOOP 5.2 f t.
Type A 109%
7 12 XY A2 MU/HP!
YZ B2 HP[
ZW B1 HP!
OCONEE 2 WX B44 B1 HPI 2 CCW/ LOOP 5.2 f t.
Type A 112%
5 4
XY B41 82 HPI YZ 840 A2 MU/HPI ZW B46-Al MU/HP!
OC0 TEE 3 WX B44 B1 HPI 2 CCW/ LOOP 5.2 f t.
Type A 112%
5 4
XV (See Note 4)
B2 HP1 YZ B40 A2 MU/HPI ZW (See Note 4)
Al MU/HPI THI 1 WX B44 HP!
2 CCW/ LOOP 5.2 f t.
Type A 109%
7 12 XY B41 MU/HPI YZ B40 HP' ZW B46 HP:
TMI 2 WX XY YZ ZW CR 3 WX B44 A2 HP.
2 CCW/ LOOP 5.2 ft.
Type B 112%
5 9
XY B41 Al MU/HP:
YZ B40 B1 HP!
Z's B46 B2 HP!
ANO 1 WX B44 C
HP:
2 CCW/ LOOP 5.2 f t.
Type B 112%
5 9
XY B41 D
MU/HP' I_
YZ 240 B
HP.
ZW B41 A
HP ;
RWCHO WX B44 D
HP:
2 CCW/ LOOP 5.2 f t.
Type A 116%
5 4
SECO XY B16 C
=
YZ 310 A
MU/HPI ZW B41 B
HPt 7
MIDLAND 1 WX A
- 2 CCW/ LOOP 5.2 ft.
Type B
- 100%
- 5
+9 XY B
MU/HP!
YZ C
HPI ZW D
HPI MIDLAND 2 WX C
- 2 CCW/ LOOP 5.2 ft.
Type B
- 100%
- 5
- 9 i
XY D
HPI YZ A
MU/HPI ZW B
HP!
1 CW & 1 CGn t
DAVIS WX B56 A2 HP!
per LOOP 9.1 f t Type C 114%
5 9
BESSE 1 XY B61 Al HP:
(Ses YZ B59 B1 MU/HP.
Note 5)
ZW B44 32 HP' r
(a) See Attachments i
l-
TABLE 1 MATRIX OF FACTS PAGE 4 MINIMlM ALLOWABLE RG TOTAL MAKEUP TOTAL MAKEUP IOTAL HPI FLOW PLANT RV PIPE CUST.
NO2ZLE PRESSUPE TO PROVIDE FLOW WITH 1 MU FLOW WITH 2 MU WITH 1 PUMP SITE COLD LEG ASS'Y IDENT.
TIPE NPSH FOR RC PtNPS AT PtNP OPERATION PLNP OPERATION OPERATION AT 160" F (2/2)
AT 2150 PSIG AT 2150 PSIG 1500 PSIG OCONEE 1 WX Al MU/HPI 300 PSIG 157 GPM 186 GPM 360 GPM XY A2 MUiHPI YZ B2 HPI ZW B1 HPI OCONEE 2 WX B44 B1 HPI 170 PSIG 157 GPM 186 GPM 360 GPM XY B41 B2 HPI YZ B40 A2 MU/HPI ZW B46 Al MU/HPI OCONEE 3 WX 844 B1 HPI 215 PSIG 157 GPM 186 GPM 360.GPM XY (See Note 4)
B2 HPI YZ 840 A2 MU/HPI
=
ZW (See Note 4)
Al MU,HP1
=
=
TMI 1 WX 844 HPI 290 PSIG 145 GPM 165 GPM 405 GPM XY B41 MU/HPI YZ 540 HPI ZW B46 HPI TMI 2 WX XY YZ ZW (c)
CR 3 WX B44 A2 HPI 230 PSIG 147 GPM 185 GPM 410 GPM XY B41 Al MU/HPI YZ 840 B1 HPI Td B46 B2 HPI ANO 1 WX B44 C
HPI 142 GPM 180 GPM 405 GPM XY B41 D
MUiHPI fZ 840 B
HPI ZW B41 A
HPI RANCHO WX 844 D
HPI 102 PSIG 192 GPM 288 GPM 405 GPM SECO XY B46 C
HPI YZ 340 A
MU/HPI TJ B41 B
HPI MIDLAND 1 WX A
- 265 PSIG NOT
- 420 GPM TOTAL
)-
XY B
MU/HPI for minimum 140 GPM AVAILABLE YZ C
HPI seal staging Td 0
-HPI QIDLAND 2 WX C
- 265 PSIG NOT
- 420 GPM TOTAL XY D
. HPI for minimum 140 GPM AVAILABLE Y2 A
MU/HPI seal Staging TJ B
HPI (c)
DAVIS WX B56 A2 HPI 190 PSIG 164 GPM 264 GPM 300 GPM BESSE 1 XY B61 Al HPI (See YZ B59 B1 MU/HPI Note 5)
Td B44 B2 HPI 0
(c) at 260 F
m TAE.E 1 MATRIX OF FACTS PAGE 5 TOTAL HPI FLOW TOTAL HPI FLOW BW5T PLANT RV PIPE CUST.
N0ZZLE WITH 2 PtMP WITH 3 PUMP RECIRCULATION TEMPERATURE FULL POWER REACT 05 SITE COLD LEG AS5'Y IDENT.
TYPE OPERATION AT OPERATION AT CONTROL MEANS (NORAL YEARS TRIPS 1500 PSIG 1500 PSIG OPERATION)
(b)
BLOCK ORitIGL OCONEE 1 WX Al MU/HP!
720/540 GPM 900 GPM (N0 ESFAS ISOL.)
80* F -
5.1 87 XY A2 MU/HPI YZ B2 HPI ZW B1 HPI (b)
BLOCK Of.lP ILL
~
OCONEE 2 WX B44 81 HPI 720/540 GPM 900 GPM (NO EVAS ISOL.)
80" F 4.82 53 XY B41 B2 HPI YZ B40 A2 MU/hPI ZW B46 Al MU/HP!
(b)
BLOCK ORitIct OCONEE 3 WX B44 B1 HPI 720/540 GPM 900 GPM (NO ESFAS ISOL.)
80" F 4.99 47 XY (See Note 4)
B2 HPI
=
=
=
YZ B40 A2 MU/HPI
=
=
=
ZW (See Note 4)
Al MU/HPI
=
=
=
TMI 1 WX B44 HPI 810 GPM FLOW ORIFICE 7'*#
3.51 18 XY B41 MU/HP!
E B40 HPI ZW B46 HPI TMI 2 WX XY YZ ZW CR 3 WX B44 42 HPI 790 GPM 1130 GPM FL0h ORIFICE 2.66 56 XY B41 Al MU/HPI YZ 840 B1 HPt ZW B46 B2 HPi ANO 1 h1 944 C
HP' 780 GPM FLOW ORIFICE 4.63 56 Xf B41 D
MU/HP:
YZ 840 B
HP:
ZW B41 A
HP '.
RANCHO WX Bil 0
HPi 585 GPM 650 GPM FLOW ORIFICE 3.87 52 SECO XY 356 C
HPI YZ B40 A
MU/HPI ZW Bd 1 8
HP ~.
0 MIDLAND 1 WX A
HP!
'675 GPM NOT
- FLOW ORIFICE 40 F-110 F 0
0 XY B
MU/HPI TOTAL AVAILABLE (DEPENDING YZ C
HPI ON THE ZW D
HPI WEATHER) 0 0
MIDLAND 2 WX C
- 675 GPM NOT
- FLOW ORIFICE 40 F-110 F 0
0 XY D
HP!
TOTAL AVAILABLE (DEPENDING l
YZ A
MU/HPI ON THE ZW B
HPI WEATHER)
DAVIS WX B56 A2 HP:
600 GPM FLOW ORIFICE 2.01 46 BESSE 1 XY B61 Al HP.
(See YZ B59 B1 MU/HP.
Note 5)
ZW B44 B2 HPI r
l (b) 2 pump operation for CNS-I!! can either be:
l 1 HP1 Train with 2 pumps or 1 HPI Train with 1 pump and 1 HPI Trali with 1 pump
-e--+
--y, 9
.-w-%-
y
--.91---+
p
.y 9
-y--.-w y,.-
9 s-
-v-
TABLE 1 MATRIX OF FACTS
'PAGE 6 PLANT RV P!PE CUST.
N0ZZLE EST. MAX.
EST. HPI MU/HPI SITE COLD LEG ASS'Y IDENT.
TYPE HPI N0ZZLE TO CONNECTION ACT.
N0ZZLE OCONEE 1 WX Al MU/HPI (20) 87 PIPE / PIPE XY A2 MU/HPI 87 YZ B2 HPI ZW B1 HPI OCONEE 2 WX B44 B1 HPI (13)
PIPE / PIPE XY B41 B2 HPI YZ B40 A2 MU/HPI 53 ZW B46 Al MU/HPI S3 OCO. NEE 3 WX 844 B1 HPI (17)
PIPE / PIPE XY (See Note 4)
B2 HPI YZ B40 A2 MU/HPI 47 ZW (See Note 4)
- Al MU/HPI 47 TMI 1 WX 844 HPI CIECK VALVE XY B41 HU/HPI YI B40 HPI ZW B46 HPI TMI 2 WX XY YZ ZW CR 3 WX B44 A2 HPI 39 CHECK VALVE XY B41 Al MU/HPI 49 YZ B40 B1 HPI 36 ZW B46 B2 HPI 37 ANO 1 WX 244 C
HPI (17)
ELBOW XY B41 D
MU/HP!
56 YZ 840 B
HPI ZW B41 A
HPI l
RSNCHO WX B44 0
HPI (31)
ELBOW SECO XY B46 C
(
YZ 840 A
MU/HPI 52 ZW B41 B
HPI MIDLAND 1 WX A
- 0
- 0 SEVERAL FEET XY B
MU/HPI YZ C
HPI ZW D
HPI MIDLAND 2 WX C
- 0
+0 SEVERAL FEET XY D
HPI YZ A
MU/HPI ZW B
ELBOW BESSE 1 XY 861 Al HPI (See YZ B59 B1 MU/HPI 46 Note 5)
ZW B44 B2 HPI l
i
NOTES:
1.
SHOP ASSEMBLIES WERE CLEANED TO CLASS C PER SPECIFICATION S-107 E.
2.
INSPECTION RESULTS NOMENCLATURE A.
SAFE END CRACKED, SLEEVE LOOSE / WORN / MISSING B.
SLEEVE INDICATED SOME LOOSENESS / WEAR - N0 SAFE END CRACKING C.
CIRCUMFERENTIAL CRACK OR MARK OK - NO ABNORMAL INDICATIONS 3.
INFORMATION MUST BE OBTAINED FROM SITE RECORDS 4.
INFORMATION FOR THIS MATRIX CONCERNING COLD LEG PIPE ASSEMBLY SERIAL N0'S. B41-209-50-1 AND B41-209-50-2 IS AVAILABLE BUT WHICH ASSEMBLY IS LOCATED IN THE B2 LEG AND Al LEG MUST BE OBTAINED FROM SITE RECORDS.
5.
WHILE TAKING MEASUREMENTS OF THE A-1 RC PUMP FIXED VANES, IT WAS DISCOVERED THAT THE THERMAL SLEEVE IN THE HPI LINE N0ZZLES WAS LOOSE. ALL THERMAL SLEEVES WERE REROLLED. THE FOLLOWING INFORMATION WAS RECORDED AT THE SITE.
CUST.
THERMAL SLEEVE ID IDENTIFICATION IN EXPANDED AREA A2 TH. SLEEVE TIGHT 1.5086 Al TH. SLEEVE LOOSE 1.5060 B1 TH.. SLEEVE TIGHT 1.5178 B2 TH. SLEEVE TIGHT 1.5162 THERMAL SLEEVE ID AFTER RER0LLING IN EXPANDED AREA A2 TH. SLEEVE TIGHT 1.5162 Al TH. SLEEVE TIGHT 1.5190 81 TH. SLEEVE TIGHT 1.5178 B2 TH. SLEEVE TIGHT 1.5183
)
TYPE A H = 4' 5 3/16" l = 12' 6" FOR NSS 3,4,5,9,11 TYPE B H = 4 9 3/1 s" t = 13 '
FOR NSS 7.8
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E
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5 O-e amm0
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1 z
l 1
l 5
5 h
=
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=
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)
TYPE C a
c=
f,er 5 NSS-14
=,
=
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_ _ _ _. _ _