ML20079M555
| ML20079M555 | |
| Person / Time | |
|---|---|
| Site: | Millstone  |
| Issue date: | 12/31/1983 |
| From: | NORTHEAST NUCLEAR ENERGY CO. |
| To: | |
| References | |
| NUDOCS 8401270360 | |
| Download: ML20079M555 (163) | |
Text
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'N DOCKET NO. 50-336 LICENSE NO. DPR-65 NORTHEAST NUCLEAR ENERGY COMPANY MILLSTONE NUCLEAR POWER STATION UNIT NO. 2 4
4 THERMAL SHIELD DAMAGE RECOVERY PROGRAM FINAL REPORT
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December 12,1983 Docket No. 50-336 B10968 Director of Nuclear Reactor Regulation Attn: Mr. James R. Miller, Chief Operating Reactors Branch #3 U. S. Nuclear Regulatory Commission Washington, D. C. 20$55
References:
(1) W. G. Counsil letter to J. R. Miller, dated September 15, 1983. (2) E. J. Mroczka letter to T. E. Murley, dated July 1,1983. Millstone Nuclear Power Station, Unit No. 2 Thermal Shield Damage Recovery Program Final Report The Millstone Unit No. 2 nuclear steam supply system vendor informed Northeast Nuclear Energy Company (NNECO) in late March,1983 of reactor internals degradation at a plant similar in design and operation to Millstone. Specific details of the degradation in the thermal shield and thermal shield support system were disseminated to NNECO and other licensees at a meeting with the NRC Staff on April 12,1983. At that time, NNECO informed the Staf f of plans for an extended refueling and maintenance outage commencing on May 28, 1983 during which detailed examination of the reactor vessel, and its internals, was scheduled. i l The core support barrel and thermal shield assembly were removed from the l reactor vessel on June 30,1983 and at that time it was apparent that darr. age to l the thermal shield support system had occurred. Preliminary notification of this damage was made by Reference (1). l i Representatives of NNECO met with the Staff on July 21, 1983 to discuss the reactor vessel internals inspection results and to outline preliminary plans supporting return to power operation. The thermal shield damage necessitated the removal of the component from the core support barrel, and evaluations of operation without the thermal shield. Reference (2) elaborated on the plans and l schedules for removal of the thermal shield and return to power operation. The analytical ~ef forts planned to support continued operation were also outlined. Af ter removing the thermal shield from the core support barrel, NNECO performed extensive nondestructive examination of the core support barrel to determine the extent, if any, of damage to the component. These inspections were completed on October 30, 1983. NNECO met with the Staf f on November 9,1983 to present the results of the core support barrel inspections and to
outline che plans and schedules for return to power operation of Millstone Unit No.2. At that time, we also outlined to the Staff the documentation which NNECO would provide before startup. The attached report provides the basis and supporting information for return to power operation of Millstone Unit No. 2 and fulfills our commitment made to the Staff at the November 9,1983 meeting. A d. aft of this report was provided to the Millstone Unit No. 2 Project Manager on November 23,1983. The contents of the report include the following: Chapter 1 Summary and Chronology of Events Chapter 2 Reactor Internals Review Chapter 3 Pressurized Thermal Shock Chapter 4 Nondestructive Examination Techniques Chapter 5 Nondestructive Examination Inspection Results Chapter 6 Failure Mechanism Analysis Program Chapter 7 Reactor Internals Stress Analysis and Core Support Barrel Structural Integrity Chapter 8 Safety Analyses Chapter 9 Inspection and Monitoring Information has been included in the attached report which was not available for presentation at the November 9,1983 meeting. Specifically, during a review of the examination results, an additional non-through-wall crack was identified on lug 4 of the core support barrel as discussed in Chapter 5 of the attached report. NNECO reinspected all thermal shield support lugs with ultrasonics to verify the inspection results provided to the Staff on November 9,1983. No additional indications were identified. The non-through-wall crack has been repaired as discussed in Chapter 7. Chapter 6 provides a discussion of the failure mechanism analysis program, in which NNECO has participated with the Florida Power and Light Company. The similar design and operating characteristics between Millstone Unit No. 2 and St. Lucie Unit 1 permit the application of the pertinent results and conclusions of the analysis performed for St. Lucie Unit I to Millstone Unit No. 2. Although some plant specific evaluations are ongoing, NNECO does not expect the results to alter the conclusion documented in Chapter 6 regarding the failure mechanism attributable to the Millstone Unit No. 2 thermal shield damage. _ _ __._ m_ _ _ _ _ _ _ _ _ _ _ _
Chapter 8 of the attached report discusses potential effects on the fuel from additional bypass flow which will result from the core support barrel repairs. Flow tests and vibration analyses have been conducted to confirm the integrity of the fuel under the flow conditions expected as a result of the core support barrel repairs. A detailed test report will be submitted as Appendix A to this document to supplement the information provided in Chapter 8. NNECO expects to provide this information to the Staf f by the end of this year. As was documented to the Staff on November 9,1983 no license amendments are required to support plant startup without a thermal shield and with the core support barrel as repaired. NNECO recognizes the fact that current heatup and cooldown curves of Section 3/4.4.9 in the Technical Specifications will no longer be applicable for the time period specified due to increases in neutron fluence at the reactor vessel without a thermal shield. Revised curves have been prepared as discussed in Chapter 3 and a license amendment will be requested to incorporate the new curves into the Millstone Unit N o. 2 Technical Specifications. It is NNECO's position that the analyses presented herein include the appropriate conservatisms and that continued operation of Millstone Unit No. 2 is prudent and justified. The attached report is currently being reviewed pursuant to 10 CFR50.59 with no unreviewed safety question identified to date. The Millstone Unit No. 2 Plant Operations Review Committee (PORC) and Nuclear Review Board (NRB) have also reviewed the attached report and concur with this determination. As such, NNECO is providing this document for the Staff's j information. A final determination regarding 10CFR50.59 will be docketed by i December 22,1983. The current date for plant heatup is January 3,1984. Refueling of the reactor is presently ongoing to support this schedule. l l We trust you find this information satisfactory. As is always the case, my Staf f remains available to assist you in any way in this matter. Very truly yours,
- NORTHEAST NUCLEAR ENERGY COMPANY s A W. G. Counsil i Senior Vice President
i NORTIMAST UTltJTIES Genmi OHiceo Se Shut. Betin. Connecticut w ca.new w .om co .'" P.O. BOX 270
.a.oa ..m om co. HARTFORD. CONNECTICUT 06141-0270 g g g aw aryyco, (203) 666-6911 December 28,1983 Docket No. 50-336 B10981 Director of Nuclear Reactor Regulation Attn: Mr. James R. Miller Operating Reactors Branch #3 U. S. Nuclear Regulatory Commission Washington, D. C. 20555
References:
(1) W. G. Counsil letter to 3. R. Miller, dated December 12, 1983. Gentlemen: Millstone Nuclear Power Station, Unit No. 2 Thermal Shield Recovery Program Northeast Nuclear Energy Company (NNECO) provided the NRC Staff in Reference (1) the final report detailing the thermal shield damage recovery program for Millstone Unit No. 2. At the time NNECO docketed the Reference (1) report, NNECO had not submitted its final determination with respect to the ' criteria delineated in 10 CFR 50.59(a)(2). Pursuant to 10 CFR 50.59, NNECO has completed the review of the plant design change to remove the thermal shield and repair the core support barrel as described in Reference (1). It has been concluded that no unreviewed safety question results from the modifications at Millstone Unit No. 2. This is consistent with the preliminary findings reported in Reference (1). NNECO is also providing additional information regarding the results of the fracture mechanics evaluation of the flaw tolerance for the core support barrel. This information is provided as revised Section 7.3.5 to be incorporated into the final reports provided to the Millstone Unit No. 2 Project Manager on December 12,1983. In addition, Section 7.2.1 and Tables 7.2-2 and 7.2-3 have been revised and updated to include values not provided in previous transmittals. Regarding the potential effects of the thermal shield removal on reactor internals vibration, the following information is provided. As discussed in Section 8.2 of Reference (1), the removal of the thermal shield from the reactor core support barrel will result in a net decrease in reactor vessel pressure drop which in turn will result in a small increase in reactor coolant system flow rate. NNECO estimates this increase in flow will be less than one percent (1%) of reactor coolant system flow rate. The increase in reactor coolant system flow rate due to this modification offsets, to a small degree, the decrease in system flow experienced as a result of plugging and sleeving steam ger.crator tubes.
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. NNECO has currently plugged over 1600 steam generator tubes (~9% of the total number of steam generator tubes) and sleeved approximately 2000 tubes. These changes have resulted in a decrease of approximately 3.6% in the reactor coolant
- system flow rate as measured during initial startup testing at Millstone Unit No.
- 2. The increase- in flow expected from removing the thermal shield is approximately one percent for a net decrease of approximately 2.6% in reactor coolant flow since initial plant startup.
The core inlet flow distributions have been evaluated for plants similar in design to Millstone Unit No. 2 which do not have thermal shields. The reactor internals . design, namely tr.e flow skirt, lower support structure bottom plate and core > support plate, ' effectively flattens the reactor coolant flow as it exits the downcomer annulus insuring a uniform core inlet flow distribution. A review of reactor internals vibration monitoring of plants without a thermal j shield whose reactor coolant systems are geometrically identical to Millstone Unit No. 2 and whose measured reactor coolant flow rates were greater than i that expected at Millstone Unit No. 2 has been completed. The review did not 4 identify any abnormal reactor internals vibration. NNECO concludes that the removal of the thermal shield from the Millstone Unit No. 2 core support barrel is not expected to result in reactor internals - vibration from the resulting change in flow characteristics, either flow rate or distribution. Furthermore, the thermal shield remvval will most likely remove a suspected source of internals vibration as was ' discussed in Chapter 6 of Reference (1). We trust you find this information satisfactory. Very truly yours, NORTHEAST NUCLEAR ENERGY COMPANY i W. C. Counsil L Senior Vice President i l l
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. 1 NORTHEAST NUCLEAR ENERGY COMPANY MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2 THERMAL SHIELD DAMAGE RECOVERY PROGRAM - -
FINAL REPORT CONTENTS CHAPTER 1
SUMMARY
AND CHRONOLOGY OF EVENTS CHAPTER 2 REACTOR INTERNALS REVIEW CHAPTER 3 PRESSURIZED THERMAL SHOCK CHAPTER 4 NONDESTRUCTIVE EXAMINATION TECHNIQUES
' CHAPTER 5 NONDESTRUCTIVE EXAMINATION INSPECTION RESULTS -
CHAPTER 6 FAILURE MECHANISM ANALYSIS PROGRAM > CHAPTER 7 REACTOR INTERNALS STRESS ANALYSIS AND CORE SUPPORT BARREL STRUCTURAL INTEGRITY CHAPTER 8 SAFETY ANALYSES CHAPTER 9 INSPECTION AND MONITORING i . I > L-i i
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TABLE OF CONTENTS CHAPTER 1 SECTION SUBJECT 1.0
SUMMARY
AND CHRONOLOGY OF EVENTS 1.1
SUMMARY
1.2 CHRONOLOGY OF EVENTS d 1.2.1 IDENTIFICATION OF PROBLEM 1.2.2 VISUAL EXAMINATION 1.2.3 THERMAL SHIELD REMOVAL EVALUATION i 1.7.A REACTOR INTERNALS INSPECTION 1.2.5 CORE SUPPORT BARREL EXAMINATION 1.2.6 CORE SUPPORT BARREL REPAIR l 1.2.7 CORE SUPPORT BARREL STRUCTURAL EVALUATION i l i l l 8
4 17/5/R3 1.0 SUMARY At!D CHRONOLOGY OF EVENTS l 1.1 SUMARY Northeast Nuclear Energy Company (NNECO) comenced a refueling out' age on May 28, 1983, at Millstone 2. The original schedule of maintenance 1 and refueling activities included steam generator channel head decontamination and tube sleeving. The additional outage time needed to support these activities afforded the opportunity to conduct the ten-year inservice inspection of the reactor coolant system and its associated systems. The ten-year inservice inspections included the reactor vessel and its interface points with the reactor internals components. NNECO was.aade aware of reactor internals degradation at two plants having similar internals configuration to Millstone 2. The degradation involved the reactor vessel thermal shield and its support system. At the time this information was disseminated, NNECO met with the NRC staff along with the other potentially effected owners with plants having similar components. As a result of the information on thermal shield support system degradation and the implications to Millstone 2, the planned reactor internals inspections were expanded to include additional detailed examinations of the thermal shield support system. The core support barrel and thermal shield assembly was removed from ! the reactor vessel on June 30, 1983, at which time it became evident that the thermal shield and its support system had experienced damage. Subsequent evaluation led to a decision to remove the thermal j shield. 1.1-1
l. 12/5/83 damaged areas have been inspected using nondestructive examination techniques, and repair methods have been formulated to insure core support barrel integrity. An analysis of the repaired barrel has been completed which shows that the original design criteria are met; " stress levels remain within those allowed by the American Society of Mechanical Engineers (ASME) Code, Section III,1977 edition plus addenda through Summer 1978 A e e 1.1-2
r , . 12/5/83
-1.2 CHRONOLOGY OF EVENTS 1.2.1 Identification of Problem During the refueling outage which connenced in May,1983, the 10-year in-service inspection was perforned for the reactor vessel and its internals. The thermal shield and its support system were of particular interest since another unit of similar design had identified damage to its thermal shield and support system in March, 1983.
1.7.2 Visual Examination A visual examination of the core support barrel / thermal shield assembly disclosed the thermal shield support system to be damaged. A few thermal shield support pins were damaged and/or missing and damage 1 to the thermal shield was visible. Pieces from the thermal shield were found in the reactor vessel and two positioning pins which had become dislodged were found in the reactor vessel. 1.2.3 Thermal Shield Removal Evaluation i An evaluation of the thermal shield support system concluded that refurbishment was impractical. A decision was made to remove the thermal shield. Analyses were performed to evaluate operation of the plant without a thermal shield for its remaining i design life. They indicated that replacement of the thermal shield I was not necessary. I 1 1.2-1 l
!~ 12/5/R3 1.2.4 Reactor Internals Inspection l The reactor internals interfaces within the reactor vessel were examined and were found not to exhibit evidence of excessive vibration of the reactor core barrel and upper guide structure. A visual examination of a selected number of fuel assemblies did not disclose detrimental effects attributable to the degraded thermal shield support system. From these inspections it was concluded that the damage was confined to the core support barrel and thermal shield. 1.2.5 Core
- Support Barrel Examination Upon remov.a1 of the thermal shield from the core support barrel, a nondestructive examination of the core support barrel was conducted.
The examination consisted of visual, eddy current and ultrasonic inspection techniques. Danage was in evidence at two of the nine lug locations. Throughwall cracks were confirmed a(iacent to two damaged lug locations. 1.2.6 Core Support Barrel Repat r-A repair process for the core support barrel was formulated. f Underwater machining of the core support barrel in the damaged areas was used to reduce stress concentrations. Through-wall cracks were l relieved by crack arrestor holes; non-through-wall cracks were removed by machining. l l I . i l l 1.2-?
l 12/5/83 1.2.7 Core Support Barrel Structural Evaluation A structural evaluation of the repaired core support barrel and the reactor internals without the thermal shield was performed. The component stresses under normal operating, fatigue, seismic, ard loss of coolant loads were evaluated and found to be within the limits of Section III, Subsection NG of the ASME Code. i l 1.2-3 l l l t
I 12/5/83 I l 1 TABLE OF CONTENTS CHAPTER 2 1 SECTION SUBJECT
2.0 DESCRIPTION
OF THE REACTOR INTERNALS AND REACTOR VESSEL INTERFACES
2.1 INTRODUCTION
2.1.1 CORE SUPPORT ASSEMBLY 2.1.2 CORE SUPPORT BARREL 2.1.3 CORE SUPPnRT PLATE AND SUPPORT COLi!MNS 2.1.4 THERMAL SHIELD AND THERMAL SHIELD SUPPORT SYSTEM (PRE-REMOVA 2.2 UPPER GIIIDE STRUCTURE ASSEMBLY 2.2.1 UPPER GUIDE STRUCTURE SUPPORT PLATE 2.2.2 CONTROL ELEMENT SHROUD ASSEMBLIES 2.2.3 FUEL ASSEMBLY ALIGNMENT PLATE 2.2.4 HOLDDOWN RING 2.3 REACTOR VESSEL INTERFACES 2.3.1 FLOW SKIRT 2.3.2 CORE STOPS ? _. _. - - - . . . _ _ .-.- - -. . - - . - . - . . . - - - - .- - - - - - - - - - -
2.0 nESCRIPTIDF 0F THE REACTOR INTERNALS AND REACTOR VESSEL INTERFACES
2.1 INTRODUCTION
This chapter describes the reactor internals and interfaces prior to damage and the subsequent repairs. The components of the original reactor internals are shown in Figure 2.1.1. A more detailed figure showing the core support barrel and connecting structures are shown in Figure 2.1.2. The structure is divided into two major components consisting of the core support barrel assembly and the upper guide structure assembly. The flow skirt and core-stops, although in the coolant flow path, are separate from the internals and are welded to the bottom head of the pressure vessel. 2.1.1 Core Support Assembly The core support is the major support member of the reactor internals assembly. The assembled structure consists of the core support barrel, the lower support structure, the core shroud, and the thermal shield. The major material for the assembly is Type 304 stainless steel. The core support assembly is supported at its upper end by the upper flange of the core support barrel which rests on a ledge in the reactor vessel flange. The lower flange of the core support barrel supports and positions the lower support structure. The lower support structure provides support for the core by means of a core support plate supported by columns resting on beam assemblies. The core support plate provides support and orientation for the fuel assemblies The core shroud which surrounds the fuel assemblies is also 2.1-1
u, af o.- supported by the core support plate. The lower end of the core support barrel is restrained laterally to the pressure vessel by six snubbers (Figure 2.1-3). 2.1.2 Core Support Rarrel The core support barrel is a right circular cylinder with a nominal inside diameter of 14R inches and a minimum wall thickness of 1-3/4 inch. It is suspended by a A-inch thick flange from a ledge on the pressure vessel. The core support barrel supports the lower support
- structure upon which the fuel assemblies rest. Four alignment keys located 90 degrees apart, are press fitted into the flange of the core support barrel slots in the reactor vessel, closure head, and upper guide structure assembly flanges at locations corresponding to the alignment, key locations and prwide proper alignment between these components in the vessel flange region.
The core support barrel is about 27 feet long. Snubbers are installed on the outside of the core support barrel near the bottom end, as shown in Figure 2.1-3. The snubbers consist of six equally spaced double lugs around the circumference and are the growes of a
" tongue-and-growe" assembly; the pressure vessel lugs are the tongues. The small clearance between the two mating snubber surfaces between the core support barrel and the reactor vessel is accomplished l by machining to as-built dimensions determined during initial field i insta11atien. nuring assembly, as the core support harrel is lowered
- into the vessel, the pressure vessel snubber lugs engage the core l support harrel snubber growe in the axial direction. Radial and l
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axial expansion of the core support barrel is accommodated, but 1ateral movement of the core support barrel is restricted. The pressure vessel snubber lugs have bolted, Inconel shims and the core i support barrel snubbers are hardfaced with Stellite to minimize wear. 2.1.3 Core Support Plate and Support Columns The core support plate is a 147-inch diameter 2-inch thick, Type 304 stainless steel plate into which the necessary flow distributor holes I for the fuel assemblies have been machined. Fuel assembly locating pins (fcur for each assembly) are shrunk-fit into this plate. Columns and support beams are located between this plate and the bottom of the
, core support barrel in order to provide support for this plate and transmit the core _ load to the bottom flange of the core support barrel. -
2.1.4 Thermal Shield and Thermal Shield Support System (Pre-Removal) The thermal shield is a 3-inch thick, 304 stainless steel cylindrical structure with an inside diameter of 156-3/4 inches and a height of l 137-3/4 inches (see Figure 2.1-2). The thermal shield is supported at i the top by nine equally sp',ced a support lugs welded to the outer periphery of the core support barrel. Support pins are fit during assembly to position the thermal shield on the support lugs. This is shown in Figure 2.1-4 The support pins are welded to the thermal shield and have a .0005 to .002 inch clearance on the sides to permit relative thermal expansion of the core support barrel and thermal shield. The thermal shield is positioned radially by a total of twenty-six positioning pins. Nine j of the pins are located approximately 15 inches below the top of the 2.1-3
support lugs and the remaining seventeen positioning pins are located approximately 21-1/4 inches from the bottom of the thermal shield. The positioning pins thread into the thermal shield and are preloaded against the core support barrel. Lock bars are inserted into slots in the positioning pins and lock welded to the thermal shield. d e I
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12/5/R3 2.2 UPPER GUIDE STRilCTURE UGS ASSEMBLY The UGS assembly, Figure 2.2-1, consists of the upper guide structure plate control element assembly shrouds, a fuel assembly alignment plate and a holddown ring. 2.2.1 Upper Guice Structure Support Plate The upper end of the assembly is a structure consisting of a support plate welded to a grid array of 24-inch deep beams and a 24-inch deep cylinder which encloses and is welded to the ends of the beams. The periphery of the plate contains four accurately machined and located alignment keyways, equ11y spaced at gn-degree intervals, which engage a the core barrel alignment keys. The reactor vessel closure head flange is slotted to engage the upper ends of the alignment keys in the core barrel. This system of keys and slots provides a means of aligning the core with the closure head. The grid aligns and supports the upper end of the CEA shrouds. 2.2.2 control Element Shroud Assemblies The control element assembly shrouds extend from the fuel assembly alignment plate to an elevation about 3 feet above the support plate. There are single and dual type shrouds. The shrouds are bolted to the fuel assembly alignment plate. At the upper guide structure support plate, the single shrouds are connected to the plate by spanner nuts which permit axial adjustment. The spanner nuts are tightened to i proper torque and lock welded. The dual shrouds are attached to the upper plate by welding. I 2.2-1 l l
l 12/5/83 i.- 2.2.3 Fuel Assembly Alignment Plate The fuel assembly alignment plate is designed to align the upper ends of the fuel assemblies and to support and align the lower ends of the CEA shrouds. Precision machined and located holes in the fuel assembly alignment plate aliga the fuel assemblies. The fuel assembly alignment plate has four equally spaced slots on its outer edge which engage with Stellite hardfaced lugs protruding from the core shroud to limit lateral motion of the upper guide structure assembly. 2.2.4 Holddown Ring The holddown ring consists of a martensitic stainless steel ring approximately 2-1/4 inches thick with an inside diameter of 156 inches and an outside diameter of 16R-1/2 inches. The botton surface of the holddown ring at the inner edge rests on the flange of the upper guide structure. The top surface of the holddown ring at the outer edge contacts the seating surface of the reactor vessel head. The holddown ring is aligned to the upper guide structure flange by the four alignment keys protruding from the core support barrel flange.
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12/5/81 2.3 REACTOR VESSEL INTERFACES The remaining interfaces with the reactor vessel not previously described are the flow skirt and core stops. 2.3.1 Flow Skirt The flow skirt (see Figure 2.1-1) is a perforated right circular cylinder that is attached to the reactor vessel by welding. The flow skirt material is Inconel. 2.3.2 Core Stops Nine equally spaced core stop lugs are welded to the reactor vessel and are located at the periphery, but below the core support barrel lower flange. The core stops limit the vertical drop of the core support barrel assembly to approximately one inch in the event of a postulated failure. l l 1 l l l l l 2.3-1 l l [
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TARLE OF CONTENTS CHAPTER 3 SECTION SilRJECT 3.0 PRESSt!RIZE0 THERMAL SHOCK 3.1 CALCtlLATION OF NEllTRON FLUENCE 3.1.1 DISCUSSION 3.2 VESSEL MATERIAL PROPERTIES ANALYSIS 3.2.1 CHEMISTRY AND TOUGHNESS PROPERTIES 3.2.2 ADJilSTED RTNDT PREnICTION METHODOLOGY 3.2.3 PRECTCTED ADJilSTED RTNDT 3.3 CALCl1LATION OF ENERGY.0EPOSITION RATE DISTRIBUTIONS IN VESSEL 3.A CALCilLATION OF VESSEL WALL TEMPERATURE DUE TO HEAT GENERATI0s 3.5 EFFECT OF GAMMA HEATING ON PTS RESPONSE 3.5,1 METHOD OF EVALUATION l 3.5.2 RESULTS l 3.5.3 CONCLilSION , 3.6 NEllTRON FLIIX MONITORING t j 3.7 TECHNICAL SPECIFICATIONS
i4 3 A r./ :)/ n.1 3.0 PRESSURIZED THERMAL SHOCK
, y. .
i 3.1 CALCULAT1DN OF t'EbiRON FLUENCE s The RTNDT, which is an index of the degme of embrittlement of the material, is determined as a function of radiation using the fast neutron (En > 1.n MeV) fluence. For the evaluation of the Millstone 2 reactor pressure vessel, the vessel fluence distribution was calculated using the Sn technique as. incorporated into the DDT 4.3 computer code (Reference 1). Azimuthal, fluence profiles were computed both with and without a thermal shield. The axial profile of the fluence was determined in Reference 6.a A long-term' average axial peak of 1.1A was used. The effect of holes in the core support, barrel on the vessel fluence d,istribution was expressSd in ter:ns of a localized penetration to the fluence distribution corresponding to a core support barrel with no holes. The results of theiD' OT calculations were normalized to the value of the. fluence ' determined 190m the evaluation of surveillance capsule W-97 reported in Reference S. ~ 3.1.1 DISCUS Ob , The three-dimensional fast neutron fluence distribution was synthesized from tie separate results of 'the RQ-00T and RZ-DOT calculation. This syr. thesis relies!on the separability of the distribution into azinuthal add' axial parts basr:d on the existing symetry in the geometric model of s t'he core to reactor vessel system. The RD geometric model assumed l ' eighth core synmetry in the azinuthal dimension. The RZ model extended from 200 centimeters (78.7 inchesi below to 200 centimeters (78.7 inches) above the active core mio-plane. Figure 3.1-1 shows a sketch of the RD (geometric model . The RZ model used the same radial geometry except that t e core is cylindricalized to a radius of 172.,7 centimeters (6R inches). The DDT 4.3 calculations used so Sg approximation for the S n hadrature and/J P3 representation of the order of scatter in the a 1 1 l , i 3.1-1 l l
-- - L . . _ . - - . . - - - - . - - - . - ._
12/5/R3 l energy group-to-group transfer cross section. The radiation transport cross sections were based on the DLC-23 " Cask" cross section library. ( The neutron energy range from 15.0 MeV to thermal energies is represented by 22 energy groups in this library. The macroscopic cross sections were calculated for use in DnT 4.3 with the GIP (Reference 2) computer code.
.The fast neutron (En > 1.0 MeV) fluence was determined using a response -
function tailored to the DLC-?3 energy group structurc. Since the DLC-23 { cross section set does not have an energy group boundary at 1.n MeV an approximate 1/E spectrum was assumed for the neutron spectrum within the
; energy group that includes the value of 1.0 MeV. The fraction of the spectrum above 1.0 MeV in this group was then multiplied by the flux value for the entire group to obtain the contribution of this group to the En > 1.n flux. Below this group the neutron flux values are not included in the fast neutron flux. Above the break-point group, the flux values for each group were summed to obtain the fast flux.
The fixed sources for the R0 DOT calculations were based on time averaged pin power distributions calculated with PD0 (Reference 3) for a fuel management strategy similar to that used in Millstone 2. These calculations represent an approximation to the time-averaged power distribution expected during the lifetime of the reactor. The assembly average power corresponding to the time averaged source distribution is shown in Figure 3.1-3. The effect on the vessel fluence of the thermal l shield removal after the end of cycle 5 was determined based on the comparison of the vessel flux results of two RO DOT calculations. The RO DOT models were identical except for the removal of the thermal shield in one of the monels. The removal of the thermal shield was calculated to cause an increase in the peak flux on the vessel by a factor of 1.73. The thermal shield worth, with respect to fast neutron flux incident on the reactor vessel, is a function of azimuthal position on the vessel and the above ratio applies to the peak flux location being at the n' (opposite the centerline through the core flat) location. i 1.1-2 6 l l
. . . . - ,. . ..-- _.__., . . . , . . _ . . ~ , _ . , , _ . . , . . . - . _ , _ _ . . _ _ - _ ~ - , _ . _ . . . _ _ _ . _ _ _ . _ . . _ . .
l
.. .. 12/5/R3 The fixed sources for the RZ-DOT calculations were based on an approximation to the expected time averaged axial power distribution for the perpheral assemblies in the reactor core. Figure 3.1-4 shows the axial power distribution used in the DOT RZ model. The power distributions were normalized such that the fixed source distributions in the RO DOT calculations represent a plane in the core with a plane average to core-average power ratio of unity.
The normalization of the results of the combined R0 and RZ DOT calculations was obtained from the ratio of the average fast flux calculated from the measured activi':les at the vessel wall dosimeter location to the fast flux calculated for the same position with the DOT R0 and RZ models for the operation from Bnl to the E003, i.e., the time period of exposure for the surveillance capsule. The peak flug and fluence values resulting from the R0 and RZ DnT calculations are summarized in Tables 3.1-1 and 3.1-2. The time averaged azimuthal distribution of the fast neutron fluence for the 32 EFPY irradiation time case are shown in Figure 3.1-5. The axial distribution for the same case is shown in Figure 3.1-6. The results discussed thus far apply to geometric models which assume a uniform core support barrel. . In order to account for the perturbations caused by the crack arrestor h' oles in the core support barrel, an additional R0 DDT calculation was performed to develop a fast flux adjustment factor for the vessel flux distribution. The location and site of the holes in the core support barrel invalidate the separability assumption of the axial and azinuthal fluence distributions on the reactor vessel. The effect of an individual hole in the core support i barrel on the reactor vessel fluence distribution was considered as a local perturbation to the three-dimensional distribution calculated for j the uniform core support barrel situation. l l 3.1-3 l l
In order to calculate the adjustment to the localized vessel fluence distribution, two 00T 4.3 calculations were performed using the same models except that one of the models included a hole in the core support barrel as shown in Figure 3.1-7. The ratio of the neutron fast flux at the vessel-clad interface for the calculatica with a slotted core support barrel to the fast flux for a uniform core support barrel provided the desired adj ustment factors. Figure 3.1-8 shows the adjustrent factor as a function of distance from the center of the hole for a 6.6 inch diameter hole. For smaller holes, this adjustment factor provides a conservative estimate of the increase in the vessel flux due to the placement of the hole in the core support barrel. The adjustment factor was applied based on the distance from the hole centerline to the inner surface of the reactor vessel. Each hole was considered as independent of the other holes and the total adj usted fluence was obtained in a simple sum of the fluence adjustments due to perturbation of each hole in the core support barrel. l l i i 1.1 A l l l l
, . . . 12/5/83 REFERENCES (SECTION 3.1)
- 1. Rhoades, W. A., and R. L. Childs, An Updated Version of the DOT 4 One and Two-Dimensional Neutron / Proton Transport Code, ORNL-5R51, Oak Ridge National Laboratory, Oak Ridge, Tennessee, April,1982.
- 2. Rhondes W. A., The GIP Program for Preparation of Group Organized Cross Section Libraries, RSIC Computer Code Collection - 320, p. 94
- 3. PD0 - Caldwell, W. R., "PDO-7 Reference Manual" WAPD-Tm-678, January 1968
- 4. MACKLIB IV, DLC 60B, Radiation Shielding Information Center, Oak Ridge National Laboratory, Oak Ridge, Tennessee.
- 5. C-E Report #TR-N-MCM nnR, NUSCO Millstone Nuclear Unit Number 2 Evaluation of Irradation Capsule W 97. April,1982.
6 C-E Report aCEN-189 Appendix E, Fvaluation of Pressurized Thermal Shock Effects Due to Small Break LOCA's with Loss of Feedwater for the Millstone 2 Reactor Vessel, December,1981. 3.1-5 l
12/5/01 l l 1.2 VESSEL MATERTALS PRnPERTTFS ANALYSTS This ioction reports the predicted '?EFPV DTFDT values for each plate and weld in the vessel beltline region. The maximum predicted RTPnT at EOL for a vestal weld was POS*F in girth weld 9-?n1 which j oins the lower and intermediate shell courses. The limiting plate was C ;05 ? This plate is in the intermediata shall with an E0L RTFDT of 107 F. course. The locations of tne cora support barrel repairs affect plate C-509-3 which, with the local fluence peak will still laq the RTNDT of plate C-505-2. 3.7.1 Chemistry and Toughness Properties The pre irradiation properties of tha reactor vessel plates and welds are sumnarized in Table 1.? 1 The chemistry data are the same as those reported in Reference 1 4 A " map" of the cylindrical portion of the Millstone tinit ? reactor vessel is given in Figure 3.?-1. it shows the locations of the plates and welds listed in Table 1.7-1 and their c.ncrasponding valuas of initial RT NnT(F*\ located within a rectangle on the figure. RT NnT values for the vertical weld seams (designated 1 ?nt,? ?ni, and 3-7035 are shown at a single seam but apply to all three vertical seams ir a given shell course. Included in the figure are the locations of the inlet and outlet nozzles, the core midplane, and the extrenities of tha active core. 1.2.2 Ad.iusted RT t'nT prediction Methodology The methodology used for predictin'g the neutron radiation induced changes in toughness properties of the Millstone tinit 2 reactor vessel J 3.2-1
, - - - , ,m~ , ~,, . - , - , - e--,.n., ,, e,- .r---
1?/5/R3 is the same as that used in the NRC staff evaluation of pressurized thermal shock (Reference 3) as summarized below. Adj usted RTNDT = RTNDT ^ ARTNDT + 2 I U where: RTNDT = reference transition temperature RTNDT = mean initial RTNDT ARTFDT = mean radiatien induced shift in RTNDT o = standard deviation , The value of RTNDT to be used is one of the following: ai Actual value in accordance with NR-??00, ASME Code; bi Calculated value based on Branch Technical Position MTER ; 7,
" Fracture Toughness Requirements for older plants";
c) Estimated value based on generic data. All three methods were used for the Millstone tinit 2 vessel materials as indicated in Table 1.?-1. In accordance with the FPC guidelines, the lower value of the following two expressions is used for the predicted adjusted RTNnT: Acij . RT N nT
- Nt 'DT + ( 10 + 47n cu + 1;n CuNii.
l9 30.77 p [g32 ,$2)1/2 [4/10 i Adj . RTt 'DT " IIFnT + 283 [o /1019'0* * + 2o1 l l 3.2-2
1?/E/R1 where: Cu, Ni = weight percent copper and nickel, respectively
& = neutron fluence, E>1MeV c4 = standard deviation for initial RT NDT (0 for initial RT NDT based on actual data, 17'F for estimated RT NDT for generic welds) o = standard deviation for RTNDT prediction, ?A*F.
Note that the second expression was not found to be controlling for the Millstone tinit 2 vessel materials. 1.2.1 Predicted Adiusted RT- NDT Fiqure 1.2-2 and Tahle 3.2-? show the predicted 17 EFDY adj usted RT HDT valuas at the inner surface of tha Millstone linit 2 reactor vessel. The vessel neutron fluence at 1? effective full power years at 7700 Mwt used in the predictions was MxinlUn/cm2 (E>1.MeV). This value reflects operation with the thermal shield in place during the first five cycles and operation with the thermal shield removed for the remainder of plant life to 37 EFPY. The initial RT NDT and chenistry data from Table 3.?- 1 and the azimuthal and axial flux profiles from Section 3.1 were used to generate the adj usted RTNDT values; differences in relative azimuthal and axial flux profiles before and after thernal shield removal as well as local peaks in the repair regions were taken into account in the fluence calculations. The values of adj usted RTNnT are located in rectangles adjacent to the plate and weld sean designations. The RTyp7 values apply to the inner surface of the reactor vessel. l The maximum adj usted RTNtiT at 19 FFPY is ?n5'F in girth weld a-?n3 which j oins the internadiate and lower shell courses. This value, and the highest value for plate material is well helow the criteria established in Reference 3 for PTS cnnterns. 3.2-2
~ . - .
17/5/R1 The holes machined in the core support barrel are predicted to result in about a 2A" increase in neutron flux at the inside surface of the reactor vessel adj acent to the holes; 1.e., at restricted locations in the intermediate shell course plate C 605 1 No welds are directly adjacent to the core support barrel holes; the nearest weld seam (?-203\ will be exposed to no increase in flux. Plate C 505-3 will have an EOL RTNOT of 170 F, and will always be bounded by the 'ead plate, C-505-2, and weld 0-201 I l 3.2 2 4
- -- - , - - - - - - . . - , , , = - - , . . , , . , - -e, , , --- , - .--.
F e..,..,.... REFERENCES (SECTION 3.2)
- 1. " Evaluation of Pressurized Thermal Shock Effects due to Small Break LOCA's with Loss of Feedwater for the Reactor Vessels", CEN-189; Appendix E, December 19RI.
- 2. " Evaluation of Pressurized Thermal Shock Effects due to Small Break LOCA's with Loss of Feedwater for the Combustion Engineering NSSS", CEN-189 December 1981.
- 3. "NRC Staff Evaluation of Pressurized Thermal Shock", November 19R2 (SECY-R2465).
1 4 l l l \ 3.2- 5 i
3.1 Calculation of Energy Deposition Rate Distributions in Vessel The rate of energy deposition due to radiation interactions with the material in the reactor vessel acts as a spatielly dependent source term in the heat transfer equation used to calculate the temperature distribution in the reactor vessel. A mgj or portion of this source term is caused by gamma ray interactions with the reactor vessel, however, some of the energy deposition rate (e.d.r.) is caused by neutrons. The gamma ray component can be sub-divided into portions from prompt fission, delayed fission, fission products, activation gamma rays and gamma rays due to non-fission neutron interactions with the material enclosed by the reactor vessel. In the outer region of the reactor vessel some of the e.d.r. would be caused by gamma rays due to neutron interactions with material outside the reactor vesset. The Sn method was used for the calculation of the reactor vessel energy deposition rate distributions in this analysis. RD (azimuthal) and RZ (axial) geometry models were constructed for application with the DOT 4.3 computer code. A three-dimensional synthesis of the e.d.r. distribution in the reactor vessel was obtained by combining the results of the axial and azimuthal calculations. The geometric models are the same as those used in the vessel fluence distribution 1 calculations and are shown in Figure 3.1-3 of Section 3.1. The fixed source was represented by the power distribution shown in terms of assembly average powers in Figure 3.1-3. The calculations used a more detailed spatial representation of the power distribution based on a transformation of pinwise edits of PD0 calculated i power distributions. For the EnR calculations an adjustment was made to l remove the effect of boron in the coolant in order to account for end of l cycle conditions. ( In order to correctly account for the ganma rays other than fission and i fission product gamma rays, the spatial and energy distribution of the a j 3.3-1 l l i
lS *
- 12/5/83 i
! neutron population must also be determined. The calculation of both the neutron and gama ray distributions was accomplished using the S n I method 0 logy with an 58 approximation to the quadrature set and a P3 representation of the energy group to energy group transfe! ross sections. The coupling between the neutron and gama ray distributions was accomplished using the energy group transfer matrix. cross sections as the link between neutron and gama ray energy groups. The DLC-23E cross section set was applied to this calculation using the first 22 energy groups of the set to account for the neutron energy range between 15.0 MeV and thermal energies. The last 18 energy groups of the set covered the gama ray energy range of 10.0 to 0.05 MeV. Once the neutron and gamma ray energy dependent flux distributions were , obtained,"they were converted to energy deposition rate distributions using the KERMA factor approximation and the MACKLIB IV kerma factor library collapsed to the 22 neutron and IR gama ray energy group structure of DLC-23E. The approximation of the neutron-to-neutron energy group transfers by a P3 approximation is generally valid over the energy range of interest for this problem. The approximation would be less valid if the results were dominated by interactions of high energy neutrons with hydrogen nuclei because of the sharp forward peaking of the neutron scattering interaction. The source distribution of the gamma rays generated by neutron interactions including fission was approximated by only a Pn component in the transfer matrix. This approximation is a standard l treatment for these interactions. Anisotropic sources may exist, however, i l it is assumed that the anisotropic component would be eliminated by the randomness of the orientation of the nuclei and the interacting neutrons. The transfer matrix for energy groups corresponding only to gamma rays was also approximated using the P3 approximation. In general, for source problems with only gamma rays this is not a good approximation if deep penetration is involved. However, in this application some of the
- 3. ~t- 2 l
12/$/83 . quantity of interest is due to neutrons and much cf the gamma ray component is due to gama rays that have not penetrated very far through the intervening material. Since the dominant gamma rays will typically travel only a few mean free paths before interacting with the reactor vessel, the P3 approximation for the gama ray transport was felt to be a reasonable approximation. The results of the calculation of the reactor vessel e.d.r. distribution are shown in Table 3.3-1 rnd Figures 3.3-1 to 3.3-3. These results assume that the core support barrel is a uniformly thick annulus of stainless steel, the thermal shield has been removed, and no boron included in the coolant. For these calculations, the effect due to gama rays generated external to the reactor vessel was not included and as a result, the e.d.r. for the region near the outside of the reactor vessel is an underestimate of the total e.d.r. An additional calculation employed the same model except for the addition of the thermal shield. The results for this case are also given in Table 3.3 1 and Figures 3.3-1 to 3.3 3 At the reactor vessel surface, the ratio of the peak energy deposition rate without the thermal shield to that with the thermal shield is 5.2. The removal of the thermal shield increases the ' 9amma ray component of the e.d.r. at the vessel due to gama rays originating from the core and core support barrel, however, the component due to gama rays originating in the thermal shield from neutron [ interactions with the thermal shield material has been eliminated. f In order to account for the effect of the holes in the core support barrel on the e.d.r. distribution for the reactor vessel, an adj ustment factor was l calculated using an RQ DOT model as shown in Figure 3.1-7 of Section 3.1 and an R0 ImT model with a uniform core support barrel. The ad ustment factor calculation uses identical DOT models except for the slot in the l core support barrel. Figure 3.3-4 shows the aQ ustment to the e.d.r. at the core support barrel as a function of the distance from the hole 3.3-3 I
12/5/83 centerline on the reactor vessel inner surface. As in the vessel fluence calculation, the adjustment due to each hole to a given point on the vessel surface is considered as an independent quantity and the total adjusted e.d.r. at a point on the vessel inner surface is obtained by suming the e.d.r. adjustments due to each hole in the core support barrel. e l l 3.3 4 l _
* - ' - 12/5/83 3.4 CALCULATION OF VESSEL WALL TEMPERATURE DOE TO HEAT GENERATION Removal of the Millstone 2 thermal shield rc3ults in an increase in the reactor vessel neutron and gamma induced heat generation rates and a corresponding increase in vessel ter.perature over the corresponding radiation induced heat generation rates during plant operation with the therwel shield. The temperatures found in the thermal analysis were based on the estimated energy deposition rate distribution (Table 3.4-1) which is conservative with respect to the calculated values in Section 3.3.
4 The impact of this increase in heat generation rates on vessel temperature was determined using classical heat transfer methods. A one-dimensional slab model of the vessel wall with exponential decay of the heat generation rate between the inner and outer vessel surfaces. An adiabatic boundary condition was imposed at the outer surface of the vessel, while the forced convection boundary condition associated with normal reactor operation was imposed at the inner surface of the vessel. Results fron these calculations, shown in Figure 3.4-1 indicate that the increase in heat generation rates causes a maximum increase of 22'F in local vessel temperature at the outer surface of the vessel; incredses of 17.5'F and 1.4*F occur, respect 4vely. 25?, of the way through the vessel wall and at the inner vessel surface. The effects of this increase in i temperature on PTS response are discussed in Section 3.5 3.A-1 l
ut nt ~ 3.5 EFFECT OF GAMMA HEATING ON PTS RESPONSE In order to evaluate the effect of neutron and gama ray heat generation on vessel integrity, two pressurized thermal shock transients were used. The first transient was a Steam Lire Break which has a rapid cool down and subsequent gradual heat up. The second transient was an exponential cool down to a final temperature. Both transients were run with constant normal operating pressure of ??50 psi. These transients have the major thermal components of any pressurized thermal shock scenario and can be used to measure the effects of radiation heat generation. Figures 3.5-1 and 3.5-2 show the thermal transients used. 3.5.1 Method of Evaluation Each transient was run with and without the radiation heating effect. The OCA-I program (Ref. 3.5.1) was used for each transient. Cladding was also considered using the OCA-II (Ref. 3.5.2) program. The initial temperature distribution for radiation heating is given in Section 3.4 In addition to the specified radiation heating effect, twice the radiation induced increase in wall temperature was also evaluated. This latter case makes the initial (gradient effects greater. Consideration of cladding was also included in separate runs. A theoretical high copper (.357,) vessel material was asstroed to evaluate the effect of radiation heating on PTS effects. (A high copper theoretical vessel naterial was chosen simply to study the effect of radiation induced heating on PTS results. Millstone 2 vessel materials have lower copper content than ( assumed in these evaluation and are thus more resistant to PTS effects.) l l For each transient, with and without radiation heating, the fluence was increased until a crack would be initiated. Fluence at incipient initiation was the criterion used in this evaluation of the effect of radiation heating. (The fluence values were arbitrarily increased until 3.5-1 l l
- u. ;, s . . . .
j incipient crack extension was calculated simply to study this effec', l .The fluence values calculated are not related to the actual operation of Millstone 2). 4 3.5.2 Results For each thermal transient case, incipient crack initiation was indicated later in life (higher fluence) when radiation heating was included. For more radiation heating, i.e., higher temperatures at the , outer wall, incipient initiation was indicated even later. Figures 3.5-3 and 3.5 A are critical crack depth diagrams for the steam line break transient, with and without radiation heating. Figure 3.5-3 shows the start of initiation without radiation heat. Figure 3.5 4 shows, at the same fluence, no initiation when radiation heat is included. 4 Figures 3.5-5 and 3.5-6 are critical crack depth diagrams for the steam line break transient with cladding considered using OCA-II, with and without radiation heating. Comparing these two figures, the initiation curve for the case with radiation heating starts later in time when compared to the case with no radiation heating. Therefore the treatment of cladding using either OCA-I or OCA-II does not change the conclusion concerning the slight benefit of radiation heating. i Figures 3.5-7 and 1.5-8 are critical crack depth diagrams for the exponentially decaying transient, with and without radiation heating. Comparing these two figures, the initiation curve for the case with radiation heating starts later in time when compared to the case with no j radiation heating. l 3.5.3 Conclusion { While the gradient effects lead to higher thermal stresses, the higher temperatures improve the material properties, KIC and Kig. The l material in the vessel wall is always at a higher temperature and in l 3.5-2 l l l
our ..o - spite of initial higher thermal loading, the improved fracture toughness properties deferred initiation. This demonstrates that neglecting consideration of radiation heating is a conservatism in pressurized thermal shock vessel integrit'y evaluations. 4 e i I i 3.5-3 i
REFERENCES 3.5.1 OCA-1, A Code For Calculating the Gehavior of Flaws on the Inner Surfaces of a Pressure Vessel Subjected to Temperature and Pressure Transients, Iskander, S. K., Cheverton, R. D., Ball, D. G. 3.5.2 OCA-II, A Code for Calculating the Behavior of Flaws on the Inner Surfaces of'a Pressure Vessel Subjected to Temperature and Pressute Transients, Iskander, S. K. , Cheverton, R. D., Ball, n. G. e e i { l l i l 3.5-4
12/b/M.5 - 3.6 NEllTRON FLtlX MONITOPING i In order to establish a solid benchmark for the new fluence levels which will result from the removal of the thermal shield, a flux monitor has been l installed. This monitor was installed in the 97* location where capsule W-97 was removed in 1980 l The replacement in-vessel dosimetry capsule inserted at the 97* position is similar to the capsule previously removed from this location; the principal difference being minor dimensional modifications to facilitate remote installation and the use of blanks in the compartments where the surveillance pressure vessel material samples would be encapsulated. The three sets of neutron flux measuring foils will be in the same position as the previously removed samples. The sulfur threshold flux monitor will be replaced with a cadmium covered Np-237 flux monitor. The short half-life of the p32 usually does not produce any usable information due to the usual long delay from reactor shutdown and the counting of the samples. The substitution of NP-237 will extend the measurement range of the neutron spectra from the present 0.7 MeV down to about 0.1 MeV. Pertinent characteristics of the reconmended neutron flux foils in each of the three positions of the surveillance capsule are summarized in Table 3.6.1. This, capsule will ha pulled with the next reactor surveillance capsule to allow accurate evaluation of both current fluence, flux and vessel material properties. l 1 I i
- 3. 6- 1
12/5/83 L 3.7 TECHNICAL SPECIFICATIONS The pressure-temperature limit curves of Section 3/A.4.0 of the Millstone , 2 Technical Specifications have been reviewed for applicability to future plant operation without a thermal shield. It has been determined that updated curves will be required although startup from the current outage utilizing the present curves is acceptible. The present curves had previously been reviewed to determine what changes, if any, were required as a result of the surveillance capsule (W-97) analysis. The review of any additional impact of removing the thermal shield and the attendant fluence increase at the reactor vessel wall has also been
, cumpleted.
The results'of these reviews, performed in accordance with the Bases for l Section 3/4.4.9, show that the pressure - temperature limits of Figure 3.4-2b are appropriate for 5.3 EFPY. As such, they are applicable for an additional 0.3 EFPY. NNECn has generated revised heatup and cooldown curves for Millstone Unit No. 2 to support continued operation beyond 5.3 EFPY. The revised curves incorporate the information obtained from the surveillance capsule analysis and the revised fluence predictions due to the therma' shield removal. The curves have been constructed utilizing current methodology accepted by the NRC Staff with appropriate conservatisms. Although NNEC0 maintains the applicability of the present curves in Figure 3.4-?b, the revised curves, where more limiting, will be l administrative 1y imposed for future operations. I l l l A license amendment incorporating the new curves into the Millstone Unit l No. 2 Technical Specifications will be requested in sufficient time to i allow appropriate reviews prior to the expiration of the present Figure 3.4-2b applicability. l 3,7-1 l
- e. . .
12/5/R3 Table 3.1.1 Millstone 2 A erage Fast Neutron Flux (E 2 1 MeV) (n/cm .s) (AtPeakAxialandAzimu>thalLocationi BOL to E005 ROL to REFPY k g. BOL to 3?EFPY Avg. Vessel / Clad Interface 3.0+10 3.9+10 5.0+10 1/4 Thru Vessel 1.6+10 2.1+10 2.5+10 1/2 Thru vessel 7.4+9 1.0+10 1.2+10 3/4 Thru Vessel 3.2+4 4.440 5.3+9 Table 3.1.2 2 Fast Neutron Fluence (E 1 MeV) (n/cm ) (AtPeakAxialandEz>imuthalLocation) BOL to EOC5 2 BOL to BETPY' BOL to 32EFPY 3 (Actual) Vessel / Clad Interface 4.R+1R . Q R+1R 5.0+14 1/4 Thru Vessel 2.5+1R 5.4+18 2.S+19 1/2 Thru vessel 1.2+1R 2.5+18 1.2+19 l 3/4 Thru Vessel 5.2+17 1.1+1R 5.3+18 l 1 Numbers shown as 3.n+10 are to be interpreted as 3.0 x 10I0 2 Approximately 5.0 EFPY (1.58 + Bs) P ?700 ht. 3 Full power is defined as ?70n ht. 4 The estimated uncertainty for all fluence values in Tab?es 3.1-1 and 3.1-2 is +307,.
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* ' 12/5/83 TABLE 3.2.2 fH11 stone 2 Designation RT NOT Fluence (10I 'n/cm2 )
Plate C 504-1 137 1.0 Plate C-Sn4-2 135 1.0 Plate C-504-3 129 1.0 Plate C-Sn5-1 159 5.0 Plate C-505-2 197 5.0 Plate C-505-3 170 6.2 Plate C-506-1 1R6 5.0 Plate C-506-2 1A0 5.0 Plate C-Sn6-3 156 5.n Weld 1-203 171 .75 Weld 2-203 94 5.0 Weld 3-203 94 5.0 Weld 8-203 145 .75 Weld 9-203 205 5.0 I I l i
. . . 12/5/83 .
TARLE 3.3.1 Reactor Vessel Peak Energy Deposition Rate Radial Position Long Term Average
- Long Term Average *
(without thermal shield) (with Thermal Shield) (w/cci (w/cc) Ves:tel/ Clad Interface .300 .077 1/4 t .n65 .017 1/2 t 016 004 3/4 t 004 0013
- Power level at 270n Mwt and boron level is at 0 PPM, axial peaking of 1.14 is included. An uncertainty factor of 1.30 is included.
f
~ - - - - - - - - + . - - - - - - . - - . , ,, , _ _
TABLE 3.4.1 Vessel Wall Heat Generation Rates location Heat Generation Rate ("/cm31 Without Thermal Shield With Thermal Shield Vessel Clad Interface n.64n 0.146 1/4 t n.In7 n.n2A
' 1/2 t n.018 n.nna 3/4 t n.ona n.onn7 i
l i l 1
IX/b/H:4 TABLE 3.6 1 . . Material for Neutron Flux Monitors Material Reaction Threshold Energy (MeV) Hal f-Life Neptunium *** Np237(n.f)Cs137 0.5 30.2 years Uranium
- U23R(n.f)CSl37 0.7 30.2 years Iron Fe68(n.p)Mn"# A.n 314 days Nickel ** NiSR(n.p)Co6A 5.0 71 days Copper ** Cu63(n,)Co"0 7.0 5.3 years Titanium T146(n.p)ScA6 R.0 R4 days Cobalt
- CoS9(n,)Co60 Thermal 5.3 years
- Cadmium shielded and bare (uranium encapsulated in vandaium for shielded monitor) ,
** Cadmium shielded
***Encapsulatted in vanadium, cadmium shielded.
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4 i 12/5/83 , FIGURE 3.2-1 MILLSTONE #2 REACTOR PRESSURE VESSEL MAP Initial RT NDT in F 4 o o ' j m S R Outlet , Inlet Inlet m ,S,,
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12/5/83 TABLE OF CONTENTS CHAPTER 4 SECTION SUBJECT 4.0 NONDESTRUCTIVE EXAMINATION TECHNIOUES
4.1 INTRODUCTION
4.2 VISUAL EXAMINATIONS 4.3 EDDY CURRENT EXAMINATIONS 4.3.1 EOUIPMENT 4.3.2 CRITERIA 4.3.3 OtlALIFICATION 4.3.4 SIGNAL ANALYSIS 4.3.5 CALIBRATION 4.3.6 SCANNING SEQUENCE l l 4.4 ULTRASONIC FXAMINATIONS i 4.4.1 E0UIPMENT 4.4.2 CRITERIA AND CALTBRATION 4.4.3 OUALIFICATION 4.4.4 SCANNING SE00ENCE 4.5 POST-REPAIR MACHINING NDE
1.7/5/R3 A.0 NONDESTRUCTIVE EXAMINATION TECHNinUFS A.1 INTpnnt!CTTOP in support of the repair, several nondestructive exanination (NDET testing techniques were utilized to detect, locate, and quantify cracking of the core support barrel. A four step inspection process was used: 11 Visual inspection to locate outer surface cracking on critical areas of the core support barrel.
?\ Eddy current tasting to confirm the visual results and detect potential nonvisible flaws around the lugs.
34 Illtrasonic testing to measure the depths of observed cracks, to detect cracks under the lugs, and tn detect inner surface core support barrel originating cracks, d) Repair process eddy current testing to direct the cutting tcol to the crack tip and to confirm complete crack removal. Each testing method was qualified in laboratory tests using realistic mockups including test samples with artificially induced fatig.se cracks. These test Methods were independently reviewed by outside consultants to provide additional assuranca that the technique is valid. Cracking from the outer surface was detected on only two of the nine lugs. No indications of inner surface originating cracks were observed in any of the areas examined. The two damaqad lugs were nachined off and eddy current and ultrasonic tests were performed over the area to assure that no cracks were obscured by the lug. i 4.1-1 l
I?/5/R3 Visual, eddy current, and ultrasonic examinations were governed by quality procedures that were previously approved by Northeast Nuclear Energy Company. Procedures and personnel certifications were written in accordance with the ASME Code Section XI. e 4.1-2 _. .__ D
12/5/83 4.2 VISUAL EXAMINATIONS Remote visual examinations of the core support barrel were conducted
- utilizing an underwater camera and lighting system that was maneuvered into position by a Remote Activated Floating Television System (r.-E RAFTS). These examinations were video taped in order to provide a record of the examinations. All of these remote visual examinations
' were conducted in accordance with Section XI of the ASME Roiler and ,
Pressure Vassel Code (1474 Edition through Summer 1075 Addendai and as such are capable of resolving a 1/17 inch black line on an IR", neutral grey card Ireference ASME paragraph IWA-?210). The video tape record was made of the core support barrel lug areas and the lower hardface pad areas. The coverage includes special attention to the lug regions and upper hardface pads. Still photographs were used to augment the videotape record, and the entire package of test results were reviewed by certified visual inspection personnel and cognizant engineering personnel. Detailed drawings of all regions with visible indications were made for further evaluation. A.2-1 i t 1 -.-- -.4 - - , . -- -
i I 4.3 EDDY CURRENT EXAMINATIONS Eddy current surface testing was performed to 1) confirm flaws found by visual examination, 2) detect flaws not found by visual examina-tion, 3) locate the coordinate (X,Y) positions of the crack tips, and
- 4) determine the angle of the crack relative to the horizontal or 2
vertical axis. The eddy current method was set to have a detection threshold of one-fourth of an inch long by 0.030" deep cracks for the initial scanning. The repair process eddy current test was demonstrated to confirm crack removal within 0.005". 4.3.1 Equipment s-All of the data acquisition sleds holding the examination instruments were posit'ioned by a mechanical system suspended from the refueling pool wall adj acent to the core support barrel. A central mast provided the vertical positioning for a motorized X table system. The mast was able to traverse from the work platform above the pool to the bottom of the core support barrel. The X table could be adjusted to l any position within fifteen inches of the centerline of the mast. All coordinate locations were measured by means of an encoder system. All coordinate movements and 90* sled rotations were controlled from a remote operator station using programmable motor controllers. The American Society of Non-Destructive Testing Level II and Level III operators directed the system operation via an underwater television / j communication link between the remote trailer location and poolside activities. I Conventional eddy current equipment consisting of a Zetec MIZ-12 display and tester, a Hewlett Packard 8-channel Magnetic Tape l Recorder, and a 2-channel strip chart recorder and an X-Y recorder were used. Remote placement of the data recording equipment in a trailer was made possible by using remote amplifiers and approximately 500' of coaxial cable located outside containment. Eddy current 4.3-1
17/5/83 i coils, 0. A5" in diameter, were enclosed in a stainless steel jacket designed to minimize the effect of wear, hydrostatic pressure and radiation. A r-scan system based upon the X/Y travel of the oositioning system was developed. The system received signals from the encoders and transmitted them to a receiver at the data recording station. Additional electronics perforced a digital to analog conversion of the
' signals for use in driving the X/Y ree. order. The same electronic instrument also acted as an adjustable " gate" for the analog vertical component of the eddy current signal. This feature allowed only those signals having a voltage defined by the inspection criteria to be written to the X/Y chart. The net result was a one-third scale plot of the actual flaws.
.Two addy current transducers were mounted with a sprina load in Sled 41 (Figure A.311 to minimize the lift off effects. The design allowed for transducer positioning against the lug weld for Inn *,
coverage of these areas. A.1.? Criteria l The minimum size flaw required to he detected on the surface by eddy current was a one half inch linear indication. This was based on preliminary analyses which established one half inch through-wall cracks as acceptable. More detailed analyses perforned recently have indicated substantially greater flaw tolerances than were utilized f in the design and calibration of equipment. 4.1.3 Oualification Laboratory qualification testing was performed on fatique cracked 3na I stainless steel and EDM notched blocks tn determine the li datectability of actual cracks, 71 indexing required for 100" coverage, 31 calibration flaws and reference voltages and di geometric interference with probe sensitivity. The flaws in the test blocks were milled out in successive staps interspersed with eddy current and 4.3-2
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12/5/83
- dye penetrant testing until completely removed. This information was
.used as the basis for the calibration voltage and to qualify the eddy current test to determine complete crack removal following repair machining. An independent review of the examination process, qualification tests, and equipment was provided.
4.3.4 Signal Analysts Throughout the qualification, the signals obtained from cracks,
- s. dents,and scratches in the test blocks'were observed by the operators who were to perform the actual' examinations and analyze the data in the field. This' familiarity witb the various signal types allowed for the discrimination of false inoications.
4.3.5 Calibration i The calibration voltage (A.0 volts) was set.'on the MIZ-12 by passing
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the sled mounted ET probes over a 1" EDM noten (.040" Deep x .005" wide) in a 304 stainless steel block (Figure 4.1-2). With this set-up, a one fourth inch EDM notch gave a reference voltage of 1.6 volts. From this result, the ET gate recording threshold on the X-Y recorder was set at 50", of ore fourth match response (n.R volts) giving a 4 fold conservatism in testing based on the criteria established for-this test. Rased on the recording threshold above, the indexing increment was f established by passing the probe both perpendicular and parallel to ( the one-fourth inch EOM notch in an incremental manner to determine the maximum increment which would not yield less than 50", of the reference voltage. For the probes used, the increment was 0.3". Since the examinations were being performed on components having l radiation fields in excess of 2 x 106 R it was anticipated that the test sled and probes might become contaminated with high levels of radiation. As a result, a reference check probe was included in the 4.3-3 -
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system. Using this probe, frequent checks were made on the electronic portion of the test system. , 4.3.6' Scanning Sequence (Scanning was n.onitored by an ASNT Level II or Level III operator as I 2
- l.he tests were perforced. The datum point was established by
, centering a sled wheel on the left side centerline of the upper hardface pad. The X/Y coordinates of this sled position were recorded. This information coupled with the as-built dimensions of the sled allowed for the accurate reconstruction of the flaw relative 4
to the lug. Encoder calibrations were checked for by returning to this point during the examinations. Scans were performed out to 4" on either side of the lug, over the area above the lug, and below the hardface. One scan was made below the lug. Cracks beyond the 4" scan area were followed until the crack tip was located. All data was recorded on the X/Y plotter, magnetic tape and strip chart for analysis. Data analysis was performed off-line from the testing. The process included review of the X/Y plot and strip charts to verify that the checklist required scans were performed, review of the magnetic tape j to analyze the signals in order to code signals as either cracks or L false indications, determining which X/Y plot signals were cracks and making a composite X/Y plot from individual plots. l l 4.3-4
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j 12/5/R3 t 4.4 ,Ultratonic Examinations Ultrasonic volunetric testing was performed to determine 1) the depth of flaws found by visual and eddy current examination, 2) flaws under the lug and hardface pad, and 3) inside rurface initiated flaws. The ultrasonic method was set to have a detection threshold of 0.035" deep f, laws (27, of T) for the inner and outer surfaces. The ultrasonic method was demonstrated to have a depth measurement accuracy of ;t 15%. 4.4.1 Equipment The ultrasonic test system consisted of a Krautkramer-Branson Model 6000 ultrasonic instrument; a J. C. Technical remote pulser and line dr1ver, Model 5000; and a C9mbustion Engineering Programmable Signal Generator, Model PSG-16 The latter is an electronic calibration simulator. The remote pulser allowed pl? cement of the equipment outside containment. The ultrasonic transducers (Panametrics Model A3275) were 2.25 MHz, 0.5" diameter, in-ersion type and were radiation hardened. J The ultrasonic display and the X and Y positional encoder displays were videotaped for analysis and for a pernanent record. Three multi-transducer sleds were used to perform ultrasonic tests; Sled #1 with 0* (Figure 4.3-1), Sled #2 with AS* (Figure 4.4-1) (relative to surface perpendicular) and Sled #3 (Figure 4.4-2) with dual element sound beams. Since some of the cracks were at angles other than 0* and 90*, Sled #7 examinations were supplemented by scans < l utilizing 6 rotatable transducer arrangement incorporated into Sled
#7. This allowed more accurate determinations to be made of the crack
- through wall coordinates. Complex calculations to correct the data i relative to the datum point were performed by an oa site computer programmed with the relevant coordinate corrections.
4.A-1
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12/5/R3
, t 4.4.2 Criteria and Calibration The ultrasonic examinations were used to measure the depths of -
cracks. Angle beam (45') transducers were calibrated using the 2", notch and the one-eighth inch diameter side drilled hnles of the calibration block (Figure 4.4-3) for the sensitivity and Distance-Amplitude Correction (DAC) curve, respectively. The na transducers were calibrated on the side drilled holes for detection of a laminar tearing and far backwall monitoring for thickness checks. The dual focus transducers were calibrated using the side drilled holes and the circular slots of the calibration biock for the DAC curve and
' sensitivity,respectively. During scanning, the sensitivity was a minimum of 6dB greater than the reference sensitivity.
4.4.1 . 0uali fication Testing was performed on cracked 304 stainless steel blocks and the calibration block to determine li the detectability and depth measurement of actual cracks and 2) calibration sensitivities and procedures. Actual depth of the cracks were determined by milling out the test blocks and verifying the depth when completely removed with eddy current testing and dye penetrant testing (after a macro-etch to l open remaining flaws). Comparison of actual depths to the depths determined by the dual focus transducers showed a 15% underestimate for this UT method. Depth determination using the angle beam j transducers was based upon conventional UT methods using the side drilled holes, notches, and screen position. An independent review of the examination process, qualification tests, and equipment was provided. 4.4-2 l l
e** l 1?/5/83 4.A.4 Scanning Sequence Scanning was monitored by an ASNT Level I? or Level III operator as the tests were performed. The datum potnt was established in the same nenner as in the eddy current sled (i.e., left centerline of hardface pad). Cata analysis was performed off-line from the testing. The process
- included 1) plotting the areas scanned to assure full coverage 2) reviewing all of the videotapes to record the screen positions and X/Y coordinates of the flaws 3) correcting the X/Y data to compensate for the transducer effects and rotations of the sleds and 4) plotting the data on a' lug drawing. Wherever possible, the data was correlated with the geometry of lug cracks as determined from visual and eddy current examinations. .
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12/5/83 4.5 POST.REPATR MACHINING WDE Eddy current testing was used in cordunction with the milling machine operation to verify the position of the crack tips relative to the datum point and complete removal as required. The system was calibrated in the same manner described above for the crack detection test and the probe was mounted into an air spring on the cutting tool. The readout coordinates for eddy current was corrected by the actual mechanical offsets to the cutting tool so that the latter is operated in the correct location. The crack tip location was reconfirmed by eddy current prior to
, machining. After the initial milling, the eddy current test was repeated to confirm that the subsurface crack did not extend beyond the surfate crack tip. Additional machining was completed after confirmation. For less than through wall cracks, the eddy current test was repeated periodically during the milling operation to confirm complete elimination of the crack over the entire length of the milled slot.
4.5.1 Oualification
- For the qualification of this test, the fatigue cracks that had been made for the ultrasonic testing were machined out in a manner similar to the repair machining, Crack removal was confirmed by acid etch and penetrant testing periodically during machining along with the eddy current testing. The results of this testing demonstrated the efficiency to confirm crack removal within 0.005 inch. (Note:
Dualification Testing Program for the Ultrasonic Testing is discussed in 4.4.3 and for the Eddy Current Testing in 4.3.31.
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12/5/83 TABLE OF CONTENTS CHAPTER 5 SECTION SUBJECT - 5.0 NONOESTRUCTIVE EXAMINATION INSPECTION RESULTS 5.1 INTRODUC110N 5.1.1 THE INSPECTION PROGRAM 5.2 DE3CRIPTION OF REACTOR VESSEL INTERNALS INTERFACE INSPECTION RESULTS
5.3 DESCRIPTION
OF THERMAL SHIELD INSPECTION RESULTS 5.3.I THERMAL SHIELD ann UPPER POSITIONING PINS INSPECTION 5.3.1.1 THERMAL SHIELD POSITION 1/R (110') 5.3.1.2 THERMAL SHIELD POSITION 2/S (150') 5.3.1.3 THERMAL SHIELD POSITION 3/T (190') 5.3.1.4 THERMAL SHIELD POSITION 4/U (210') 5.3.1.5 THERMAL SHIELD POSITION 5/V (270*) 5.3.1.6 THERMAL SHIELD POSITION 6/W (310*) 5.3.1.7 THERMAL SHIELD POSITION 7/X (350') 5.3.1.8 THERMAL SHIELD POSITION 8/Y (30') 5.3.1.9 THERMAL SHIELD POSITION 9/Z (70') 5.3.2 LOWER POSITIONING PIN INSPECTION
5.4 DESCRIPTION
OF CORE SUPPORT BARREL INSPECTION RESULTS 5.A.1 CORE SUPPORT BARREL LUG INSPECTION RESULTS 5.4.1.1 LUG 4/U (30') (FIGURE 5.4-1) 5.A.1.2 LUG 5/V (270') (FIGURE 5.4-2) 5.5 FUEL INSPECTION RESULTS l
i 12/5/R3 5.n NONnESTRilCTIVE EXAMINATION INSPECTION RESULTS
5.1 INTRODUCTION
The inspection program with the core support barrel removed from the reactor vessel was performed as part of the routine in-year inservice inspection progran and as a result of dansge identified to this component at a unit of similar design. The inspection program consisted of a visual examination of the thermal shield and the thermal shield supports using a remotely controlled underwater TV camera. This inspection was performed immediately after the core support barrel, with the thermal shield attached, was removed from the reactor vessel. The areas around the support lugs and support pins, Figure 2.1 4 of Chapter 2, and the upper and lower positioning pins were scanned from the outside of the thermal shield and from within the core support barrel to thernal shield annulus to characterize damage to the interfaces. . During the initial inspection program the reactor internals interface with the reactor vessel was exanined to determine if the relative motion of the thermal shield had caused excessive motion of the core support barrel and resulted in wear or damage. The reactor vessel outlet nozzles, surveillance capsules, reactor vessel snubbers, alignment keys, and keyways were examined. The reactor vessel lower head, the flow skirt, and the plate surface of the core support barrel assembly below the core support plate were examined for debris. 5.1-1 i
n, .- 1?/5/83 4 += The objective of the initial inspection was to characterize the damage to the thermal shield and surrounding observable components and to establish the initial course of action. 5.1.1 The Inspection Program A program was designed to document the damage to tha thermal shield and the core support barrel. The purposes were to assist in the confirmation of the failure, provide data to determine structural integrity and to formulate a repair procedure. The thermal shield was examined using underwater photography. Remote cameras were located on the outside diameter, in the themal shield support ba'rrel annulus. Upon removal from the core support barrel, the thermal shield segments were examined on the inside diameter. The core support barrel was inspected using three NDE techniques. A visual (television) examination of the core support barrel provided an averall view. Eddy current testing was used to verify crack length and search for cracks not seen during visual examination. Ultrasonic
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testing was used to quantify crack depth. (See Section 4.0 for a
-detailed discussion of these techniques.)
l l 5.1-? l i
12/5/R3
5.7 DESCRIPTION
OF REACTOR VESSEL INTERNALS INTERFACE INSPECTION REStlLTS An inspection of the reactor vessel and internals interfaces was completed using remote TV cameras, and the components were found to be free of damage. The areas examined were the interfaces on the flow skirt, reactor vessel snubbers, keyways, outlet nozzles, upper guide structure to core support barrel, and the upper guide structure to the core shroud. e 5.2-1 L
N- 17/5/81 l 5.3 . DESCRIPTION OF THERMAL SHIELD 1NSPECTION REstfLTS This Section documents the damage to the thermal shield and its support system. Information is prwided on the upper support areas, the lower positioning pins, and the thermal shield debris recovered. 4 5.3.1 Thermal Shield and Upper Positioning Pins Inspection As shown in Figure 5.3-1, the thermal shield in the area of the 0 support lugs and support pins was heavily damaged near the 0 - 1R0 0 axis, less damaged near the 90 0
- 2700 axis. Damage to the-thermal shield occurred at positions 25, 3T, 6W and 7X, causing both the support pin and large pieces of the thermal shield to break loose. Two upper positioning pins and lock bars were missing. A detailed discuss.on of each thermal shield position is found below.
0 i 5.1.1.1 Thermal Shield Position 1/R (110 ) l l Thermal Shield : No damage occurred. Support Pin : There is a circumferential crack where the pin is
- - welded to the thermal shield. Extreme wear occurred at the pin-to lug interface. The top of the lug has rolled over.
Upper Positioning Pin : The pin and Icek bar were in place with no damage, though the pin had shifted downward on the core support barrel hardface due to relative movement between the core support barrel and the thermal shield. 0 5.3.1.2 Thermal Shield Position 2/S (160 ) \ Thermal Shield : A triangular section on the left of the support f pin was missing from the top of the thermal shield. l l l
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12/5/A3 Support Pin : A significant portion had broken loose and was missing. Upper Positioning Pin : The pin and lock bar were in place with no damage. 0 5.3.1.3 Thermal Shield Position 3/T (190 1 Thermal Shield : Two major cracks on both sides of the support pin
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with both pieces of shield appearing ready to break away. SuDport Pin : Sections of pin cracked and e*Deared read) to break s away. Upper Positioning Pin : The pin and lock har were in place with no damage. 0 5.1.1.4 Thermal Shield Position A/U (230 ) Thermal Shield : No damage observed. Support Pin : Evidence of'.some wear on top front edge. No damage cbserved l tipper positioning Pin : The pin and lock har were in a place with no damage. 4 0 5.1.1.5 Thermal Shield Position 5/V (2701 Thermal Shield : No damaged observed. Suoport Pin : A small indication at top left edge. . ! Upper Positioning Pin : Both the pin and lock har are missing from the hole. 5.3-?
12/5/A3 0 5.3.1.6 Thermal Shield Position 6/W (310 ) Thermal Shield : A triangular section on the left of the support pin is missing. A major crack on the right with the section almost broken loose. Support Pin : A small section on the left is cracked and pulled out slightly. Upper Positioning Pin: The pin and the lock har were in a place with no damage 0 5.3.1.7 Thermal Shield Position 7/X (350 ) Thermal Stiield : A triangular section on the right of support pin is missing. Support Pin : A major portion of the pin is missing. Upper Positioning Pin : The pin and the lock bar were in a place with no damage. 5.3.1.R Thermal Shield Position 8/Y (3n0) Thermal Shield : Weld between thermal shield and support pin is cracked. No other damage observed. Support Pin : Significant wear occurred at the pin-to-lug intersection. Crack from the upper right corner of the slot to the edge of pin.
' ace with Upper Positioning Pin : The pin and lock bar were no damage.
0 5.~4.1.0 Thermal Shield Position 4/Z (70 ) Thermal Shield : No damage observed. 5.3-3
12/5/R3 Support Pin : Significant wear at the lug and pin intersection. Evidence of upset metal and small cracks at the top edge. Upper Positioning Pin : Both the pin and the lock bar were missing from the hole. 5.3 ? Lower Positioning Pin inspection Position N: Lock bar missing hut weld was intact. Position L: Lock bar missing and the positioning pin was backed out.
, Position X: Weld at one end of lock bar was broken, d
e l 5.3 4
6, . . 12/5/83
5.4 DESCRIPTION
OF-CORE SUPPORT RARREL TNSPECTION RESULTS This section discusses the damage to the core support barrel in a manner similar to the thermal shield damage description. 5.4.1 Core Support Barrel Lug Inspection Results Non-destructive examinations using visual, eddy current and ultrasonic techniques revealed damage to the core barrel at lug locations 4 and
- 5. The core barrel was found to be free of any flaws at other lug locations.
5.4.1.1 Lug A/U (300 ) (Figure 5.4-1) One through wall crack was found which runs across the bottom of the lug from the left to the right. A non-through wall crack was also found which runs vertically down along the right side of the lug and curves to the right and merges with the through wall crack. The configuration of the through wall crack is shown in Figure 5.4-1. 5.4.1.2 Lug 5/V (2700 ) (Figure 5.4-2) The through wall crack was found. This crack runs down the left side of the lug, across its face and out into the core barrel on the right l of the lug. The crack configuration is shown in Figure 5.4-?. l l 1 l l 5.4-1
I?/5/n 5 i. l 5.5 FtlEL INSPECTION RESULTS A comprehensive fuel inspection program was conducted following Cycle 5 operation of Millstone 2. The motivating factor for the inspection program was evidence of fuel pin leakage indicated by primary system activity. The inspection program included fuel assembly sipping, visual examinations, ultrasonic examinations of fuel pins, a review of plant operating history and a reciew of the fuel manufacturer's records. NNECO met with the NRC Staff on October 12, 1983, to discuss the fuel pin failufes and fuel assembly component failures identified during i the inspections outlined above. A sumary of the information presented to the Staff at the meeting was provided in a letter to J. R. Miller from W. G. Counsil, dated November 4,14R3, Docket No. 50-336. The cause of failure of the fuel pins has not been attributed to any
. single factor. Although debris was observed in the fuel, the majority was non-metallic. NNECn has concluded that of the sources identified as potential contributors to the fuel pin failures, none is indicative of a situation that would lead to continued degradation of the fuel clad.
The fuel assembly ccmponent failures, namely the broken holddown springs and damaged holddown flowers, are not 7,uspected to be related l to the events associated with the thermal shield support systen degradation. These failures are considered to be related to the design of the fuel. The design of the Cycle 6 reload fuel and future reload fuel has been modified to reduce the probability for similar type failures to occur in upcoming fuel cycles. There has been no evidence identified which would indtcate any abnonnality existed in the core as a result of the thermal shieldg d.5-1
1 .. 12/5/R3 support degradation. There have been no instances of Control Element Assemblies (CEA's) either failing the CEA drop time requirement or becoming immovable or untrippable as a result of friction or interferences. In sumary, there has been no evidence identified to date suggesting degraded fuel performance resulted from the thermal shield support degradation. i e l l l 5.5-7 1
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12/5/A3 TABLE OF CONTENTS CHAPTER 6 SECTION SilRJECT 6.0 FAILilRE MECHANTSM ANALYSTS PR00, RAM 6.1 INTR 00tlCTION 6.2 ST. LtICIE 1 FAILilRE MECHANISM ANALYSIS PROGRAM 6.3 POSTillATEn SE0llENCE OF EVENTS 6.A CnNCLIJSIONS s 4 L
12/5/R3 6.0 FAlltlRE MECHANISM ANALYSIS pRnGRAM
6.1 INTRODUCTION
Following the discovery of the thermal shield and thermal shield support system degradation at St. Lucie 1, an analysis program was initiated to determine the mechanism that caused the damage. Northeast Nuclear Energy Company participated in this program for Millstone 2. The similarities in plant design, operating conditions and observed damage patterns provide a basis from which it can be concluded that the results of the failure mechanism analysis for St. Lucie 1 are expected to be applicable to Millstone 2. T'he similarities in design and operating characteristics between Millstone 2 and St. Lucie 1 include the following pertinent factors: o identical reactor vessel internals design o identical reactor coolant system layout o identical reactor conlant pumps design o equivalent operating power levels o similar reactor coolant flow rates at operation A description of the St. Lucie 1 failure mechanism analysis program and the postulated failure mechanism.are presented in the following sections. 6.1-1
12/5/R3 6.2 FAILURE MECH 4NISM ANALYSIS PROGRAM FOR ST. LUCIE The program was divided into three phases, shown schematically in Figure 6.1-1. The three phases are: (1)thecompilation, analysis, and evaluation of data related to the failure,-(2) integration and evaluation of this information to select possible mechanisms for the initiation and progression of the failure, and (31 structural analysis to quantitatively test hypothesized mechanisms against supporting data, and selection of a sequence of wants which best explains the observed condition of the components. The first phase of the program examined a number of areas potentially related to the thermal shield system degradation. These included:
- Hydraulic Loads;
- Structural Response Characteristics;
- Design, Fabrication, and Installation Data;
- Vibration Monitoring Programs and Visual Inspection of Damage;
- Loose Parts Monitoring and Internals Vibration Monitoring Data;
- Metallurgical Evaluations.
Periodic and random hydraulic loads were evaluated for normal operating conditions. Dynamic response calculations show that these loads are not sufficient to cause damage to the shield or the supports in the as-designed condition. Extensive free vibration structural response calculations were performed on a detailed finite element model of the core support barrel-thermal shield as a coupled system. These calculations were done for the system in its normal, as-designed, condition and in assumed degraded conditions. 6.2-1
12/5/A3 Separate analyses to evaluate the effect of a number of parameters on positioning pin preload, i.e., installation procedures, pin torquing sequence, sMeld weight, static and dynamic pressure loads, stress relaxation, and thermal expansion were performed. This evaluation showed that a combination of parameters and operating conditions are important to the retention of positioning pin preload. Preoperational data were r~4wec' for potential effect on the integrity of the support system. No indications of abnormal wear, contact, or incipient damage sites were detected. Limited inservice visual inspections were performed prior to 1983 for the purposes of 4 locating loose parts but none were found. Loose part's Monitoring (LEM) and Internal Vibration Monitoring (IVM)
-data were reviewed and then reanalyzed using state-of-the-art techniques to determine if there were quantifiable changes during operation. LPM data showed changes in magnitude and frequency of the impulsive signals and changes in location of these signals with operating time. The TVM data also exhibited changes in characteristics with operating time. These indications suggest an uncoupling of the thermal shield-core support barrel. Radiological examinations of two positioning pins which became dislodged from thtir original locations indicate that one may have been dislodged in a time
! frame consistent with the LPM data. Results of metallurgical examination of portions of the damaged thermal shield were combined with results of the analytical I investigations to provide a postulated failure mechanism. 6.2 ?
12/5/R3 The results permit construction of a postulated sequence of even+.s which may have led to the thermal shield degradation. The last phase of the program considered forced dynamic response of the thermal shield-core suppc?t barrel coupled system to yarious loading conditions with simulated changes in support integrity. The analyses indicate that loads on the functional supports increase as some support components are removed from the miel. The increased loads likely contribute to wear of the remaining support components. a e f 6.2-3
12/5/A3 6.3 POSTULATED SEnt1ENCE OF EVENTS A postulated sequence of events which may have led to the St. Lucie 1 thermal shield damage has been selected using judgment, examinations, and analyses. The conc 1prions are necessarily inferential due to the absence of direct observation of the behavior cf the thermal shield and related components. It appears certain that the damage to the thermal shield was caused by hydraulically induced loads acting on the thermal shield. It is very likely that this was made possible by deterioration of the thermal shield support system. The deterioration was probably preceeded by loss of preload on positioning pins. The reasons for the loss of preload h#ve not been specifically identified, but several factors have been exanined and found to be capable of contributing. It is believed that a combination of the deterimental factors is the most [ reasonable explanation. With a loss of preload, the thermal shield has been calculated to uncouple from the core support barrel and this is supported by IVM data. The resulting response to hydraulic loads would increase the relative motion between the shield and the core support barrel, causing wear between the lugs and the shield. l With sufficient wear on the lugs, the thermal shield may have reached l an unstable dynamic situation. Analyses show a tendency for the ! ' thermal shield to become unstable with sufficient reduction in the effective stiffness of the lugs. The resulting large motions could cause thermal shield damage in a relatively short period of time. Changes in the magnitude and frequency of LPM alarms and changes in the IVM system data support this scenario. I-6.3-1 l
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12/5/83 6.4 CDNCLUSIONS The results of the failure mechanism analyses performed for St. Lucie 1 indicate that the thermal shield damage was caused by hydraulically induced loads acting on the thermal shield made possible by the deterioration of the thermal shield support system. The similarities between Millstone 2 and St. Lucie 1 support this scenario as the cause of the Millstone 2 - thermal shield degradation. NNECn believes that the core support barrel damage is attributable to the thermal shield support system degradation and the resultant loads in the thermal shield / core support barrel system. This is supported by the fact that core support barrel damage was confined to the thermal shield attachment points. NNECO is confident that removal of the thermal shield from the reactor internals will relieve a maj or source of loading on the core support barrel. 6.4-1
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- 12/5/83 FIGURE 6.1-1 :
PROGRAM TERMAL SHIELD FAILURE IIECHANiSM ANALYSIS TEST O HYDRAULIC LOADS PERIODIC, RAND 0M O, , ECHANISMS 9 QUANTITATIVELY INTEGRATE STRUE"llRA'" '3C8 ^E'I"RFSPONSE (PERFORM l O O RF SUPPORT SNCMDITiogg EVALIJATE DESIGN, FABRICATION SELECT O a inSTAttATi0n DATA O> CANDIDATE FAILURE 5 CECK COMPLIANCE
- REPORT PVMP & DAMAGE VISUAL O insPECTi0n DATA O INITIATION wiTu a SUPPORTING DATA
'"" * " O PROGRESSION O DATA ECHANI2S SELECT CREDIBLE s
ETAllU O CAL EXAM. Os ECHANISM j e e
12/5/33 TABLE OF CONTENTS CHAPTER 7 SECTION SUBJECT 7.0 REACTOR INTERNALS STRESS ANALYSIS AND CORE SUPPORT BARREL STRUCTURAL INTEGRITY 7,1 REACTOR INTERNALS STRESS ANALYSIS RESULTS 7.2 CORE SUPPORT BARREL STRESS ANALYSIS 7.2.1 STRESS FIELD IN CORE BARREL AT LUG ELEVATION 7.2.2 DYNAMIC AND TRANSIENT LOADING CONDITIONS 7.2.3 THERMAL AND HYDRAULIC CONSIDERATIONS 7.3 CORE SUPPORT BARREL REPAIR 7.3.1 FLAW CHARACTERIZATION 7.3.2 REPAIR MACHINING OBJECTIVES 7.3.3 REPAIR MACHINING PROCESS 7.3.3.1 NON-THROUGH WALL FLAWS 7.3.3.2 THROUGH WALL FLAWS 7.3.4 REPAIR CRITERIA 7.3.4.1 NON-THROUGH WALL FLAWS 7.1.4.2 THROUGH WALL FLAWS 7.3.5 FRACTURE MECHAh.CS EVALUATION w -, y -~ , -, _ - . , w -~-r--.--.,---,-y-----,e- e,- - , < - - = , , . - - ,
12/5/83 7.0 REACTOR INTERNALS STRESS ANALYSIS AND CORE SUPPORT BARREL STRUCTURAL INTEGRITY 7.1 REACTOR INTERNALS STRESS ANALYSIS RESULTS The effects of thermal shield removal on the reactor internals were determined to be: (a) A change in the hydraulic loads on the internals. (b) A negligible change in the frequency of the corc support barrel assembly. Evaluation of the above effects on site specific seismic analysis and asymmetric 16 ads analysis and comparison with the original stress analysis showed that the reactor internals with the thermal shield removed meet the ASME Code allowable stresses using the original design criteria for normal operation, upset (Level A and B), and pipe rupture (Level D) conditions. 1 I l 7.1-1
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l 7.2 CORE SUPPORT BARREL STRESS ANALYSIS L A comprehensive stress analysis of the core barrel in its undamaged condition with the thermal shield removed has been performed. Effects of the cracks in the core barrel are found to be local and are accounted for in Section 7 3.4 and 7 3.5. In this section, stresses for the core barrel at the lug locations for various loading conditions are provided. 7.2.1 Stress Field in Core Barrel at Lug Elevation dx c'y Txv Static + 407 11 Pressure - 1,284 Seismic 1 852 273 Thermal OD1 +34,000 +34,000 ID - 8,000 - 8,000 OD2 +17,000 +17,000 ID -17,000 -17,000 Plant fluctuations 1 2,345 1 2,345 Pump Induced Pulsation 1,763 1,233 109 l LOCA 128,000 1 5,092 Cfx = Longitudinal Stress (PSI) d'y = Hoop Stress (PSI) SFxy = Shear Stt ess (PEI) Notes:
- 1. This stress is based upon actual thermal gradient and is a peak value at the surface.
- 2. This stress is based upon a lineariacd thermal gradient 7.2-1 l
i 1, .. 12/5/R3 7.2.2 Dynamic and. Transient Loading Conditions Following is a list of cyclic events which have to be included in the fatigue evaluation. Type Cycles Heat up/cooldown 500 5* Load / Unload 15,000 Step Load Increase / Decrease
+ Normal Plant Variations 10 6 Reactor Trip 4n0 los of Flow / load R0 Operating Basis Earthquake 200 Dynamic Loading 2x10 ll 7.2.3 Thermal and Hydraulic Consideratfons 7.2.'s.1 Thermal l
Core support barrel temperat.ure is an important consideration in the integrity analysis because these temperatures give rise to thermal stresses that must be ennsidered in the core support barrel fatigue analysis. A one-dimensional slab heat transfer model was used to determine temperatures for evaluation of core support barrel integrity. A natural convection boundary condition was imposed on the inner surface (i.e., side facing the core shroud) of the core support barrel, based on a conservatively high coolant temperature. A forced convection boundary condition was imposed on the outer surface (i.e., surface l facing the reactor vessel) based on the vessel inlet coolant temperature. 7.7-2
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\ 12/5/R3 Heat generation rates associated with full power operation were used in the 1-D' analysis. These heat rates were based on the sum of average axici peak power and maximum azimuthal peak power. 7.2.3.2 Hydraulic Loads The normal operating loads, generated for use in the stress analysis, consist of the following categories of loads:
- 1. Static Hydraulic loads, ,
- 2. Pump Induced loads, and
, 3. Turbulence Induced loads.
These are discussed in detail in the following sections. Static Hydraulic Loads The static hydraulic loads acting on the portion of the core support barrel extending from the thermal shield lug elevation down to its bottom are given in Table 7.2-2 and Figure 7.21. Loads were , calculated for the two sets of conditions given in Table 7.?-3; the mcst adverse loads were chosen as input to the integrity analysis. The axial hydraulic load in Table 7.2-2 is based on the operating conditions in Table 7.2-3 and calculated loss coefficients for the flow path segments between the inlet nozzles and the upper region. The maximum radial delta p across the core support barrel wall in Table 7.2-2 is also based on calculated loss coefficients. The lateral loading distributions on the core support barrel given in Figure 7.2-1 are based on measured total pressures and flow kinetic heads in the downconer region of a scaled flow model of the St. Lucie 1 reactor and are applicable to Millstone 2. 7.2-3
12/5/R3 Pump-Induced Loads Pump-induced accoustic loads acting on the core support barrel were calculated at an inlet temperature of 548*F at the following four pump characteristic frequencies:
- 1. Rotor speed,15HZ
- 2. 2 x rotor speed, 30HZ
- 3. Blade passing frequency, 75HZ
- 4. 2 x blade passing frequency,15nHZ The pump-induced loads on the core support barrel are determined using two hydrodynamic models:
- 1. The first model evaluates the propagation of pump-induced pressure pulsations in the cold leg water column from the pump discharge to the inlet nozzle on the vessel.
- 2. The second model evaluates the propagation of pump-induced pressure pulsations in the downcomer water column in the reactor v essel . The output from the first model is used to drive the second.
The wave equation for a compressible, inviscid fluid is set up and solved for each model. For the case of the downcomer, the series solution for the resulting 3D wave equation was solved by means of a C-E computer code, DPVIB. The output from the downcomer model consists of a description of the pressure distribution on the core support barrel wall, P g (RCSR' 0,2). Typically, a pressure distribution is generated at each pump driving frequency for the case of a single operating pump with a noninal unit fluctuating inlet nozzle pressure. The resulting pressure distribution is described by the series: 7.2 4
12/5/83 Pg (Rcso,9.Z)=jH,COSm0 where: H, ={Cnms [J, (Anmsd + nnnt Y,(Anmsr)] COSdn I m = Circunferential wave number n = Axial wave number s = Radial wave number C nms = Fourie. coefficient ) n 2 2 nms = Eigenvalue = (W 2n ms/Co - n )1/2 W nms = Liquid natural frequency Co = Speed of sound in liquid n = Variable related to the axial waves = nn/L r = Radius J M,Ym = B,essel functions of first and second type Z = Axial position 0 = Azimuthal position referenced to the zero degree position for the operating pump L = Length of downcomer annulus The pressure distribution Po (RCSB, Q, Z) based on the nominal unit psi inlet pressure, is-scaled by the calculated inlet nozzle pressure that is output from the model for the cold leg. P(RCSB, Q, Z) = Pinlet x Po (RCSB' 0' 2) where: P inlet = calculated pump induced pressure fluctuation at the vessel inlet nozzle; values are given in Table 7.2-4 To obtain the overall pressure distribution P (RCSB,Q,Z)onthe core support barrel, for rultiple pump operation, the pressure distribution for the single pump case was superimposed at the appropriate azimuthal positions for the particular operating pumps. To maximize pressure fluctuations on the downcomer, the phasing between 7.2-5
i' E ' 12/5/R3 the operating pumps was selected to produce the most adverse loading . condition on the core support barrel. Turbulence-Induced Loads Hydraulic exuitation of the core support barrel due to random turbulence was calculated from a power spectral density vs. frequency plot based on turbulent pressure fluctuation measurements in a scaled PWR model and coherence areas determined from laboratory and field test data inside a PWR. The normalized power spectral density (PSD) plot developed from those a sources for the downtomer annulus is given in Figure 7.2-?. The parameters, associated with the PSD are defined as:
$ - normalized power spectral density for the turbulent pressure fluctuations PSD - power spectral density for the turbulence, PFS 2/HZ I - coolant density, lb/ft 3 -
U - average coolant velocity FPS U cony - turbulence convect, ion velocity, FPS 6 - radial gap of' annulus, ft F - frequency HZ 2 Acoherence - coherence area, ft The coherence area for the turbulent pressuras was calculated from the square of the radial gap of the downcomer annulus. 7.2-6
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- 12/5/R3 7.3 CORE SUPPORT RARREL REPAIR 7.3.1. Flaw Characterization w
As described in Section 5.a of this report, the flaws in the core support barrel were through wall as well as non-through wall. All flaws appeared to propogate from the lug region on the core support I barrel into the barrel it'self. For complete details of flaw orientation, see Figures 5.4-1 and 5.4-2. 7.3.2 Repair Machining Obj ectives The 'obj ective in repairing a11 flaws was to assure the suitability of i the core support barrel for continued use considering normal operating and accident loads. Flaws which were not through wall were blended out in a milling operation. Those flaws which did go through wall were stabilized by crack arrestor holes located in the vicinity of the crack tips in the milled slot. 4 7.3.3 Repair Machining Process 7.3.3.1 Non-Through Wall Flaws The process for removing non-through wall flaws involved milling slots l in the core barrel along the length of the flaw. The process removed
.125 inches of material in each cutting pass and was followed by an eddy current inspection of the slot surface. The process was continued I
until the flaw was completely removed from the slot.
- l. 7.3.3.2 Through Wall Flaws Flaws which were through wall for some portion of their length were arrested by drilling a 1-1/R" diameter hole. The positioning of this hnle was accomplished using cn inspection process coupled with milling the core barrel at the surface within the crack tip region. The prncess removed .175 inches of material in each pass followed by an eddy.
1 4 7.3-1
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12/6/R3 current inspection of the slet. This process was continued until assurance of crack tip recession to the lug was achieved. With this assurance, a 1-1/8" hole was drilled at the end of the inspection slot. Figure 7.3-1 depicts the configuration of the slot and hole and shows where a typical crack tip may appear in the slot. 7.1.4 Repair Criteria 7.3. A.1 Non-Through Wall Flaws The removal of a non-through wall flaw from the core support harrel resulted in a milled slot of a depth sufficient to remove the flaw. The criteria which were applied to the slot reflected concerns of both stress and fatigue. The evaluation of fatigue assured there would be no crack initiation from the slot for normal operating and transient conditions. The remaining ligament requirenents are based upon menbrane stress calcuiations for differential pressure across the core barrel which would result fom asynmetric loadings. 7.1. A.7 Through Wall Flaws ! The repair of through wall flaws involved milling an inspection slot I and drilling a through wall crack arrestor hole. The design considerations for this repair involve fatigue and crack stability. The fatigue concern focused around the possibility of propagation of a new flaw from the arrestor hole. This hole was sized to keap the ! intensified stresses below the endurance limit from the ASME code curves (Ref. I., Figure 1.0.2.21. With stresses helow this endurance i limit, the normal operating loads are not expected to propogate a flaw from the arrestor hole. The concern regarding crack stability deals with an unceacked liganent - which is left at the inspection slot (See Figure 7.3-11 The section was reduced to a minimal thickness to assure that any crack 7.3-2
- propogation would occur within the slot and thus terminate at the crack arrestor holes.
7.3 5 Fracture Mechanics Evaluation An evaluation of flaw tolerance was conducted by Westinghouse Electric Corporation for the Millstone 2 core
. support barrel. The result of their calculation indicates that fatigue crack growth is limiting in determining the flaw tolerance for this structure.
In the evaluation of operational transients it became apparent that they will not govern the flaw tolerance of the core support barrel as their number of occurrences is small in comparison to other loadings. The key consideration, is the pump excitation, because the number of expected occurrences is approximately 2 x 1011 cycles in the design life. These cycles are dealt with by assuring that the range of applied stress intensity factor due to these loadings falls below the threshold for fatigue crack growth for the 304 stainless steel. Study of the literature indicates this threshold is 6 Ksi3/Inl which results in an allowable flaw size of 8 inches in length. This means that a flaw in this size range could exceed the threshold and therefore be subject to extensive fatigue crack growth, and therefore, is beyond the limit of allowable size. Consideration of fatigue crack growth due to normal operational cycles will reduce the allowable below this 8 inch length requiring the determination of what size flaw could grow to 8 inches in the design life. A fatigue crack growth analysis of the region showed that a flaw 3.6 inches in length would grow to 8 inches in length in the design life due to normal operating loads. When assessing crack extension due to normal operating cycles, the governing loading case was the thermal stress due to the temperature difference across the core support i barrel. This temperature profile includes fluid temperature differences and internal heat resulting from gamma energy deposition. This analysis treated the thermal bending stressee as membrane tension which results in a very , conservative estimation of fatigue crack growth. The conservatively calculated allowable flaw size of 3.6 inches is considerably larger than the flaw detection capability as described in Section 4. For those lug areas which were not machined off, an area inaccessible for inspection exists below the lug. A potentially undetected flaw in this area would be significantly smaller than the allowable flaw size providing margin between inspection capabilities and flaws which could challenge the integrity of the structure. 7 3-3
1?/5/R1 REFEREtiCF (SECTinN 7)
- 1. ASME Section 111,10R3 Friition, Appandix 1.
I l
- 7. 3 - - _ . _ , . _ _ . _
BEj)2 0193] TABLE 7.?-2 Normal Operatin,g Static Hydraulic Loads On The Core Support Barrel Type of~ Load Loading value Loading Condition Axial Uplift Load 569,000 lbs. (See Table 7.2-3 Condition No,'2 Radial Pressure 22.4 psi Condition No. 1 Differential Across (Directed Radially CSB Wall at Thermal Inwards) Shield Lug Elevation ' d e o l l l l l 5
. , , , _ . , _ _ - - _ - - _ , - , , _ . - . . _ - . . . . . . - - , - , . . . . _ . . ~ . . . . . . . . . _ _ _ _ , . . . .
l 30201983 TABLE 7.2-3 L Normal Operating Conditions [ For _ Calculating Hydraulic Loads Condition No. 1 Condition No. 2 Parameter for Maximum Loading For Minimun Loading Inlet Temp Snn'F SAR*F System F1ow Rate 422,000 gpm 377,000 gpm Power Level Zero Power 2700 MWT i h l
12/5/83 TABLE 7.2-4 Pressure Fluctuations At The Inlet Nozzle Station T = h4R'F Pump Characteristic Peak Pressure Frequency Fluctuation, Pinlet Description Value psi
- 1. Rotor Speed 15 ,10.115
- 2. '2x Rotor Speed 30 ,10.05
- 3. Blade Passing 75 +0.6A
,4 ?x Blade Passing 150 +0.n8 l
l l l l
j r munt / .e-1 NORMAL OPERA-*NG LA ERAL HvDRAUL:0 LOA!;5 ON "HE 00RE $UPSOR* SARREL 8 Tu = 500 F; O = 120 of QDt3 3 As,139ltM LgvRL .c7 Ax!s THACUGH NOT LESS A a m t 4 . I r o' G2 w. I 09 3 Lt p/tu gg3
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I l PMBURE 7.2.2 ' l ANNULAR SPACI BETWEEN Rv/C33 NORMAUZED 12/5/83 PRESEURE POWER SPECTRAL DENSITY w l NCRMAUZED PREQUENCY l 1 l . If P *pseny
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1 12/5/83 FIGURE 7.3 1 Outline of Proposed Slot
" Radius h ,- % h 'i
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Flaw Surface PLAN VIEW AT CSB SURFACE PRIOR TO MILLING T f '
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I?/5/83 TABLE OF CONTENTS CHAPTER R SECTION ' SUBJECT 8.0 SAFETY ANALYSIS
8.1 INTRODUCTION
8.2 EFFECT OF THERMAL SHIE'to REMOYAL 8.2.1 PRIMARY SYSTEM FLOW RATE 8.2.2 CORE INLET FLOW OISTRIBUTION 8.2.1 IMPACT ON SAFETY ANALYSES 8.3 EFFECT OF CORE SUPPORT BARREL REPAIR 8.3.1 BYPASS LEAKAGE RATES
- 8.3.2 ASYMMETRIC L'0 ADS EVALUATION 8.3.3 BYPASS FLOW FUEL IMPINGEMENT CONSIDERATIONS 8.3.3.1 IMPINGEMENT CONSIDERATIONS OUTLINEO 8.3.3.2 FLOW TESTING 8.3.3.3 FLOW IMPINGEMENT ACCEPTANCE CRITERI A l 8.3.4 IMPACT ON SAFETY ANALYSES 1
8.3.4.1 NON-LOCA ANALYSES l 8.3.4.2 SMALL BREAK LOCA ANALYSIS 8.3.4.3 LARGE BREAK LOCA ANALYSIS i i
-s 12/5/R1 8.0 SAFETY ANALYSIS A.1 INTRODUCTION Removal of the thermal shield and repairs to the core support barrel were evaluated for impact on primary system flow rate, core inlet flow distribution, and core bypass leakage rates. The details of these evaluations are discussed in Sections A.2 and A.3.
The effect of the thermal shield removal and the core support barrel repairs on the core thermal 'and hydraulic conditions is a slightly higher core flow due to the removal of the thermal shield and a slightly lower core flow due to the increased bypass flow through the crack arrestor a holes. In any event the net result on actual operation is that the core thermal hydraulic conditions change by an insigr.ificant amount. Sections A.2,3 and A.1.4 show that the removal of the thermal shield and the core support barrel repairs do not significantly impact the core thsemal hydraulic conditions and that the Cycle 6 safety analyses submitted in References 1 and 2, as updated in References 3 and 4, remain applicable. , e l 8.1-1
12/5/83 REFERENCES (SECTION 8.1)
- 1. W. G. Counsil letter to R. A. Clark, dated October 22, 1982.
- 2. W. G. Counsil letter to R. A. Clark, dated April 13,~1983.
- 3. W. G. Counsil letter to J. R. Miller, dated November 2,1983.
- 4. W. G. Counsil letter to J. R. Miller, dated November 17, 1983.
A e R.1-2
17/6/R3 R.2 EFFECT OF THERMAL SHTELD REMnVAL R.2.1 Primary System Flow Rate Removal of the thermal shield results in a net decrease in reactor vessel pressure drop at normal operating conditions with a total reactor system coolant flow rate of 35n,nno gpm / minimum guaranteedi. The reduced resistance in turn results in a small increase in reactor coolant system flow rate of less than one percent. The net effect will be somewhat less due to the reduction in primary system flow due to steam generator plugging and sleeving. Tha core hydraulic uplift forces following removal of the thermal shield are bounded by the results of evaluations performed for Cycle A operation when NMErn changed fuel vendors for Millstone 7 Tha reactor coolant system flow rate for Cycle 4 was greater than that expected for Cycle 6 due to additional steam generator tube plugging. The effects of reactor coolant system flow on the reactor internals due to the thermal shield removal ara expected to ha bounded by operating characteristics of plants identical to Millstone 7 which do not have a thermal shield. R.7.? Core inlet Flow nistribution Removal of the thermal shield has a negligible impact on the core inlet flow distribution. This conclusion is raached based on examination of scaled reactor flow model test results and on operating experience with several C-E reactors that do not have thernal shields. Scaled flow model tests have been run on a C-E a-pump reactor ' configuration, both with and without a thermal shield. Examination of the core inlet flow distributions show that resulting differences in the assembly inlet flow factors are very small. Further, there are no systemmatic trends in the changes in th< flow factors that could be related to the removal of the thernal snield. It is concluded from these 8.2-1
12/5/83 test results that the thermal shield removal has no systematic impact on core inlet flow distribution. In other reactor flow model tests on C-E 3410 Nt class reactors, which are very similar in geometry and layout to Millstone 2, the core inlet flow distribution was also measured without a thermal shield. The resulting core inlet flow distribution was relatively uniform and indicated that an acceptable flow distribution can be achieved without the presence of a thermal shield. The relatively uniform core inlet flop distribution is attributed to the flattening effects of (1) the hydraulic resistance of the flow skirt, (2) the lower support structure bottom plate, and (3) the core support plate, all located between the exit of the downcomer annulus and the core inlet plane. Finally, successful operation of other 2560 Mwt class reactors, which are sinilar in construction to Millstona 2 and also operate at 2700 Nt (" stretch" rating 1 as well as the 3410 Nt class reactors indicates no unusual or unacceptable thermal hydraulic performance for C-E reactor designs without thermal shields. 8.2.3 Impact on Safety Analyses The impact of removing the thermal shield from Millstone ;f on the docketed safety analyses has been assessed and presented to the NRC Staff in References 3 and A of Section R.I. A summary of those assessments are provided herein for completeness. It is concluded that operation without ! a thermal shield is acceptable based on the results of the currently docketed safety analyses. The effect of the removal of the thermal shield on non-LOCA transients l has been assessed. The primary results of this modification is to enlarge the downcomer region flow area and volume. The enlarged downcomer flow area and volume produces less resistance to flow in the downcomer area which increases core flow velccities. Increased flow results in a slight penalty for the system transient cooldown events and R.2-2
17/5/83 a steamline break, however this slight penalty would be more than offset by the resulting benefit with respect to DNB. Increased flow would also tend to cause a more rapid pump coastdown impacting transients such as the loss of flow / locked rotor transients, however this effect would be more than offset by the DNB benefit resulting from higher flow. The impact of this modification on the safety assumptions and results presented in the BSR and subsequent reanalysis is negligible; therefore,the conclusions of the previous safety analysis remain valid and support Cycle 6 operation without the thermal shield. An evaluation of the impact on the Small Break LOCA ECCS performance due to removing the thermal shield from Millstone 2 has been completed. Removing the thermal shield from the reactor vessel annulus improves the
- small break LOCA response in that approximately 6000 lbs of additional y
j- liquid coolant will be present in this region during the transient, j Thus, following a small break LOCA, more water inventory is present in the annulus which would delay corp uncovery for all of the breaks analyzed in the small break LOCA spectrum for Millstone 2. That is, an additional amount of liquid would have to be "hoiled off" in the core to reduce the core level to the same elevation as that for the case with the shield present. The additional time to core uncovery will reduce the amount of decay energy in the fuel resulting in lower peak clad temperatures prior to core recovery by the safety injection system. The Cycle 6 large break LOCA analysis has been evaluated to determine ' the impact of the removal of the thermal shield at Millstone 2. The blowdown transient and the reflood transient portions of the accident scenario were specifically evaluated. Thermal shield removal impacts the following portions of the blowdown model.
- 1. Reduces downcomer metal heat and heat transfer surface.
- 2. Enlarges downcomer region flow area and volume.
R.2-3 . l l .- ~ -r r s ,-----w~.-w, . . - - --. e--g,.~- - -> ,e-,-e.,,,.e, n.. m-evv mev,e,.,-- w--,w-, m,----,we,--~ ,,~,---ww.-nv --,.-e+~e--, a w , ,,
12/5/R3 Note that since the peak clad temperature (PCT) for Millstone 2 occurs du-ing the reflood portion of the transient, the impact of the modification during the blowdown portion of the transient will generally have a second order effect on the ultimate PCT. Reducing the downcomer metal heat provides for cooler flows through the core resulting in improved core heat transfer prior to core flow rev ersal . After core flow reversal, the break discharge flow to the containment will be at a slightly lower enthalpy, which tends to reduce energy discharged to the containment. Enlarging the downcemer flow area (additional 10.5 ft )2 and volume (additional 120 ft 3) by removing the shield has the following effects
= on the large break LOCA:
- 1. Less resistance to flow in downconer allowing reversed core flow velocities to increase which benefits core heat transfer.
- 2. The additional 12n ft 3of downcomer fluid is discharged to containment increasing containment pressure in the early portion of the transient. (This effect more than offsets cooler discharge flows due to reduced metal heat).
t
- 3. Larger downcomer flow area potentially provides for earlier end of bypass and reduced safety injection tank (SIT) deficit, due to easier penetration of SIT water into the downtomer.
l The expected net result during blowdown of the effects described above would be a negligible to small decrease in calculated maximum clad I temperature at the end of the blowdown phase of the accident. The thermal shield removal impacts the following reflood model assumptions: R.2 4
i 12/5/A3
- 1. Reduces metal heat in downcomer.
- 2. Enlarges downcomer volume and flow area.
The major effects of these differences on the large break LOCA are as follows:
- 1. Less metal heat in the downcomer increases the lower plenum fluid subcooling, which improves core reflooding rates and acts to reduce the PCT.
4
- 2. Increased downcomer volume (120 3ft additional) requires more STT water to fill the downcomer early in the reflood transient. However, an evaluation of the current Millstone 2 worst large break LOCA f
analysis demonstrates that ample additional SIT water is available, such that the dcancomer will still be filled by the SIT, providing additional SIT cold water that will not spill to the containment. The net effect will be an increase in the containment pressure which improves the core reflooding rate and reduces PCT. A slight delay in the filling of the downcomer (1-2 seconds) early in reflood, due to the additional volume, has a negligible effect on the reflooding transient that follows. The expected net result of the effects during reflood described above is a negligible to small decrease in the ultimate PCT during reflood. Based on this evaluation, it is conservatively concluded that removal of t the thermal shield at Millstone 2 can be expected to have a negligit,le effect on the overall calculated PCT for the large break LnCA analysis. ! Furthermore, a PCT benefit from this plant modification could be expected. The Cycle 6 large break LOCA analysis predicts a PCT of 2n;5'F demonstrating significant margin to the acceptance criterion specified in 10CFR5n.'46 Based on the negligible impacts of the thermal shield removal on the large break LOCA analysis results, it is concluded that 4 the exi:, ting large break LOCA analysis of Reference 1 (Section 8.1) remains limiting. 8.2-5 1
12/5/83 8.3 EFFECT OF CORE SUPPORT BARREL REPAIRS 8.3.1 Revised Bypass Leakage Rates As discussed earlier, following removal of the thermal shield, close inspection of the core support barrel revealed cracks requiring repair. Specifically, two through wall cracks were identified one each on Lugs 4 and 5. To prevent further propagation of the cracks under continued service conditions, crack arrestor holes were machined through the core support barrel at the tips of the cracks. Four crack arrestor holes, 1.125 inches in diameter, will remain unplugged in the core support
. barrel. The four holes represent an additional bypass flow path of
- approximately 3.97 in2 . The bypass leakage through this flow path is calculated to be 0.167, of vessel flow at normal operating conditions.
A summary of'the core bypass flow at normal operating conditions following core support barrel repairs is given in Table 8.3-1 and shows that the total bypass flow remains bounded by the bypass flow assumption of 3.77. utilized in the safety analyses. 8.3.2 Asymmetric Loads Evaluation An asymmetric loads evaluation was performed for Millstone 2 as part of a ! generic owners group activity in response to Unresolved Safety Issue A-2. , The evaluation for asymmetric loadings includes the effects of l differential pressure across the core support barrel on the reactor internals and fuel. The addition of four,1.125 inch diameter holes in the core support barrel has no significant effect on the calculated pressure difference across the core support barrel. The small potential difference between the results calculated for an intact core support I barrel and the Millstone 2 core support barrel, as repaired, will be in a direction reducing the differential pressure across the core support barrel. i R.1-1 l l t
12/b/8.5 The remaining ligament section requirements for the non-through wall crack in Lug 4 which was machined out are based on membrance stress calculations for the differential pressure across th core support barrel as a result of asymetric loads. 8.3.3 Bypass Flow Fuel' Impingement Considerations 8.3.3.1 Impingement Considerations Outlined The repair of the Millstone 2 core support barrel consistr, of drilling four,1.125 inch diameter crack arrestor holes at the tips of the two through wall cracks. One crack was located on Lug 4 and one on Lug 5 of the core support barrel. The crack location at Lug 4 of the core support barrel is in the vicinity of a seam in the core shroud. The core shroud is designed as four 90 degree sections with a maximum seam size of 15 mils between sections. Eight former plates surround the shroud circumferential1y and provide structural support. Figure 8.3-1 illustrates the core shroud design. One crack arrestor hole on Lug a will provide a bypass flowpath in the vicinity of one of the core shroud seams. The Millstone 2 design is such that coolant flows from within the core, through the core shroud seams, into the core shroud / core support carrel annulus. The one crack arrestor hole provides a potential for bypass flow impingement on the core shroud seam, through the seam and onto the fuel pins in the vicinity of the seam. Figure 8.3-2 illustrates the geometry described i abov e. l Analytical assessments were performed and actual flow tests were conducted to determine the effects of the core barrel repair on the fuel in the vicinity of the core shroud seam. These efforts are discussed in the following sections. 8.3-7
~_
l 12/5/83 8.3.3.2 Flow Testing A core shroud / core barrel mock up is being tested by Creare R A D, Inc. The tests are designed to determine the worst expected jet magnitude and profile which could result from a 1-1/8" crack arrestor hole drilled in'the core barrel. The target of this jet is the fit-up seam in the core shroud haffles and ultimately fuel rods interior to this seam. The driving pressure for the jet is consistent with the expected downcomer pressure. Creare has also been calculating velocities and jet profiles with their proprietary computer code FLUENT. The preliminary findings of both the testing and calculations show that the worst expected jetting condition is bounded by the fuel vendor criteria for fuel damage due to haffle jetting. The test mock-up consists of a modeled section of the core shroud and core barrel. The shroud assembly models three former plates and six baffle plates. The model approximates 3n' of the shroud circumferentially and represents two vertical sections. A 1-1/8" hole was drilled in the core barrel mock-up just below the middle former ~ plate. A plenum was attached over the hole and water was pumped into the plenum to .naintain jet driving pressure. The baffle plate seam was set at .n15". A manual traversing of the seam with a pressure probe established an approximate magnitude and a jet profile. The profile measured was consistant with the analytical prediction. Stationary pressure probes were then located in the vicinity of the profile peak i to establish a more accurate local profile. The measured velocities e and profiles for the limiting case were evaluated by the fuel vender to determine the impact on the fuel due to baffle j etting. Finalized velocity profiles from testing and computer simulation are expected by early December,1483. Variations from the preliminary profiles are not expected to be significant and are not expected to invalidate the fuel vendor criteria for fuel baffle jetting. Appendix A to this report provides additional details regarding the flow testing performed to support the core support barrel repairs. 8.1.1 i
12/5/83 8.3.3.3 Flow Impingement Analytical Evaluation As a result of the decision not to plug the crack-arrestor holes placed in the Millstone 2 core support barrel, the potential exists for water jets to impinge on a shroud seam, resulting in flow jetting on the fuel rods adjacent to the seam. Because such cross flows have been known to result in fuel rod failures at other plants, an analysis was performed to assure that the expected cross flow jetting would not lead to fuel rod instability and subsequent fuel rod failures. Utilizing experimentally determined velocities and velocity profiles, the fuel vendor has analytically determined vibration mode shapes and performed a stability analyses. It was determined that for the cross flow jetting conditions analyzed, the fuel rods would not fail. Additional details of the fuel integrity analyses performed to support the core support barrel repairs are provided in Appendix A. 8.3.4 Impact on Safety Analyses The nnn-LOCA and 10CA analyses have been reviewed to determine what, if any, impact the core support barrel repairs have on the safety analyses results docketed to support Cycle 6 operation. Specific references for these results are provided in Section 8.1. The impact of removing the thermal shield on these analyses was addressed in Section 8.2.3. 8.3.4.1 Non-LOCA Analyses As was presented in Section R.3.1, the increase in core bypass flow l I resulting from the core support barrel repairs is 0.16",of reactor vessel flow during normal operation. The total best estimate core i bypass flow following core support barrel repairs is presented in Table 8.3-1 as 2.67",. This value is bounded by the current design bypass flow rate assumption of 3.7% used as input to the safety analyses. It is concluded that the current non-LOCA transients and accident results remain appropriate for Cycle 6 operation. ' R.3 4
.)
12/5/83 8.3.4.2 Small Break LOCA Analysis A review was conducted to assess the effect of four 1.125 inch diameter holes in the core support barrel on small break LOCA's for Millstone 2. The addition of these holes represents a new but small flow communication path between the vessel downtomer and core shroud bypass region actj acent to the core. The total flow area (3.97 in2), however, is very small and the flow resistance very high. Therefore the mass transfer rate is slow. At steady state normal operating conditions the flow through these holes is about 0.16", of the total vessel flow. Following a small break, the reactor coolant pumps are tripped following SI AS causing a much lower and reversed pressure gradient across the core barrel and a much lower and reversed flow rate through the holes. The small amount of mass transfer which might occur through these holes during a small break does not alter the mass lost from the system since the total loss is controlled by the critical flow conditions of the break. The holes, however, would allow a small amount of liquid redistribution from the cor,e to the downcomer until the liquid level falls below the holes. The ' level falls below the holes soon after the core begins to uncover since the holes are only about 13 inches below
- the top elevation of the fuel. As shown by the current analysis of a spectrum of small breaks for Millstone 2, whenever the mixture level in the core is at or above this elevation, steam cooling of the uncovered portion of fuel is adequate to keep clad temperatures low (below 1100'F). This is more than R00'F below the worst case (0.1 ft2 break) peak clad temperature. Therefore, even if all the liquid above the holes were assumed to redistribute, the result is acceptable.
l The peak clad temperature for these small breaks occurs when the minimum core mixture level occurs. This minimum level is below the elevation of the holes and is unaffected by the holes. The minimum 8.3-5
. 12/5/R3 level is simply the result of achieving a match between treakflow and safety inj ection inflow. Hence, the predicted peak clad temperatures for the current analyzed spectrum of small LOCA's is unaffected.
The current licensing analysis remains a conservative representation of compliance with the acceptance criteria of 10CFR50.46 for Millstone 2. 8.3.4.3 Large Break LOCA Analysis An evaluation of the impact of four 1-1/8" -diameter holes in the core barrel on the large break LOCA peak clad temperature (PCT) for Millstone Unit No. 2 is complete. The barrel holes function as an additional by-pass flow path during the blowdown transient. The axial core flow rates are predicted to be reduced by -0.3% due to the presence of the barrel holes. This small blowdown flow reduction is estimated to be an approximate 20F penalty on the LOCA PCT. The impact on the reflood transient is negligible since the downcomer head remains unchanged. It is concluded that the Impact of four 1-1/8" -diameter holes on the core barrel is expected to be less than a 200F PCT penalty for the large break LOCA. Since the current LOCA PCT for Millstone Unit No. 2 is 2055 F, 10CFR50.46 limits will.not be exceeded due to this modification. Therefore, the current large break LOCA analysis for Millstone Unit No. l 2 remains applicable. 8.3-6
1?/5/83 l TABLE R.3-1 Millstone ? Core Byp a s Flow Distribution FLnW PATH CORE RYPASS FLnW-
- OF VESSEL FLOW
- 1. Alignment Keys n.no P. nutlet Nozzle Gap n.11 3 Core Shroud Holes n.?7 seams n.23 A. Guide Tubes 1.6n
- 5. Surveillance Holes n.10 6 Four 1-1/R in, dia, holes in CSR n.16 Best Estimate Total Bypass Flow 7.67 nesign total hypass flow utilized in 1.7 safety analyses i
i (
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( 12/5/R3 TABLE OF CONTENTS CHAPTER 9 SECTION SUBJECT 9.0 MONITORING AND INSPECTION PROGRAMS
9.1 INTRODUCTION
9.2 MONITORING ORJECTIVES 9.3 EXCORE MONITORING 9.3.1 SYSTEM DESCRIPTION 9.3.2 MONITORING PROCEDURE 9.3.7.1 BASELINE MEASt!REMENTS Q.3.2.2 SURVETLLANCE MEASUREMENTS 9.3.2.3 CORRECTIVE ACTION 9.4 LOOSE PARTS MONITORING l 4.4.1 SYSTEM DESCRIPTION 9.4.2 DATA ACOUISITION 9.a.3 DIAGNOSTICS 9.5 INSPECTION PROGRAM 9.6 STARTilP CALIBRATIONS
l 12/$/nS 9.0 MONITORING AND INSPECTION PROGRAMS
9.1 INTRODUCTION
Two diagnostic systems are used for nonitoring the integrity of the reactor coolant system internals during normal power operation. These diagnostic systems are the Internals Vibration Monitoring (IVM) system and the Loose Parts Monitoring (LPM) system. The IVM system monitors the signal output or " noise" from six summed (upper plus lower) linear power range excore neutron flux detectors. Any motion of the core / core support barrel causes a neutron level modulation which is monitored by the IVM system. The IVM system provides the signal conditioning and statistical processing necessary to provide estimates of core / core support barrel vibration. The LPM system monitors the output from accelerometers mounted on the external surface of the reactor vessel and steam generators. Metal to metal impacting within the primary system causes high frequency impulse vibractions on the reactor system shell which are detected by the LPM system. l l l l 0.1-1 i l
~, _ _ _ ., _
( i 12/5/83 9.2 MONITORING OBJECTIVES Both the IVM and LPM systems provide diagnostic information about the integrity of the reactor internals. The primary monitoring objective is to use this information to provide as early an indication as possible of degradation of failure of reactor internals components. The state of the art for these monitoring systems has not progressed to the point where they can provide an inmediate and absolute indication of an internals failure. However, the systems can provide considerable infornation to personnel trained and experienced in diagnosis of " reactor noise monitoring." There is a trade-off between monitoring sensitivity and the incidence of expected false indications. The monitoring philosophy described here is to maintain a high degree of sensitivity and accept some false indications which will require disposition by an experienced diagnostician. l t I i l 9.2-1
I i 12/5/R3 9.3 EXCORE MONITORING . Excore monitoring are used with the Internals Vibration Monitoring (IVM) system. During the current outage, the system has been thoroughly checked and tested and found to be in satisfactory working condition. We are also evaluating the replacement of certain components with current state-of-the-art components. 9.3.1 System Description The IVM is an automated, computer-based system specifically designed to acquire, analyze, and interpret small variations in excore nuclear detector signals for the purpose of identifying reactor core and core support barrel motion. The system is based on a Hewlet Packard 5451-B Fourier Analyzer capable of accepting analog information, digitally converting it and processing it in both the time and frequency domains. The IVM system provides descriptions of the motion in the frequency domain, its amplitude probability distribution, the l azimuthal character of the motion, and a quantitative measure of the root mean square (RMS) vibration amplitude. The system also provides the capability of selectirig a wide range of statistical analysis functions providing considerable flexibility when diagnosing anomalcus signals. The IVM system has been programed for three modes of operations: (1) continuous monitoring of internals motion by comparison of calculated RMS values of excore signal amplitude to preset limits, (2) calculation and plotting of amplitude probability density ( APD) l functions for each of the six excore signals, and (3) Fourier analysis consisting of calculation and plotting of power spectral density (PSD). cross PSD, coherence, and phase measurements for each of three pairs of excore detector signals. l 9.3-1
12/5/83 Mg2 Monitoring Procedure
? comprehensive re-evaluation of the monitoring procedure will be carried out during the next fuel cycle. The primary objective is to discern the changes caused in the frequency response of the internals by the removal of the thermal shield and establish a baseline for future monitoring and diagnosis.
The monitoring procedure is in three phases: baseline measurements, surveillance measurements and corrective action. 9.3.2.1 Baseline Measuremer.ts
~ Tne ohj ective of this phase is to obtain sufficient data to: ,
a) identify the frequencies corresponding to the predominant modes of the core barrel motion. b) determine normal variations in the frequency and amplitudes (percent noise) of these modes, and c) evaluate alert and alarm limits for these modes based on allowabic variations about the baseline values. These limits will be expressed in terms of the RMS values for the frequency bands encompassing these modes. The establishing of limits for these modes will require collection of l sufficient data and off-line analysis to accurately determine the limits of normal variation in these modes. This detailed baseline phase will be for the next fuel cycle only. Subsequent cycles will use alert and alarm limits established during this cycle. Baseline tests for these cycles will consist of obtaining spectra for all cross-core and adj acent pairs of detectors as well as coherence and phase data. 9.3-? L
( 12/5/33 9.3.2.2 Surveillance Measurements The objective of the surveillance is to monitor dynamic response of the core support barrel for changes that may be indicative of changes in the motion of the core barrel. This surveillance will be done by computing the narrow band RMS values within the frequency ranges established in the baseline phase. Comparisons will be made with the alert and alarm levels. In addition, frequency spectra consisting of PSD's, CPSD's, coherence and phase of cross-ccre detectors pairs will be examined. The frequency of surveillance for the next and subsequent cycles will be established later. 9.3.2.3 Corrective Action Exceeding an alert or a known level will require timely action to determine'if the variation is spurious or not. If it is spurious, the alert / alarm levels will be adj usted to account for these changes. If not, more analysis and confirnatory evidence of abnormality will be sought and appropriate corrective actions taken. 0.3-3
I 12/5/R3 9.4 LonSE PARTS MONITORING The Loose Parts Monitoring System (LPMS) detects vibrations on the external surface of the reactor vessel and steam generators. In addition to normal " background" vibration caused by mechanical and flow induced sources, the system detects impulse like variations caused by metal to metal impacting within, on, or near the primary system. 9.4.1 System Description The original system has been completely refurbished, modifi(d, and calibrated during the Cycle 6 refueling outage. The LPMS monitors vibration levels on the reactor vessel and steam generators using high temperature accelerometers designed for radiation environments. Acceleration signals detected by the accelerometers are converted to their voltage equivalent and amplified by charge amplifiers located outside of reactor containment. The amplified voltage signals are fed to the LPMS panel in the control room for continuous evaluation. j Each accelerometer is an Endevco type 2273AM1 with a nominal sensitivity of 12pC/g. The accelerometers are stud mounted to 2 inch diameter multipole magnets. The magnet is affixed to its location by both magnetic force and high temperature ceramic adhesive. This ensures a linear frequency response from very low frequency to at least 6kHz to 10kHz. This is adequate for loose parts detection. i Signals from all eight accelerometers are first processed and amplified by B A K Charge Amplifiers, type 2634, at the outside of the containment structure. The amplified voltage signals from each charge amplifier are fed to the control room LPMS. The LPMS provides two separate functions. i 9.4-1 .
( 12/5/83 First, the signals from each channel pass through a signal amplifier. The output of this amplifier is passed through a filter and then a detection circuit. Filtered signals of sufficient amplitude and duration trigger an alarm circuit. The alarm circuit also trips a control panel annunciator. The second function of the LPMS is to continuously record the vibration signals from all 8 transducers. This is done using " endless loop" tape recorders. Input to the recorders is then output of the first stage LPMS amplifier. 9.4.2 Data Acquisition During the refueling period data has been collected on the response of the reactor vessel. This is being done using known force inputs fron a 3 pound calibration force hamer. The vessel is inspected at known locations,and signals are collected from the mounted LPMS accelerometers. The force input spectra and acceleration output spectra are divided to produce plots of vessel mobility and phase lag as a function of separation between impact location and transducer. This data relates the acceleration level measured by an accelerometer with a force caused by a moving obj ect impacting the vessel. Wave propagation velocities are also being measured between various points on the vessel using this technique. During the startup process background vibration level data will be collected for each of the operable LMPS channels. Background noise due to normal operation of the plant will be quantified through detailed mesurements. These background vibration levels will be used to determine the minimum setpoint levels which will not cause an excessive number of spurious alarms. 9.4-2 _______= .- -_ .
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. . . 17/5/R3 I
During the coming fuel- cycle detailed background noise level data will be collected. This data will form the overall baseline conditions from which all loose parts evaluations will be conducted. All containment cabling and the accelerometers will be checked for proper installation at regular intervals. The LPMS will be placed in an operational mode during plant startup.L Because of the lack of background noise level data, initially the LPMS will be set to a relatively low sensitivity. Gradually, during the startup process the sensitivity of the LPMS will be increased until substantial false alarms occur. The sensitivity will then be decreased until the false alarm rate diminishes. The background noise levels causing the false alarms will be analyzed. The LPMS will then be set to this sensitivity level during normal operation. The firm of Bolt, Reranek, and Newman is presently investigating the false alarm
-issue for NNECO. The results of their parametric study on loose part detectibility and false alarn rate will be available for guidance during the start-up process.
9.4.3 Diagnostics Since the LPH is sensitive t' o external vibration:, of the primary system, it will generate alarms for many events besides the occurrence of a loose part. The setting of absolute alarm limits which require i plant shutdown would tend to limit the effectiveness of the early indication of an internals failure, because the alarm limits would j have to be rather insensitive to preclude unacceptable frequency of false alarms. However, if persistent LPM system alarms should occt,r, the source will be evaluated if possible by plant personnel, i Experienced diagnosticians may also be consulted to evaluate such alarm indications. 9.4-3
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[ [. . 1?/5/R1 t 4.5 INSPECTION PROGRAM Damage to the core barrel was found at Lugs a and 9 The non-through wall crack at Lug 4 has been machined out and the through wall cracks at Lugs 4 and ; have been arrested by crack arrestor holes. Analytical evaluations demonstrate that these cracks will not propagate. However, in order to verify this, inspections are being evaluated for next outage. Visual examinations of the core barrel at the damaged locations are being evaluated. Feasibility of utilizing computer enhancement of video signals to obtain pictures of greater clarity are currently being investigated. i t 9.5-1
17/5/83 9.6 STARTUP CALIRRATIONS The removal of the therinal shield and the core barrel rep f rs to Lug areas 4 and 5 will result in changes to certain mr asured parameters. In particular, the reactor coolant system (RCS) flow through the core will increase by a small amount since the remcyal of the thermal shield reduces the flow resistance in the reactor vessvl downcomer. Likewise, the removal e of the three inch thick steel thermal shield will alter the response of the excore neutron detectors. NNECO plans to reestablish the measured parameters affected by the thermal shield removal and core barrel repair program. RCS flow will be recalibrated at the beginning of Cycle 6. The calibration will be performed in the same manner as in previous cycles when recalibration was required due to steam generator tube plugging. The absolute value of RCS flow is di:termined from secondary side calorimetric power and primary side temperature differences. The Reactor Protection System (RPS1 low flow trip setpoints are subsequently reset based upon the measured flow. The net change in RCS flow from Cycle 5 is expected to be very small(less than IQ l because the increase in flow due to the removal of the thermal shield will be offset by a similar decrease in flow due to steam generator tube plugging and sleeving. ~ l Nuclear power is expected to be conservatively decalibrated due to the removal of the thermal shield. The excore detector response is expected to be somewhat larger for the same power level and power distribution. The power range nuclear channels will be recalibrated in the same manner as in previous cycles at approximately 10'. , 2 0', , 30", , and Sn", power. The results of these calibrations will determine the required frequency of calibration ! above sn*, power. Axial shape index (ASI) is not expected to be affecter' by the thermal shield removal. NNECn plans to recalibrate ASI in the same manner as an previous cycles. Additionally, an induced xenon oscillation test will be performed in order to verify the shape annealling factor constants used in the calibratior. of ASI. l 4.6-1 l}}