ML20076H314

Forwards Response to Core Performance Branch Draft Ser.Draft Page Changes Will Be Included in FSAR Rev Scheduled for Jul 1983

ML20076H314
Person / Time
Site:	Limerick
Issue date:	06/13/1983
From:	Bradley E PECO ENERGY CO., (FORMERLY PHILADELPHIA ELECTRIC
To:	Schwencer A Office of Nuclear Reactor Regulation
References
NUDOCS 8306160527
Download: ML20076H314 (91)

Text

u e

9 PHILADELPHIA ELECTRIC COMPANY 23O1 M ARKET STREET P.O. BOX 8699 PHILADELPHI A. PA.19101 E DW A R D G. B AU E R, J R.

'.*o'l"!!l,".".".'ou..s h CUG ENE J. BR ADLEY assoceava esmanaL cousesst DON ALD BLANKEN CUDOLPH A. CHILLEMI

c. C. MI R K H A LL T. H. M AHER CORNZLL PAUL AUERB ACH messsvAnt asusn AL couwssk EDW ARD J. CULLEN J R.

THOM AS H. MILLER, J R.

g IRENE A. McKENN A ASSIST ARY CouMSaL Mr. A. Schwencer, Chief Licensing Branch No. 2 Division o f Licensing U. S. Nuclear Regulatory Commission Washington, D.C. 20555

Subject:

Limerick Gene ating Station, Units I&2 Response to Core Performance Branch Draft Safetyr,' valuation Report

References:

1) A. Schwencer to E. G.

Bauer, Jr.,

Letter dated March 11, 1983 2) E. J. Bradley to A. Schwencer, Letter dated May 25, 1983 File: GOVT l-1 (NRC)

Dear Mr. Schwencer:

Our May 25, 1983 letter advised you that Chapters 4, 5, 6, and 15 of the FSAR would be revised to incorporate the latest approved revision o f " General Electric Standard Application for Reactor Fuel" (GESTAR II). Attached are those draft page changes which are now scheduled to be included in the July 1983 FSAR revision. These changes will appear in the FSAR exactly as shown on the attachments. Sincerely, E en J. Bradley Boot RJS/gra/12 Copy to: See Attached Service List 8306160527 830613 ~' PDR ADOCK 05000352 E PDR

6 F Judge 'awrcuce Brenner (w/o enclosure)' L cc: Judge Richard F._ Cole (w/o enclosure) Judge Peter A. Morris (w/o enclosure) Troy B. Conner, Jr., Esq. (w/o enclod'ure) Ann P. Hodgdon (w/o enclosure) Mr. Frank R. Romano (w/o enclosure) Mr. Robert L. Anthony (w/o enclosure) Mr. Marvin 1. Lewis (w/o enclosure) Judith A. Dorsey, Esq. (w/o enclosure) Charles W. Elliott, Esq. (w/o enclosure) Jacqueline I. Ruttenberg (w/o enclosure) Thomas Y. Au, Esq. (w/o enclosure) Mr. Thomas Gerusky (w/o enclosure) Director, Pennsylvania Emergency Management Agency (w/o enclosure) Mr. Steven P. Hershey (w/o enclosure) Donald S. Bronstein, Esq. (w/o enclosure) Mr. Joseph H. White, III (w/o enclosure) David Wersan, Esq. (w/o enclosure) Robert J. Sugarman, Esq. (w/o enclosure) Martha W. Bush, Esq. (w/o enclosure) Atomic Safety and Licensing Appeal Board (w/o enclosure) Atomic Safety and Licensing Board Panel (w/o enclosure) Docket and Service Section (w/o enclosure) e e

r ( LGS FSAR D R A F a' OUESTION 490.1 Since the issuance of NEDO-20944, which is referenced in the Limerick FSAR, GE has developed enother fuel design as descr.ed in NEDO-24011-A-2. Is the applicant's fuel system design like the new fuel design (NEDO-24011-A-2)? If so, provide the information that is needed for the new fuel design.

RESPONSE

[ initial core fuel system design for Limerick Unit 1 is being revised to contain five bundle average enri m Lopposed tM Lthe typical three indicated in NEDO-20944.)chmants. i Limer'ick's fuel bundles M ll e includea in e f;;t h.ww swvielm. t: NED07NEDE-2401 - P-A-5

JAS44, eneral Electric Standard Application for Reactor Fuel."'

are-F D owing unis revision, Philadelphia Electric Company intends to modify the Limerick FSAR to reference NEDO/NEDE-24011. This is l ( (expectedtooccurduringthesecondquarterof1983. IS h M<. lxw changed +o incoqwk C MPTigs % 5; G, AND Pf cd l

• 4 NsDof d eos - 14 0ll ~ P-A-5, ' Ge~<al E l*cir.*c-Standa vi A li

[.< R, a cjo,- F u s. 0 ", infordien. 2 07/8.3 490.1-1 Rev. 82 -

s i, LGS FSAR 3 D R.. A F x :i s CHAPTER 4, EACTOR g g, cal,ep M M E e,Yc. t f A l = f' 7 "", i 4.O INTRODUCTION { (GESTAR4heLdik hhiN bd"ff*"#2$ / (UECE -2Voll-P-43 E0E'Noll4-A *E / - This chapter was prepared utilizingIth ' O ;i"- Te - e-BWR/4 -end -BWR/5 -Feel *' %M ee : pie:1 R g,th -EDO-20944 and-#EDE-*9944-P,--< lated-October ' S j

nd:d 6i-leties -Rr-itrBuchhatr,-GE -to-0.-D.-Park;--NRC, #~

- S-dered-February 45, 7 1;;0, Applicable sections of the-topical t<.l. report are referenced as noted in Sections h-2, -4r3 -and 4 A 4.:+.o - r* t h u U.. u ;..; includwo an unio cnapte & 44 7 ve I I I 9 4.1-1 y ,.-,-w-.

LGS FSAR 4.1

SUMMARY

DESCRIPTION h 3 The reactor assembly consists of the reactor vessel, its internal components of the core, steam separator and dryer assemblies, and jet pumps. Also included in this assembly are the control rods, control rod drive housings, and the control rod drives. Figure 3.9-4, Reactor Vessel Cutaway, shows the arrangement of reactor cssembly. components. A summary of the important design and performance characteristics is given in Section 1.3. Loading conditions for reactor assembly components are discussed in Section 3.9., 4.1.1 REACTOR VESSEL The reactor vessel design and description is discussed in Section 5.3. 4.1.2 REACTOR INTERNAL COMPONENTS The major reactor internal components are the core (fuel, channels, control blades, and incore instrumentation), core cupport structure (including the shroud, top guide, and core plate), shroud head and steam separator assembly, steam dryer cosembly, feedwater spargers, core spray spargers, and jet pumps. Except for the 2ircaloy in the reactor core, these reactor internals are made of stainless steel or other corrosion- ) j rcsistant alloys. Of the preceding components, the fuel casemblies (including fuel rods and ' channel), control blades, incore instrumentation shroud head and steam separator assembl l cnd steam dryers are re, movable when the reactor vessel is open.y, 4.1.2.1 Reactor Core 4.1.2.1.1 General i The design of the boiling water reactor (BWR) core (including fuel) is based on the proper combination of many design variables and operating experience. This contributes to the achievement of high reliability. r I umber of important features of the BWR core design cu rized belows a. e BWR core mechanical design i sed on conservative ap cation of stress limits perating experience, and exper tal test result The moderate pressure level character of a et cycle reactor (approximately I 1000 psia) resu moderatecladdingtemperaturesandli stress level / b. ow coolant saturation temperature, high heat / / transfer coefficients, and neutral water chemistry of s / 4.1-2 i

~_. m;~ LGS FSAR ( the BWR are significant, advantageous factbrs in Nemperature-dependentcorrosionandhydridebuildup'./ minimizing Zircaloy temp'erature.and associated The elatively uniform fuel cladding temperatures thro out the core minimize migration of the hydrides l _f to col cladding zones and reduce thermal stre es. c. The basic ermal and mechanical criteria a lied in the design have een proven by irradiation of atistically significant q ntities of fuel. The des. n heat transfer rates nd linear heat generati n rates are similar to value proven in fuel asse ly irradiation. d. The design power di ribution use in siting the core represents a worst ex cted stat of operation. e. The General Electric the al nalysis basi L GETAB, is applied to ensure that no han 99.9% of the fuel rods in the core are expected void boiling transition for l the most severe abnorma oper tional transient described in Chapter 15. The sibilit of boiling transition j occurring during no al reactor ration is insignificant. { ~f. Because of th large negative modera or density coefficient f reactivity, the BWR ha a number of i inherent vantages. These are the us of coolant flow for los following, inherent self-flatte ing of the l radia power distribution, ease of contro spatial men stability, and the ability to overri menon in o er to follow load. BWRs do ot have instability problems due to menon. Thi ghas been monstrated by special tests which have been conduct d on oper ing BWRs, in an attempt to force the reactor into men in ability, and by calculations. No menon instabilities have e er been observed in the test results. All of the indicators how that zenon transients are highly damped in a BWR due to t large negative power coefficient of reacti - 4r4-tr. -- - j Important features of the reactor core arrangement are as follows: a. The bottom-entry cruciform control rods consist of boron-carbide in stainless steel tubes, surrounded by a stainless steel sheath. Rods of this design have ace'umulated thousands of hours of service in operating BWRs without significant failure. M jixed incore fission chambers provide continuous b. powet-range neutron-flux monitoring. A guide tube in 4.1-3

~ ORAFT LGS FSAR I each incore assembly provides for a traversing ion-chamber for calibration and axial detail. Source and' intermediate range detectors are located incore and are axially retractable. The incore location of the startup and source range instruments provides coverage of the. large reactor core and provides an acceptable signal-to-noise ratio and neutron-to-gamma ratio. All incore instrument leads enter from the bottom and the instruments are in service during refueling. Incore instrumentation is further discussed in Section 7.7. Experience has shown that the operator, utilizing the c. incore flux monitoring system, can maintain the desired power distribution within a large core by proper control rod scheduling. The 'i ::1;; [ reusable channels providd a fimed flow d. ( path for the boiling coolant, serve as a guiding M ace _ for the control rods, and protect the fuel during handling operations. Mechanical reactivity control permits criticality checks e. during refueling and provides maximum plant safety. The core is designed to be suberitical at any time in its operating history, with any one control rod fully . -) withdrawn. f. The selected control rod pitch represents a practical value of individual control rod reactivity worth, and allows adequate clearance below the pressure vessel, between control rod drive (CRD) mechanisms, for ease of maintenance and removal. 4.1.2.1.2 Core Configuration The reactor core is arranged as an upright circular cylin' der containing a large number of fuel cells, and is located within the reactor vessel. The coolant flows upward through the core. The core arrangement (plan view) and the lattice configuration cre " " G l" G ' W ' 3 3 '- ~ -- - ^ ' ' --^ ^ ' ' ~ ~ - - * * ~ ' ' * - - vs'K ~~~--~~-'--"~'--~; 4.1.2.1.3 Fuel Assemb y Dejr i DeseAM.S ne bl amblda b D<4n f.2. in "As/C4 be een rom gur 4.J i and~4.3 2, he re sed fe ont 11y c pon#ts: e as mbli s a d /ca en rol rods Th fuel asse ly co r r mec ani al ec fi rati ns ( igur 4. 4 an 4.2-da, es ctiv ly) re b si 11y he ea us in D esden'1 di all sub u t B ._ F.the disc ssio is e ntaisied j Se tion 4.2. "2 4.1-4

JAAFT ^ ~ LGS FSAR ]} '] 4.1.2.1.3.1 Fuel Rod A fuel rod consists of. uranium dioxide (UO,) pellets and a A Zircaloy-2 cladding tube. A fuel rod is made by stacking pellets into the cladding tube, which is evacuated and backfil' led with f helium, and sealed by welding Zircaloy end plugs in each end of the tube. The ASME Boiler and Pressure Vessel (B&PV) Code, f Section III, is used as a guide in the mechanical design and stress analysis of the fuel rod. The fuel rod is designed to j withstand applied loads, both external and internal. The fuel pellet is sized to provide sufficient clearance within the cladding tube, to accommodate. axial and radial differential expansion between fuel and clad. Overall fuel rod design is conservative in its accommodation of the mechanisms affecting I fuel in a BWR environment. Fuel rod design bases are discussed in more detail in Section 4.2.1.1.2. l t 4.1.2.1.3.2 Fuel Bundle i i Each fuel bundle contains 62 fuel rods and 2 water rods, which i are spaced and supported in a square (8x8) array by 7 spacers and a lower and upper tie plate. The fuel bundle has two important design features ,l C' The fuel bundle design places minimum external forces on a. a fuel rod; each fuel rod is free to expand in the axial direction. I b. The unique structural design permits the removal and replacement, if required, of individual fuel rods. The fuel assemblies, of which the core is comprised, are designed to meet all the criteria for core performance and to provide ease l of handling. Selected fuel rods in each assembly differ from the, j others in uranium enrichment. This arrangement produces more uniform power production across the fuel assembly, and thus allows a significant reduction in the amount of heat transfer surface required to satisfy the design thermal limitations. (Further discussion may be found in Section 4.2. f 4.1.2.1.4 Fuel Assembly Support and Control Rod Location All peripheral fuel assemblies and their individual peripheral fuel support pieces are su' ported by the core plate. Otherwise, p individual fuel assemblies in the core rest on fuel support pieces mounted on top of the control rod guide tubes. Each guide 1 tube, with its fuel support piece, bears the weight of four assemblies, and is supported by a CRD penetration nozzle in the bottom head of the reactor vessel. The core plate.provides l lateral support and guidance at the top of each control rod guide tube. 4.1-5 i

\\ s. 4< I-(" g,w DRAFT -,SA. \\ The top guide,' mounted inside he shroud, provides lateral cupport and guidance for ea fuel assembly. The reactivity of

• thecoreiscontrolledbygr)cuciform-shaped control rods

trini.; 5::::

hydraulic drive system.:::bif;, and their associated mechanical-between fuel assemblies. The control rods occupy ~ alternate spaces Each independent CRD enters the core from the bottom, accurately positions its associated control rod during normal operation, and yet exerts approximately ten times the force of gravit mode of operation. y to insert the control rod during the scram Bottom entry allows optimum power shapin the core, ease of refueling, and convenient CRD maintenance.g in 4.1.2.2 Shroud Information on the shroud is contained in Section 3.9.5. 4.1.2.3 Shroud Head and Steam Separator Assembly Information on the shroud head and steam separators is contained in Section 3.9.5. 4.1.2.4 Steam Dever Assembly Information on the steam dryer assembly is contained in Section 3.9.5. ) 4.1.3 REACTIVITY CONTROL SYSTEMS 4.1.3.1 Operation The control rods perform dual functions of power-distribution chaping and reactivity control. Power distribution in the core is controlled during operation of the reactor by manipulation of colected patterns of rods. of the near-cylindrical reactor core, are positioned toThe rods, which enter from th counterbalance steam voids in the top of the core and effect cignificant power flattening. uced for power flattening These groups of control elements, cycle and neutron exposure, experience a somewhat higher duty than the other rods in the control cyates. Tho reactivity control function requires that all rods be evcilable for either reactor scram or reactivity regulation. Because of this, the control elements are mechanically designed to withstand the dynamic forces resulting from a scram. connected to bottom-mounted, hydraulically-actuated driveThey are tschanisms which allow either axial positioning for reactivity regulation.or rapid scram insertion. drive conne,ction permits each blade to be attached or detachedThe design of the from its drive, without disturbing the remainder of the control cystem. The bottom-mounted drives permit the entire control I 4.1-6

s. HAFT LGS FSAR peatswith'thereactor system to be left intact and operable for vessel open. 4.1.3.2 Description of Control Rods d A desce,%.F h ccmfet mas is aw.i m sec/A #2.2.1 1 e cruciform-shaped control rods contain stainless steel tubpIl 11ed with vibration compacted boron-carbide powder. The bes 'ar seal-welded with end plugs on both ends. Stainless at el bal are used to separate the tubes into individual comp taents. The stainless steel balls are held in ition by a sli t crimp in the tube. The individual tubes pro de contain nt of helium gas released by the boron-neu on capture reaction. The tubes a e' held in the cruciform array by a ainless steel sheath exten ing the full length of the tubes. A top handle, shown in Figu - 4.2-6, aligns the tubes and ovides structural rigidity at th top of the control rod. Ro ers, housed in the top handle, prov de guidance for control r insertion and withdrawal. A bo om casting is also us to provide structural rigidity, and cont 'ns positioning roll s and a parachute-shaped velocity limiter. e handle and lowe casting are welded into a single structure by m ns of a small ruciform post, located in the center of the conte 1 rod. A s ainless steel stiffener is ( located approximately at he mids n of each cruciform wing. The control rods can be pos to ed at 6-inch steps and have a i nominal withdrawal and insert n speed of 3 inches /second. The velocity limiter is a device ch is an integral part of the control rod, and protects a ains the low probability of a control-rod-drop accident. It is esigned to limit the free-fall velocity and consequent activity nsertion rate of a control rod, so that minimal fu damage wo d occur. It is a one-directional device, a that control r scram time is not l significantly affect 1 1 The control rods a e cooled by the core 1 akage (bypass) flow. l The core leakage low is made up of recire ation flow that leaks through several eakage flow paths, the nos important of which are as follows a. The area between the fuel channel and t e fuel assembly lo er tie plate i b. oles in the lower tie plate ( c. The area between the fuel assembly lower tie pla e and l the fuel support piece d. The area between the fuel support piece and the conte 1 ' ( rod gui_de tube 4.1-7 l l

L 1 LGS FSAR ( I e !~ k e_ 'Phe inrea between the control rod guid the core ~ suppo l f. The area he core su ate and the shroud.~} T ow paths are shown in Figure 4.4-3. 4.1.3.3 SuDDiementarY Reactivity Control The initial and reload core control requirements are met by use of the combined effects of the movable control rods, supplementary burnable poison, and the variation of reactor 2he supplementary burnable poison isj;t:f fuel re M coolant flow. f:lir_i_* E'N ^ } riced 1... the er:ni ; fic:id: f;:1 in ee! - v inied.:-I L 41" A duengb 8)

• A9 S8Cbin N 4.1.4 ANALYSIS TECHNIQUES 4.5* k 4.1.4.1 Reactor Internal Components-Computer codes used for the analysis of the internal components are:

a. MASS b. SNAP (MULTISHELL) c. GASP d. NOHEAT e. FINITE f. DYSEA l l g. SHELL 5 h. HEATER i. FAP 71 j., CREEP PLAST i l k. ANSYS Detailed descriptions of these programs are given below: l l ') De chanys on pys 4.1-4 4co$k 4.i-x.,] l 4.1-8 l

( ~ 4.1.4.1.11.3 History of Use t The ANSYS program has been used for productive analyses since 1970. Users now include nuclear, pressure vessel, piping, mining, structures, bridge, chemical, and automotive industries, as well as many consulting firms. ~ 4.1.4.1.11.4 Estent of Application ANSYS is used extensively by GE/NED for elastic and elastic-plastic analysis of the reactor pressure vessel, core support structures, eactor intern SM frel f g 4.1.4.2 Fuel Rod Thermal Analysis newnud in Geck 2.4 of GESTARE (li' f 4,/-lh. s F__uel rod thermal design analyses ares :rferr:d utili:in; t " "_ assical relationships for neat transfer in cylindrical j coor eometry with internal heat generation. Conditions '100% and 116 ed power are analyzed correspondin l steady-state and shor transient operati ormal l operational transients are a In=t_ ensure that the damage limit of 1.0% cla astic is not violated. l The strength the reinology, and strain-s tegories

(

presente e ASME Code, Section III are used as a gu l \\ Lt anical desian and stream analvsis of the fuel rods. 1 l 4.1.4.3 Reactor Systems Dynamics The analysis techniques and computer codes used in reactor systems dynamics are described in Section 4 of Ref 4.1-10. Section 4.4.4.6 also provides a complete stability analysis for the reactor coolant system. 4.1.4.4 Nuclear Encineerino Analysis The analysis techniques are described and referenced in Section 3 ef t.2.2. Se eM:: er M != +he errlyrie = ~: E GastAR.2T(afg./-1) Comnuter Code Function Lattice Phys odel Calculates average few-group cross ections, bundle reactivities, an re fuel rod powers wi he fuel bundle BWR Reactor Simul tar C.u % tes three-dime nal nodal power distributions, exposures, hermal-hydraulic characteristics as a n o Q uenup i h 4.1-17 ,.-,..<.__-,.,,,,.--------,.,-y-,.,.-.,--,,_m__

i LGS FSAR O RA, m.cau 3 ri 4.,1.4.5 Neutron Fluence Calculations I Vessel neutron fluence calculations were carried out using a one-dimensional discrete ordinates Sn transport code with general ~ anisotropic scattering. This code is a modification of a widely used discrete ordinates code which solves a wide variety of radiation transport problems. The program solves both fimod source and multiplication problems. Slab, cylinder, and spherical geometry are allowed with warious boundary conditions. The fluence calculations were done as a fixed source problem in cylindrical geometry. The fixed source was developed from neutron fission distribution data prepared as l a part of'the core physics data. Anisotropic scattering was considered for all regions. The cross sections were prepared l with a 1/E flus weighted and P matrices for anisotropic l scattering, but did not include resonance sele: shielding factors. Fast neutron fluxes at locations other than the core midp&ane were calculated using a two-dimensional discrete ordinate code. The two-dimensional code is an extension of the one-dimensional code. Th.is calculational procedure was also used on Philadelphia Electric Company's Peach Bottom Atomic Power Station Units 2 and 3. These units are similar in design, and are of the same core ) size and power rating as~the Limerick units. The neutron fluz at these units has been measured using a threshold foil technique on the wire samples contained in the vessel surveillance samples. Flux greater than 1 MeV was determined to be 1.3 10' n/cm -see a for Unit 2, and 9.8 x10s n/cm -see for Unit 3. This is a equivalent to a fluence over 40 years of 1.3 x10:e nyt and 9.8 x10" nyt. The calculated value using these methods is 1.4 x1018 nyt. 4 1.4.6 4 desced,. Thermal Hydraulic Calculations * *F the 4h.cnol-hplic woodsts is awen i W computer program uses a parallel flow path model to perform s the s - tate BWR reactor core thermal-hydraulic an 1 Program inpu es the core geometry, o

power, pressure, coolant flow nd iatt

.cna py, and power - distribution within t from the program includes core pressu , coolant flow dis n, critical power rati axial variations of quality, density, halpy for channel type. E 4.

1.5 REFERENCES

(fR7L. Crowther, " Xenon Considerations in Design of 4.1-1 Bolling Water Reactors," APED 5640, June 1968. %W Elecfric.Wl d APf '58fl*Cf'r R888I'# N*le' dddiq O' E d l S Sttdu %plemed, Nf0E-24on-t-A ud NEDE-24cn-A-A-1&(ldsi qpved rwis n). 4.1-18

mmW LGS FSAR i 4.1-2 L. Beitch, "Shell Structures Solved Numerically by Using a Network of Partial Panels," AIAA Journal, Volume 5, No. 3, March 1967. 4.1-3 E. L. Wilson, *A Digital Computer Program For the Finite Element Analysis of Solids With Non-Linear Material Properties," Aeroiet General Technical Memo No. 23, Aerojet General, July.965. 4.1-4. I. Farhoomand and E. L. Wilson, "Non-Linear Heat Transfer Analysis of Axisymmetric Solids," SESM Report SESM71 6, University of California at Berkeley, Berkeley, California, 1971. 4.1-5 J. E. McConnelee, " Finite-Users. Manual", General Electric TIS Report DF 695L206, March 1969. 4.1-6 R. W. Clough and C. P. Johnson, "A Finite Element Approximation For the Analysis of Thin Shells," International Journal Solid Structures, Vol 4, 1968. 4.1-7 "A Computer Program For the Structural Analysis of Arbitrary Three-Dimensional Thin Shells," Report No. GA ( 9952, Gulf General Atomic, January 1970. 4.1-8 A.D. Burgess,# . Guide and Inaineerina Description $ V ,HEATERcComput_er. rdWash,d NWc i 1974.""" ~ L gggy _,o3,,_gy, v..-~,r, ,m,- 4.1-9 L.J. Young, "FAP-71 (Fatigue Analysis Program) Computer Code," GE/NED Desian Analysis Unit R. A. Report No. 49, January 1972. 4.1-10 L.A. Carmichael and G.J. Scatena, " Stability and Dynamic Performance of the General Electric Boiling Water Reactor," APED 5652, April 1969. 4.1-11 Y. R. Rashid, "Th: ry R:;;rt ft: Ort:;-P!rrt Cc ruter A Pr @ .,"GE;; ;0 % 5, ?2C *=e==rch.a n.veier et o R::;rt, J;;;;ry,- ^72.* Knees Mc% ! % c Cgetsol l Compdec' Program," Neoo-2asse, December 1976. l l I 4.1-19 ,w---- -,,-,,,-c,-n -,,,w-new-,-ww w, r-se n,,, w,-,en - n s -v4%----- r r w v'r ~ e m myg-- w-v~~~

l pl pu2GS,- DRAFTa ~ 3 b I 2 FUEL SY m M DESIGN / / ( In rmation covering the following subjects in Section 4.2 is ' el prov ed in topical report NEDO-20944 BWR/4 and BWR/S Desig October 1976. Proprietary information is con ined in NEDE-2 44-P as amended by letter R.H. Buchholz, GE o 0.D. areBWR/4swith251-inchvessels,764fuelassemb/,and2 plants Park, NR, dated February 25, 1980. The Limerick-1 res, loaded on C lattice. Topical report paragraph, table, and igure numbers are consist t with FSAR numbers, except that th initial digit (4) is not p sent. 4.2.1 GENERAL ID DETAILED DESIGN BASES 4.2.1.1 General sian Bases The following addit n is included for N -20944 and NEDE-20944-P, Subsect'on 2.1.1.1.D, Par graph one, line no. 17 from the top of page 2 6: "... are applied for ircaloy-2 adding. A statistical analysis is not perfo ed.for t se design analyses. In lieu of a statistical a proach, the evaluations are per-formed using models whic hav been shown to be conserva-tive when compared to dat ( ference 7). In addition, dimensions used are worst erance conditions, and the calculated results are com a ed to conservative material properties." 4.2.1.2 Detailed Desian Ba es 4.2.2 GENERAL DESIGN DES RIPTION 4.2.2.1 Core Cell l l 4.2.2.2 Fuel Assemb y 4.2.2.3 Fuel Bun 4.2.2.4 Reacti ty Control Assembly l 4.2.3 DESIGN VALUATIONS l 4.2.3.1 R sults of Fuel Rod Thermal-Mechanical E luations 4.2.3.2 Results from Fuel Design Evaluations 4.2.3 3 Reactivity Control Assembly Evaluation (Contrql Rods) 4. .4 TESTING AND INSPECTION .2.4.1 fuel. Hardware. and Assembly \\ O 4.2-1

i Fr nw :t 2 LGS FSAR g ?_ Testino and Insoection (Enrichment and Burna oison\\ d7 1 Cdfreentrations) ) 4.2.4.3 Surveil ection a estino of Irradiated Fuel Rods I 4.2.5 OPERATINn nEVEinPMENT EXPERT.EN .2.6 NCES M k s \\ b G Sl &y s~ l I l l 4.2-2

LGS FSAR DRAFT 4.2 FUEL SYSTEM DESIGN TheformatofthissectioncorrespondstoStandardReviewPla$4.2in NUREG-0800. Most of the information is presented by reference to GESTAR II (Ref. 4.2-1). 4.2.1 DESIGN BASES References to design bases are given in Subsection A.4.2.1 o~f GESTAR II (Ref. 4.2-1). 4.

2.2 DESCRIPTION

AND DESIGN DRAWINGS References to the fuel system description and design drawings are given in Subsection A.4.2.2 of GESTAR II (Ref. 4.2-1). 4.2.2.1 Reactivity Control Assembly (Control Rods) The control rod description is given in Subsection 2.2.4 and is shown on Figures 2.6a, 2.6b, and 2.7 of NEDE 20944-P-1 (Ref. 4.2-2). 4.2.2.2 Reactivity Control Assembly Evaluation The control rod evaluation is given in Subsection 2.3.3 of NEDE 20944-P-1 (Ref. 4.2-2). 4.2.3 DESIGN EVALUATION Compliance with the design bases is discussed in Subsection A.4.2.3 of GESTAR II (Ref. 4.2-1). 4.2.4 TESTING, INSPECTION AND SURVEILLANCE PLANS Descriptions of fuel assembly testing, inspection and surveillance are referenced in Subsection A.4.2.4 of GESTAR II (Ref. 4.2-1). 4.

2.5 REFERENCES

4.2-1. " General Electric Standard Application for Reactor Fuel," including the " United States Supplement," NEDE-24011-P-A and NEDE-24011-P-A-US (latest approved revision). 4.2-2. "BWR/4 and BWR/5 Fuel Design," NEDE-20944-P-1 (Proprietary) and NEDO-20944-1, October 1976, and " Amendment 1," January 1977. 4.2-1 (INSERT) PCY: rf/G05091*-1 5/9/83

DPha4F'1"/ LGS FSAR g T7dLES 4.2-1 to 4.2-Y ~ / - / The tables are contained in Topical Reports NEDO-20 '4 and/or NEDE 0944-P, as amended by letter R.H. Buchholz, GE, o 0.D.

Park, RC, dated February 25, 1980, as Tables 2-1 t 2-12.

Table No Title 4.2-1 Conditions of Design Resulting f m In-Reactor Process Conditions Combined wit Earthquake Loading 4.2-2 tress Intensity Limits 4.2-3 Es mated Number of Cyclep for Each Cyclic Con ' tion Used for Fatigge Analysis 4.2-4 Fuel D a 4.2-5 Material ropertie _ ~' 4.2-6 Post Shipme F 1 Inspection Plan 4.2-7 Inspection E ment 4.2-8 Summary o Exper nce with Production Zircaloy-Clad 00, Fue (Septemb 30, 1974) 4.2-9 Summ y of General E ectric Operating Experience wi Production Gadol ia-Bearing Fuel (September 3, 1974) 4.2-10 General Electric Developm ntal Irradiations Zircaloy-Clad 95% TD 00, P llet Fuel Rods 4.2-11 General Electric Developmenta Irradiations Zircaloy-Clad 95% TD UO, Pelle Capsules General Electric Test Reactor .2-12 Halden Irradiation Program Status pk 4"g,o-zo# .gles*.

~ DRAFT l FIG ES 4.2-1 TO 4.2-11 ARE CONTAINED IN TOPICAL REPORTS NEDO-944 AND/OR NEDE-20944P, OCTOBER, 1976, AS AMENDED BY LETT R.H. BUCHHOLZ, GE, TO O.D PARK, NRC, DATED f FEBRUARY 5, 1980, AS FIGURES 2-1 TO 2-11. / ,/ /* / Figure No. Title / / 4.2-1 C-Lattice Dimensions ,/ 4.2-2 -Lattice Dimensions (Not Ap licable to LGS) 4.2-3 S ematic of Four-Bundle ell Arrangement 4.2-4 Fue Assembl 4.2-5 Fuel As -section n go\\c, h }O 4.2-6a Co ,,49 4.2-6b C ormation Diagram for C-Lattice 4.2-6c Con d Information Diagram for D-Lattice (Not pli ab to LGS) 4.2-7 Control od Vel ity Limiter 4.2-8 Fuel adding Aver e Temperature at a Fuel Col n Axial Gap 4.2-9 C18dding Temperature v sus Heat Flux, eginning of Life 4.2-10 Cladding Temperature versu Heat Flux, End of Life / 4.2-11 j Fuel Energy Release as a Funct on of Time \\, '\\ LIMERICK GENER ATING STATION UNITS 1 AhiD 2 FINAL SAFETY ANALYSIS REPORT SECTION 4.2 FIGURES FIGURES 4.21 TO 4.2-11

~ DRAFT 'es 's*" 4.3 NUCLEAR DESIGN Most of the information of Section 4.3 is provided in the licensing topical report, GESTAR II (Ref. 4.3-1). The subsection numbers in Section 4.3 directly correspond to the subsection numbers of Appendix A of GESTAR II. Any additions or differences are given below for each applicable subsection. 4.3.1 DESIGN BASES e 4.

3.2 DESCRIPTION

4.3.2.1 Nuclear Desian Description The nuclear design description in GESTAR II is referenced in Subsection A.4.3.2.1 of GESTAR II (Ref. 4.3-1) except for the reference (initial) core loading pattern which is shown in Figure 4.3-31. The initial core uses five enrichments of fuel bundles, which are described in Figures 2-2.45, 2-2.46, 2-2.48, 2-2.53, and 2-2.54. 4.3.2.2 Power Distribution i Power distribution is referenced in Subsection A.4.3.2.2 of GESTAR II (Ref. 4.3-1). 4.3.2.3 Reactivity Coefficient 4.3.2.4 Control Requirements 4.3.2.4.1 Shutdown Reactivity Information on shutdown reactivity is referenced in Subsection A.4.3.2.4.1 i of GESTAR II (Ref. 4.3-1), except for the cold shutdown margin for the reference initial-core loading pattern, which is given in Table 4.3-1. 4.3.2.4.2 Reactivity Variations Information on reactivity variations is referenced in Subsection A.4.3.2.4.2 of GESTAR II (Ref. 4.3-1). The combined effects of the individual constituents of reactivity are accounted for in each K,ff in Tame 0 3-1. 4.3.2.5 Control Rod Patterns and Reactivity Worths Control rod patterns and reactivity worths are discussed in Section 3.2.5 of NEDE-20944-P-1 (Ref. 4.3-2). Typical control rod patterns and the l associated power distributions are presented in Appendix A of NEDE-20944-P-1 (Ref. 4.3-2). These control rod patterns are calculated with the BWR l Core Simulator. Qualification for this model is discussed and referenced l in Section 3.1 of GESTAR II (Ref. 4.3-1). I l l 4.3-1 (INSERT) PCY:rf/G05091*-2 5/9/83 l . ~.

LGS FSAR 4.3.2.5.1 Scram Reactivity Scram reactivity is calculated as described in Section S.2 of GESTAR II (Ref. 4.3-1) and is discussed in Section 3.2.5.3 of NEDE-20944-P-1 (Ref. 4.3-2). 4.3.2.6 Criticality of Reactor Durino Refueling 4.3.2.7 Stability 4.3.2.7.1 Xenon Transients 4.3.2.7.2 Thermal Hydraulic Stability l 4.3-2 (INSERT) PCY: rf/G05091*-3 j 5/9/83

~ =3 ; N LGS FSAR ~ NUCLEAR DESIGN 5 Excep or Section 4.3.2.8 which is contained herein, the informat n covering the subjects of Section 4.3 listed b low is provided i_ topical report NEDO-20944, BWR/4 and BWR/5 F el Design, Octo r 1976. Proprietary information is cont ned in NEDE-20944-P a amended by letter R.H. Buchholz, GE, O.D. Park, NRC, date ebruary 25, 1980. Limerick 1 and are BWR/4's with 251 inch ves 1s, 764 fuel assemblies, loaded C lattice. Topical report para aph, table, and figure. number correspond to the FSAR numbers belo except that the initial ' git (4) is not j present. 4.3.1 DESIGN BASES 4.3.1.1 Safety Desian Bases 4.3.1.2 Plant Performance Desio Bases 4.

3.2 DESCRIPTION

4.3.2.1 Nuclear Desian Deser ion 4.3.2.2 Power Distribu 4.3.2.3 Reactivit fficients 4.3.2.4 Cont Requirements 4.3.2.5 trol Rod Patterns and Reactivity Wo ths l l 4.3. .6 Criticality of Reactor Durina Refuelino .3.2.7 Stability ~ 4.3.2.8 Vessel Irradiations The neutron fluxes at the vessel have been calculated using the one-dimensional discrete ordinates transport code described in Section 4.1.4.5. The discrete ordinates code was used in a distributed source mode with cylindrical geometry. The geometry described six regions from the center of the core to a point beyond the vessel. The core region was modeled as a single homogenized cylindrical region. The coolant water region between the fuel channel and the shroud was described as containing saturated water at 5500F and 1050 psi. The material compositions for the stainless steel in the shroud and the carbon steel in the vessel contain the mixtures by weight, as specified in the material specifications for ASME SA 240, 304L, and ASME SA 533 grade B. In the region between the shroud and the vessel (for g conservatism), the presence of the jet pumps was ignored. A 3 4.3-4

t 0AAFT T simple diagram showing the regions, dimensions, and weight fractions is shown in Figure 4.3-29. The distributed source used for this analysis was obtained from the gross radial power description. The distributed source at any point in the co're is the product of the power from the power description and the neutron yield from fission. By using the neutron energy spectrum, the distributed source is obtained for position and energy. The integral over position and energy is normalized to the total number of neutrons in the core region. The core region is defined as a one-centimeter-thick disc with no transverse leakage. The power in this core region is set equ to the maximum power in the axial direction. The radial j'.'b_ I ___,.....'_hbN f " [..distributiong Q hown in Figure 4.3-M os The neutron fluence is determined from the calculated flus by assuming that the plant is operated 90% of th e at 90% power level for 40 years, or equivalent to 1x10' fu wer. seconds. The calculated fluxes and fluence are shown i ble 4.3-5. The calculated neutron flux leaving the cylindrical core is shown in Table 4.3-4. ~ 4.3.3 ANALYTICAL METHODS 4.3.4 CHANGES ) 4.

3.5 REFERENCES

4'3' I " General Electric Standard Application for Reactor Fuel," including " United States Supplement" NEDE-24011-P-A and NEDE-24011-P-A-US (latest approved revision). 4.3-2 "swn/4 and swr /5 Fuel Design," NEDE-20944-P-1 (Proprietary) and 16 0-20944-1, October 1976, and " Amen h nt 1", January 1977. a 4.3-J

TABLES 4.3-1 to 4.3-4 Thhse tables a contained in Topical Reports NE -20944 and/or NEDE-20944-P, as amended by letter RH.Buchholz, E, to 0.D. Park, NRC, dated February 25, 1980 as Tables 3-1 to -4. Table No. Title 4.3-1 Bh eactor re Dimensions v 4.3-2 Reacti t e Cold, Xenon Free State y0 4.3-3 Reactivit rol Fraction for Various Reactor States 4.3-4 Summar 'of BWR/4 and 5 besign Rev @ ns l \\ i N N l 1 l N l l c. l l

LGS FSAR Table 4.3-7 CALCULATED CORE EFFECTIVE MULTIPLICATION AND CONTROL SYSTEM WORTH - NO VOIDS, 20*C Beginning of Cycle-1, K-effective Uncontrolled 1.1047 Fully Controlled 0.9230 Strongest Control Rod Out (18-47) 0.9821 R, Maximum Increase in Cold Core Reactivity with Exposure Cycle-1, ak 0.0 1 \\ l l [ l (INSERT) i l PCY: rf/G05091*-4 5/9/83

FIGURES 4.3-1 TO 4.3-26 ARE CONTAINED IN TOPICAL REPORTS NEDO-20944 AND/OR NEDE-20944-P, AS AMENDED BY LETTER FROM R. H.BUC OL2, GE, TO O.D. PARK, NRC, DATED FEBRUARY 25, 1980. Figure No. Title 4.3,1 I tal Core Loading Map 4.3-2 C-Lat ice Dimensions j 4.3-3 RodType.DesignationsforEnrichmentandGadoling5 Distributi(ons in the High Enrichment, 2.19 wt% U Bundle 4.3-4 Rod Type Designations for Enrichmeurt and Gadolinia UgributionsintheMediumEnrich' ment, D 1.76 wt% Bundle N 4.3-5 Axial Enrichment a'nd Gadolin Distribution, liigh Enrichment Bundle 4.3-6 Axial Enrichment and olinia Distribution, Mediun Enrichment Bundle 4.3-7 K-Infinity a 4$ @ it Exposure at Various Void Fract IP ment, Dominant Fuel Type 4.3-8 Atom Fract ction of Exposure, High i Enrichment, Fuel Type, 40 Percent Voids 4.3-9 Fission ract n as a Function o S Exposure, High Enrich nt, Dominant Fuel Type, 40 Percent Voids g \\ 4.3-10 Delaved Neutron Fraction as a Function of Exposure at, /4 0 percent voids \\ s ,Ieutron Generation as a Function of Exposure at 4.3-11 l 40 Percent Voids s [ Voids - Beginning-of-Cycle 4.3-12 Uncontrolled Local Power Distribution at 40 Percent / / 4.3-13/ Uncontrolled Local Power Distribution at 40 Percent Voide - Beginning-of-Cycle l l 4 -14 Variation of Maximum Local Power as a Function of Exposure for Ifigh Enrichment, 40 Percent Voids, Uncontrolled a LIMERICK GENERATING STATION UNITS 1 AND 2 j FINAL SAFETY ANALYSIS REPORT l SECTION 4.3 FIGURES FIGURES 4.3-1 TO 4.3-26 SHEET 1 OF 2

um Figuro !!o. Title f O 4.3-5 variation of Maximum Local Power as a Function f Exposure for Medium Enrichment, 40 Percent Vofds, Uncontrolled 4.3-16 Uncontrolled Local Power Distribution as Percent Voids - As a Function of Exposure 4.3-17 Uncontrolled Local Power Distribution s a Function of Exposure - Beginning-of-Cycle 4.3-18 ontrolled Local Power Distributio at 40 Percent ids - Beginning-of-Cycle 4.3-19 Un ntrolled R-Facto Distribu ion at 40 Percent Voi - Beginning-cle 4.3-20 Varia on of ag R-Factor as a Function of Bundle A e for the Vacontrolled High Enri c4 g6 4 4.3-21 Radial Pow 6 4.3-22 Beginning-of e and End-of-Cycle Core Sverage Axial Power 4.3-23 Moderator Void activity Coefficient as a Function of Voids at End- -Cycle 1 4.3-24 Doppler Reactivity efficient as a Function of Fuel Expo'sure and Av age Fuel Temperature at an Average, Void Content 40 Percent Iligh Enrichment, Dominant Fuel Type / I 4.3-25 Cold /ShutdownReactivity ? 4.3-26 Ho,t Operating, End-of-Cycle - SCRAM Reactivity (S) l / / / / [ l l l l LIMERICK GEN TING STATION i UNITS ND2 l FINAL SAFETY ANALYSIS REPORT N SECTION 4.3 FIGURES l I FIGUR E S 4.31 TO 4.3-26 SHEET 2 OF 2

l L4.5 F5( DRAFT .EMMMMMM. BB3BBHE8HE8H888BE esBBBBEE8ME8HE8MB8&e eB8BBME8HE8HE8HE8ME8E8e EE8ME8ME8ME8HE8ME8MB8BB iMME8HB3HE8HE8HE8HE8HE6 E88ME8ME8ME8HE8ME8HE8BB RHE8HBBHE8HB3HBsHB3HE8 MBBMBBEE8ME8MBBMBBRE8BB MHB8H88HB8HBsHBBMB8HE Ma8E88M8sM88HE8M88MEBBB

• E888M88HB3HBBHE8MBBEla*

l l "*$888ME8ME8ME8MElBMS" BBBBBHB8HE8HBBEil: SHMMMMME" l l CESTAR II Nomenclature 94 U3 (13 P8C 09 = = 1.63 wt% U235 (160) P8CRB163 C = 2.48 wt% U235 (308) P8CRB248 = 2.78 vt% U235 ( 72) P8CRB278 Figure 4.5-31 Referen e di g Pat rn l

Q L-MakuAa f ) bu o.4 % ycab.S A &Jh &Qf/ h[5 Sdsed,h ' sWuxt =- 5 LGS FSAR $e pQQgdC5*'- 3 kases ec referesced l IMtMMAL AND HYuMAULIC DESIGN O'Y Y'I C 6ESIAN E (Ac f 4.4-/). g.* 4.4.I btFIGNBASIS 4 M d85 b "be me'u 1 *fed go c ' ll 5 ~ s limit for Y um cr*:g 43.1 Safety Desion Bases Tbt+'o (MCI'.4 ed the b ice bed, ) generatsm ete (WGig) uc, g.ve,5 j Ther 1-hydraulic design of the core establish M TokteM-y a. Actuation limits for the devices of the nuclear saf systems so that no fuel damage occurs as a result oderate frequency transient events. Specifica y, the m nimum critical power ratio (MCPR) operating imit is sp cified so that at least 99.9% of the fuel ods in the cor are not expected to experience boiling ransition duri g the most severe moderate frequency tr nsient event. b. The the al-hydraulic safety limits for u. in evaluati the safety margin relating the consequences of fuel b rier failure to public safety c. That the nu lear system exhibits no in rent tendency t'oward diver ent or limit cycle oscill tions that would compromise th integrity of the fuel r nuclear system process barrie 1 4.4.1.2 Power Generatio Desion Bases s The thermal-hydraulic desi of the core ovides the following f operational characteristicst a. The ability to achie rat g core power output throughout the design e of the fuel without sustaining premature 1 failure b. Flexibility t. djust cor output over the range of plant load load maneuv ring requirements in a

stable, dictable manner ithout sustaining fuel l

damage i 4.4.1.3 Requi ements for Steady-State nditions For purpose of maintaining adequate therm margin during normal. steady-stat'e operation, the MCPR must not b less than the required PR operating limit, and the maximu linear heat t j generatio rate (MLHGR) must be maintained belo the design linear h t generation rate (LHGR) for the plant. This does not specify he operating power, nor does it specify p king factors. rameters are determined subject to a number including the thermal limits given previo(0 These The constraints, core /and fuel design basis for steady-state operation, gly. i ;e4, MCPR LHGR limits, have been defined to provide a margin between p l 4.4-1 1

LGS FSAR the steady-state operating conditions and any fuel damage / ondition to accommodate uncertainties and to ensure that/no fue1\\ l age results even during the worst anticipated transi t i { opercon ition at any' time in life. The design steady-stat MCPR ing limit and the MLHGR are given in Table 4.4 g 4.4.1.4 eauirements for Transient Conditions i The transie thermal limits are established so t t no fuel damage is ex cted to occur during the most seve moderate frequency'tran ent event. Fuel damage is defi dasperforation[ of the cladding hat permits release of fissio products. The following mechani s can cause fuel damage in eactor transients: l caused by inadequate,l a. Severe ove eating of fue) claddin cooling i b. Fracture of th fuel cladding c used by relative i expansion of th uranium diori e (UO,) pellet iwside the' fuel cladding For design purposes, the tra lent I'mit requirement is met if at. least 99.9% of the fuel rods the core do not experience boiling transition during any a d ate frequency transient event. No fuel damage would be expected o occur even if a fuel rod actually experienced a boiling sition. l A value of 1% plastic strain f Zir loy cladding is j conservatively defined as th limit b low which fuel damage from I overstraining the fuel cla ing is not expected to occur. The l linear heat generation rat,4 required to cause this amount of cladding strain is appro 4mately 25 kW/f in unirradiated UO, j fuel, but it decreases ith burnup to app ximately 20 kW/ft fori UO, at a local exposur of 40,000 mwd /t. 4.4.1.5 Summary of esion Bases In summary, the eady-state operating limits ha e been established to most severe mo erate fr4qu'ncy transient event.sure th6 the design basis is sa {htre is sfied for the T steady-state esign M pc wr basis. An. overpower t t occurs during an i ident *I e ; prate frequency transient ent must meet the p nt trantaent'6ePR limit. Demonstration th t the transient imits are not exceeded is sufficient to conc! de that the desi n basis is satisfied. l The M R and MLHGR limits are sufficiently general so thatNno I othe limits need to be stated. For example, cladding surface j te eratures are always maintained within 50 to 150F of the \\ ~. lant temperature as long as the boiling process is in the \\ \\ c ucleate regime. The cladding and fuel bundle integrity \\ ' erion is ensured as long as the MCPR and MLHGR limits are \\ - ~ 4.4-2

@ w =" t w [ Ii ' 4 OA a LGS FSAR C/ re are no additional design criteria on codlant fraction, co ant flow-velocities, or flow eitfG'Eion, nor. are they needed. The e low mieet es and void fraction l become constraints on +-;;;;ctEn ca nd W ales _ design of J reactor co and are partially constrained by l ibility g co requirements. r -4.

4.2 DESCRIPTION

OF THERMAL-HYDRAULIC DESIGN OF THE REACTOR (e,dey%, fe rut.lv corths e aim s %e mnk su%.4.4.2 A dese,. in SdsesL A ,e r,es n s be:a. 9 An eva us on o per ormance rom a ' thermal affd hydraulic standpoint is provided in Section 4.4.3. 4.4.2.1 Summary Comoarison A summary comparison of the thermal and hydraulic design parameters of the reactor with reactors of.pimilar design is provided in Table 4.4-1. = 4.4.2.2 Critical Power Ra'tio [ earethreedifferenttypesofboilingheattransfertowath in a reed convection system: nucleate boiling, transition

boiling, nd film boiling.

Nucleate boiling, at lower heat,/ ( transfer r s, is an extremely efficient mode of heat d ansfer, allowing large uantities of heat to be transferred With a very small temperature se at the heated wall. As the heat transfer rate is increased, t boiling heat transfe (surface alternates between film and nuclea boiling, lead 16g to fluctuations in heated wall temperatures, he d of departure from the nucleate boiling region inyt e transition boiling region is called the boiling trajn irion. nsition boiling begins at the critical power and.is characterize fluctuations in cladding surface tempera.tufe. Film boiling occu at the highest heat transfer rates; it begins as transition bol comes to an end. ' Film boLlitg heat transfer is characterized by le wall j 4 y eate boiling. temper 1iitures that are higher than those experienced diiMG 2 I. 2.2.1 4 Boiling Correlations The occurr.ence of boiling transition is a fuyct on of the local steam quality, boiling length, mass flow ce, pressure, flow geometry, and local peaking pattern E has conducted extensive experimental investigations of se parameters. These parametric studies encompas e entire design range of these variables. In the experd ental investiga'tions, a boiling transition event w s 4ssociated w a 250F rise in rod surface temperature. T (critical) quality a't Qich boiling transition occurs as a netion of the distance from the 4quilibrium boiling g boundar s predicted by the GEXL (General Electric. Critical Qua y, Xe - boiling length) correlation. This cor. relation i 4.4-3

fT Y N., sed on accurate test data from full-scale prototype simula ns o eactor fuel assemblies operating under conditions du cating thos actual reactor designs. The GEXL correlat s a "best fit" to ata and is used together with a a tical analysis to ensure ad e reactor thermal margins 4.4-1). The figure of merit us e reac esign and operation is the critical power ratio (CPR). is defined as the ratio of the bundle power that would p ce rium quality equal to, but not exceeding, the co ation value (c al quality) to the bundle power at t eactor condition of int (i.e., the ratio of crit bundle power to operating bundle wer). In this defi on, the critical power is determined at same mass , inlet temperature, and pressure that exist at ( tfied reactor condition. 4.4.2.3 Linear Heat Generation Rate (LHGR) I limiting constraints in the design of the reactor core e st ed in terms of the LNGR limit and the MCPR operating 1 it for e plant. The design philosophy used to ensure that hese limits I re met involves the selection of one or more pow distribu ons that are more limiting than expected ope ting conditions nd subsequent verification that, under th e more stringent co itions,.the design limits are met. T refore, the T " design power stribution" is an extreme conditi of power. It / is a fair and st ngent test of the operability the reactor as designed to comply ith the foregoing limits. spected operating conditions are less evere than those repres ted by a design power distribution, w ch gives the maximu allowable LHGR and the MCPR operating limi for the plant. owever, it must be established that operatio with a less evere power distribution is not a necessary conditi for the fety of the reactor. Because there is an infinite umber f operating reactor states that can exist (with variatio i rod patterns, time in cycle, power level, distribution, flow etc) that are witid n the design constraints, it is not possib t determine them all.

However, constant monitoring of oper ing co itions using the available plant measurements can en re compli ce with design objectives.

The core-average ~and are given i Table 4.4-1. 4.4.2.3.1 Design Po r Distribution The thermal desi of the reactor--including t selection of the core size and fective heat transfer area, the sign steam quality, the otal recirculation flow, the inlet a cooling, and the specif ation of internal flow distribution--is ased on the concept d application of a design power distributio The desig wer distribution is an appropriately conservat e repr entation of the most limiting thermal operating sta at r ed conditions and includes design allowances for the com ned ) fects (on the fuel rod and on the fuel assembly heat flux a 4.4-4

__n - MAP 1 LGS FSAR ( mperature) of the gross and local steady-state powie d ity di ributions and adjustments of the control rods. The des power distribution is used in conjun on with flow and pressur drop distribution computations determine the i thermal condit s of the fuel and the en Ipy conditions of the coolant throughou he core. The desi axial power distribution; -used in the calculet of the MCPR eration limit is given in Table 4.4-2. This distr tion i consistent with that discussed' in Ref 4.4-1. The design power distrib$ on is ba d on detailed calculations of the neutron flux disfribution as di ssed in Ref 4.4-30. 4.4.2.3.2 Desi Linear Heat Generation Rat 6s N s s f The maximum and core average linear heat generation'tm es are shown ip-fable 4.4-1. The MLHGR at any location is the verage line heat generation rate at that location times the tot)1 }{ ing factor. 4.4.2.4 Void Fraction Distribution / The cork-average and maximum e'xit void fractions in the core at (, rated condition are given in Table 4.4-1. The axial distribution of core void fractions for the average radial channel and the maximum radial channel (end of node value) for the core are given in Table 4.4-3. The core average and maximum exit values are also provided. Similar distributions for steam quality are provided in Table 4.4-4. The core average axial power distribution used to produce these tables is given in Table 4.4-5. 4.4.2.5 Core Coolant Flow Distribution and Orificing Pattern C rrect distribution of core coolant flow among the Euel T t as blies is accomplished by the use of an accurate alibrated; fixe ifice at the inlet of each fuel assembly .se orifices i are 1cca in the fuel support piece. They ntrol the flow distribution hence the coolant condi ns within prescribed bounds throughout e design range o ore operation. The sizing' and design of the ori s en table flow in each fuel assembly during all pha operation at normal operating conditions. i The core is vided into two orificed flow es. The outer zone is a narr reduced-power region around the pe ' ery of the f core. e inner zone consists of the core center re ~ n. No othe control of flow and stream distribution, other thahsthat I in dentally supplied by adjusting the power distribution with., ( e control rods, is used or needed. The orifices can be chang R' uring refueling, if necessary. r _1 j

o 4.4-5

,o -3 w-w---.. -y-- ,,.,..ww,w-,,,-, ---,..,-,-..-,....,v. .m-m -.w.-wnwm,-.- .,,-we-w.,v--- e.-e-

LGS FSAR [ gn core flow distribution calculations are made using, desi er' distribution, which consists of a hot and a rage powered embly in each of the two orifice zones. T design bundle power d resulting relative flow distributi are given in Table 4.4-6. i The flow distribution the fuel assembli is calculated on the assumption that the press drop acros all fuel assemblies is the same; This assumption h been firmed by measuring the flow distribution in a modern g water reactor as reported l in Ref 4.4-2. i There is reasonable assur ce, therefore, hat the calculated flow distribution thro 2 out the core is in ose agreement with the actual flow dis bution of an operating r tor. l The use of th esign power distribution discussed viously ensures th the orificing chosen covers the range of rmal operati The expected shifts in power production duri core life e less severe and are bounded by the design power d ribution. 4.4.2.6 Core Pressure Drop and Hydraulic Loads i i e pressure drop across various core components under the stead T st design conditions is included in Table 4.4-1. Analyses r\\ i 1 the a t limiting conditions, the recirculation line break I the ste line break, are reported in Chapter 3. The componen of bundle pressure drop considered ar

friction, local, elevati

, and acceleration. Core plate pr ssure drop measurements hav een taken on several operat g BWR/3 and 4 plants containing 7, 8x8, and mixtures of x7 and 8x8 fuel. Table 4.4-7 compares asured and calcul ed core plate pressure drops. The measured an calculated v ues are in good agreement. i l The data are predicted wi an ave ge error of 0.04 psi. The l one sigma error is 0.86 psi. I ( The thermal hydraulic loa on t fuel rods during the steady-i state operation, trans t, and acc ent conditions are l negligible, primari ecause of the nnel conf.inement, thereby resulting in sma crossflow between rod i.e., essential constant pres e at any given elevation in e fuel bundle). The loa (i.e., horizontal) across the control b es are mini or negligible, primarily due to the flat inte annel (vefe'cityprofileasgiveninRef4.4-3. / 4.4-6

u LGS FS / (' y__ ~~ 4. .2.6.1 Friction Pressure Drop Fric ion pressure drop is calculated using the model relation: AP = wa gL ,a D A6H TPF / f 29cs H where: / / APf friction pressure drop, psi / = mass flow rate { .W = i g avitational constant t ge = i / water density s = i chan\\ / 1 DH nel hydraulic diameter = ACH channel low area = length / L = friction factor / f = {. ,aTPF" two-phase fr ion multiplier I This basic model is similar to $ hat used throughout the nuclear power industry. The formulation for the two-phase multiplier is l based on data that compare closely to that found in the open literature (Ref 4.4-4)/ GE has taken signific/ant amounts of iction pressure drop data in multirod geometries representative yf modern BWR plant fuel bundles and correlated both the friction factor and two-phase multipliers on a dest-fit basis using the above pressure drop formulation. Checks against more recent data are being made on a continuing basis to ensure that the best models are used over the full range of Anterest to BWRs. 4.4.2.6.2 ocal Pressure Drop \\ { The local, pressure drop is defined as the irrever'sible pressure loss associated with an area change such as the orifice, lower tie pla 'es, and spacers of a fuel assembly. \\ The g neral local pressure drop model is similar to the. friction pre sure drop and is expressed as: AP = wa K ,a L 2ges Aa TPL here: C'_ 4.4-7

M FSAR ) g . local pressure drop, psi 1 AP = K local pressure drop loss coefficient = A reference area for local loss coefficien e$.pg = two phase local multiplier W, ge and p are defined in the same way as for fr ction. This basic model is s ilar to that used throughout e nuclear power industry. The fo ulation for the two-phase a tiplier is similar to that re ted in the open literat e (Ref 4.4-5) with the addition of empi cal constants to adju the results to fit data taken by GE for t e specific designs f the BWR fuel assembly. Tests are pe ormed in single-hase water to calibrate the orifice in the lower ie plate and eformed in both single-and two-phase flow to arri at best-design values for spacer and upper tie plate pressur drop. e range of test vae6ebles is specified to include the r ge o interest to BWRs. To ensure that the most applicable meth a in use at all times, new data are taken whenever there i significant design change. 4.4.2.6.3 Elevation Pressure op The elevation pressure drop based the well-known relation: } / 72 A1. = 9c sf (1- ) + sgo a AL = gc where: A% el vation pressure drop, psi = AL neremental length = T average water density = average void fraction over the length AL e = sf,sg , saturated water and vapor density, respecti ely acceleration of gravity g = gravitational constant g = 4.4-8

i l LGS FSAR ( 4. .2.6.4 Acceleration Pressure Drop / A re rsible pressure change occurs when an area change / s i l encoun t f accelera ed through the boiling process. red, and an irreversible loss occu The basic crmulatioh i for the r ersible pressure change resulting from a low area change is g'ven by: ~ , APACC " (1 ~ 8') W" 1 8"h 2gepAj As where: APACC = accele tion pressure drop l A, final f1 area / = i / l } A initial flo grea = j k Other terms are as previously\\ efirIed. The basic fo N ation for the acceleration pressure changa due to density change is: /\\ l y x AP = wa 1 IN 9cAs ,N ( l in \\, out AM / '\\ i where / i 1 = X + (1_x): fp op (1-s)s 'n M g f d g,f momentum density p s X steam quality = Other terms are as previously defined. The total acceleration z pressure drop in BWRs is on the order of a few percent of the, j Q otal pressure drop. v 4.4.2.7 Correlation and Physical Data GE has obtained substantial amounts of physical data in suppcet of the pressure drop and thermal hydraulic loads. discussed W u. Sect 4en-4.6 2.44-Corre l at ions-Aa ve -been -devel to 4it--these, date-40-the-formul ions -discussed? Tbs iJom ion n Sin n86 4fFAM E. of GiESTAit X (Ref 4.4-l whe.rg, respeco m. pre &ded to NRC peshoe on Sedw. 4 n t.e.2.7." h ..u6= Lw Ca M tier.; _ WAC - GE Falr-takest nificant amoun ..of-feictierrpressure droi~dVN ' i ( in multirod e esentative of modern BWR plant fuel f pd1 correlated both the frTcTi'6frfactor-and 4wo-phas4./ n 4.4-9 .m ._-_-.__.._._--...---...._,._m_,,,____,,.,___,_.,y _m..,_ ._-,_,,c, _,7.,..y,_,___m,, y,

f W FSAR i f Itipliers on a best-fit basis using the pressure drop fo ulations~ reported ~in Sections 4.4.2.6.1 and 4.4.2.6.2." Che against more recent data are being made on a continuo i basis o ensure that the best models are used over the ful range of into t to BWRs. Tests are formed in single-phase water to calibrat the orifice and t lower tie plate and performed in bo single-and two-phase flow arrive at best-fit design value for spacer and upper' tie plate essure drop. The range of te variables is specified to inclu the range of interest t WRs. To ensure that the most applic e methods are in us t all times, new data are taken wheneve here is a signi cant design change. Applicability to the singl hase a two-phase hydraulic models discussed in Sections 4.4.2. I a 4.4.2.6.2 is confirmed by prototype (64-rod bundle) flow sts. The typical range of the test data is summarized in Ta e 4-8. 4.4.2.7.2 Void Fraction rrelation The void fraction cor lation used is a v ion of the Zuber-Findlay model (Ref .4-10) in which the conc tration parameter and void drift co f.icient are based on compar on with a large quantity of wor wide data (Refs 4.4-11 through 4-22). ) 4.4.2.7.3 eat Transfer Correlation The J s-Lottes wall superheat equation (Ref 4.4-6) is u in fu design to determine the cladding-to-coolant heat tran er efficients for nucleate boil _ing. 4.4.2.8 Thermal Effects of Operational Transients rThe evaluation of the core's capability to withstand the thermal . effects resulting from anticipated operational transients is b vered in Chapter 15. A l 4.4.2.9 Uncertainties in Estimates r certainties in thermal-hydraulic parameters are consid the atistical analysis that is performed to the fuel claddi ntegrity safety limit so that est 99.9% of the fuel rods he core are not ex to experience boiling i transition durt modera requency transient event. The statistical model an rocedure are described in detail in Ref 4.4 The conserva wer distribution used for the sta cal analysis is shown in F g in terms of a rel bundle power histogram. The uncertainties co ed heir input values for the analysis are shown in Table 4.4 l . ) ~No clwi cs mpays 4.4-ll ugb Y'Y~ IY' 4.4-10

maccOn [h LGS FSAR Design of the single-cycle BWR plant includes the abiIity to follow load demand over a reasonable range without requiring i operator action. This load-following capability is accomplished by automatic variation of reactor recirculation flow. The plant i responds to ramp load changes at a rate of up to 30% per minute without changes in control rod settings. The reactor power level can be controlled automatically by flow control over approximately 35% of the power level on the rated rod line. Load following is accomplished by varying the recirculation flow to the reactor. This method of power level control takes advantage of the reactor negative void coefficient. To increase reactor power, it is necessary to increase the recirculation flow rate that sweeps some of the voids from the moderator, causing an increase in core reactivity. As the reactor power increases, more steam is formed and the reactor stabilizes at a new power level with the transient excess reactivity balanced by the new void formation. No control rods are moved to accomplish this power level change. Conversely, when a power reduction is reovired, it is necessary only to reduce the recirculation flow rate. When this is done, more voids in the moderator automatically decrease the reactor power level to a level commensurate with the new recirculation flow rate. Again, no control rods are moved to accomplish the power (' reduction. Varying the recirculation flow rate (flow control) is more advantageous, relative to load following, than using control rod positioning. Flow variations perturb the reactor uniformly in the horizontal planes and ensure a flatter power distribution and reduced transient allowances. As flow is varied, the power and void distributions remain approximately constant at the steady-state end points for a wide range of flow variations. After adjusting the power distribution by positioning the control rods at a reduced power and flow, the operator can then bring the reactor to rated conditions by increasing flow, with the j assurance that the power distribution will remain approximately constant. Section 7.7 describes how recirculation flow is varied. 4.4.3.6 Thermal and Hydraulic Characteristics Summary Table The thermal-hydraulic characteristics are provided in Table 4.4-1 for the core and in tables of Sections 5.1 and 5.4 for other portions of the reactor coolant system. s' S see A. d (= ESTAR E (Ref 4A-N.%e resufh oNe, cple-! ,I 94i design ma is employ the thermal-hydraulic?= s ; cha acteri'sti quipmen( chara teristids, nuc ea,r,e'co'njunhtion T incorpor. ed the core'desi n in (- wit the/ plant tnst _ 'entation, d 4he reacto ot'ection syhte_m_,_isJ o r o\\ / i SWAldt a.%sts ut gwen in Table //.V-l1 a4Qwss 4.4-7 hu.sh 4 4-/0. 4.V.5 75571Ar MD vresicAT/04/ 4.4-15 a See. Sa ue.+a A.4.' ~ ~L GesTAC X (Ref <l.4-1). f

& Ai sadh 4.4.4 (4mf py$&Q f~f mb9 9 LGS FSAR h1stnofuel-damageoccurduringnormaloperationordurihg '.. ) a ormal operational transients. Demonstration that the app cable thermal-hydraulic limits are not exceeded is giv n by anal es. l 4.4.4.1 Critical Power The GEXL c tical power' correlation is used in thermal ydraulic evaluations. This correlation is discussed in more d all in Section 4.4.2. .1. l 4.4.4.2 Core Hy aulics Core hydraulic mode and correlations are discu ed in Sections 4.4.2.6, 4.4.2.7, an 4.4.4.5. 4.4.4.3 Influence of P er Distributions The influence of power dis ibutions on the hermal-hydraulic design is discussed in Ref 4-1, Appendiz 4.4.4.4 Core Thermal Response The thermal response of the core for cidents and expected transient conditions is discussed in hapter 15. } 4.4.4.5 Analytical Methods The analytical methods, thermod namic ata, and hydrodynamic data used in determining the thers and by aulic characteristics of the core are similar to thos used throu hout the nuclear power industry. 1 Core thermal-hydraulic a lyses are perform d with the aid of a digital computer progr This program mode s the reactor core through a hydraulic d cription of orifices, ower tie plates, fuel rods, fuel rod acers, upper tie plates, fuel channel, and the core bypass flo paths. I 4.4.4.5.1' React Model The orifice, 1 er tie plate, fuel rod spacers, and pper tie plate are hyd aulically represented as being separate, distinct lo:a1 losse of zero thickness. The fuel channel cross-section is represe ed by a square section with an enclosed area equal tc-the unrod ed cross-sectional area of the actual fuel channel. The fue channel assembly consists of three basic 'arial regions. The fi t and most important is the active fuel region, wh4ch ts of the 62 fuel rods, 2 non-fueled rods, and 7 fueI\\ rod cons cpa ers. The second is the non-fueled region consisting of 44 -) -fueled rods and the upper tie plate. The third region \\ n sents the unrodded portion of the fuel channel above the\\ 4.4-16

~ LGS FSAR [ up r tie plate. The active fuel region is consider.ed in 2 ind endent axial segments or nodes over which fuel thermal ties are assumed to be constant and coolant proper pro es are assum to vary linearly. The code an handle 12 fuel channel types and 10 type of bypass flow path In normal analyses, the fuel assemblies re modeled by four cha nel types--a " hot" central orifice regi n channel type, an ave ge central orifice region channel t e, a " hot" peripheral ort ice region type, and an average p ipheral orifice l region type. ually, there is one fuel assemb each of the " hot types. representing balance of the co The average types th n make up the The computer program iterates on flow thro h each flow path (fuel assemblies and pass paths) until e total differential t pressure (plenum to pl um) across each ath is equal, and the sum of the flows throug each path equa s the total core flow. Orificing is selected to o imize the core flow distribution between orifice regions as 'scusse in Section 4.4.2.5. The core design pressure is deter ined rom the required turbine throttle pressure, the steam 1 e ressure drop, steam dryer pressure drop, and the steam se ator pressure drop. The core inlet enthalpy is determined fr the reactor and turbine heat a balances. The required core f w then determined by applying the procedures of this sectio and ecifications so that the thermal limits of Ref 4.4-1 e sati ied and the nominal expected bypass flow fracti is appro imately 10%. The results of applying these methods nd specifica ions are as follows: a. Flow for each ndle type b. Flow for eac bypass path c. Core pres re drop d. Fluid p perty axial distribution for e ch bundle type i e. CPR Iculations for each bundle type 4.4.4.5.2 stem Flow Balances The basi assumption used by the code in performing th hydraulic analysi is that the flow entering the core divides itssJf s betwe the fuel bundles and the bypass flow paths so thA ,ly and bypass flow path experience the same pressur(y each ass e drop. Th bypass flow paths considered are described in Table 4.4%12 d shown in Figure 4.4-3. Due to the large flow area, the ressure drop in the bypass region above the core plate is essentially all elevation head. Thus, the sum of the core plate i e o 4.4-17

r.. r a-pa LGS FSAR /. dif rential pressure and the bypass region elevation hesd is equa to the core differential pressure. The tot core flow less the control rod cooling flow enters he lower ple m through the jet pumps. A fraction of this pas s through th various bypass paths. The remainder passes th ough the orifice the fuel support (experiencing a pressure oss) where more f1 is lost through the fit-up between the el support and the lower tie plate and also through the 1 wer tie plate holes into he bypass region. Most of the flo continues through the lower ie plate (experiencing a pressur loss), where some flow is lost t rough the flow path defined by he fuel channel and lower ti plate, and restricted by t finger springs, into the byp s region. The flow through the byp ss flow paths is exp essed by the form: W =Ci a P1 2 + AP * +C P2"- 3 Full-scale tests have been pe formed to stablish the flow ~~ coefficients for the major flo paths ef 4.4-23). These tests simulate actual plant configura 'ons hat have several parallel flow paths, and therefore the fl efficients for the individual paths could not be sepa ated. However, analytic'al models of the individual flow pat s were developed as an independent check of the tests. The odels were derived for actual BWR design dimensions a const ered the effects of dimensional variations. Thes models edicted the test results when the "as-built" dimensio s were appl'ed. When using these models for hydraulic desig calculations, nominal drawing dimensions are used. Thi is done to yiel the most accurate prediction of the expect d bypass flow. Be use of the large number of components i a typical BWR core, viations from the nominal dimensions wi tend to statistically ncel, resulting in a total bypass fl best represented by that alculated using nominal dimensions. The balance of t flow enters the fuel bundle fro.. the lower tie plate and passe through the fuel rod channel spaces' A small portion of the in-channel flow enters the non-fueled ods through three orific holes in each rod just above the lower t plate. This flow, rmally referred to as the water rod flow, mixes with the a ive coolant channel flow below the upper tie late. The uncer ainties in calculations and the resultant uncer inty in reac r coolant system flow rate are provided in Table 4-9. 4.4.4 .3 System Heat Balances Wi in the fuel assembly, heat balances on the active coolant arg p formed nodally. Fluid properties, expressed as the bundle verage at the particular node of interest, are based on 4.4-18

LGS FSAR f 4.4-7. In evaluating fluid properties, a constant p essure mo el is used. The re power is divided into two parts: an active colant power nd a bypass flow power. The bypass flow is b Eted by neutron lowing-down and gamma heating in the water, and by heat transfer brough the channel walls. Heat is also ransferred to the bypass low from structures and control elem ts that are themselves ated by gamma absorption and by (n a) reaction in the control terial. The fraction of total r ctor power deposited in t e bypass region is very nearly 25.. A similar phenomenon occu , with the fuel bundle, to he active coolant and the water ro flows. The net effect is that 96% of the core power is conducte through the fuel claddi g and appears as heat flux. In design analyses, t power is alloc ed to the individual fuel bundles using a relati power factor. The power distribution along the length of the uel bundle specified with. axial power factors that distribute t e bundle' power among the 24 axial nodes. A nodal local peak g fact is used to establish the peak heat flux at each noda loca ion. The relative (radial) and axi ower distributions, when used with the bundle flow, determin the axial coolant property distribution, resulting in su fi 'ent information to calculate the pressure drop components with ach fuel assembly type. Once the equal pressure drop cri erion h s been satisfied, the critical bundle power (the power tha would result in critical quality existing at some oint in the undle using the correlation expressed i Ref 4.4-8) is etermined by an iterative process for each fuel pe. In applying the abov methods to core desi the number of bundles (for a spec' ied core thermal power) and bundle geometry (8x8, rod diameter etc) is selected based on power density and LHGR limits. 4.4.4.6 Therma -Hydraulic Stability Analysis 4.4.4.6.1 I roduction There are ny definitions of stability, but for fee ack processes nd control systems, it can be defined as fo lows: a system i stable if, following a disturbance, the trans ent settles o a steady, noneyelic state. A sy em may also be acceptably safe even if oscillatory, ded that any limit cycle of the oscillations is less thgn a pro cribed magnitude. pr Instability, then, is either a continuaJ parture from.a final steady-state value or a greater-than-N rescribed limit cycle about the final steady-state value. 4.4-19

l = m;"M & brx W l' LGS FSAR The mechanism for instability can be explained in terms of f freq ncy response. Consider a sinusoidal input to a feedback conte system that, for the moment, has the feedback ( disconn eted. If there were no time lags or delays between input; and outp the output would be in phase with the input. / i Connnectin the_ output so as to subtract from the input (negative feedback or 800 out-of-phase connection) would result in stable l closed loop o eration. However, natural laws can cause phase' shift between tput and input and, if the phase shift reacKes 1800, the feedba signal would be reinforcing the input sfgnal l rather than subtr ting from it. If the feedback signal pere equal to or larger han the input signal (loop gain equal to one or greater), the inp signal could be disconnected and the system would continue o oscillate. If the feedback ignal were l less than the input sig 1 (loop gains less than one, the oscillations would die ou It is possible for an unstab e process to be st ilized by adding f a control system. In general, however, it is P eferable that a l process with inherent feedback e designed to se stable by itself j before it is combined with other rocesses d control systems. The design of the BWR is based on is pre se, namely, that l l individual system components are st le. l 4.4.4.6.2 Description Three types of stability considered 'n th design of BWR's are: reactor core (reactivity) stabilit, chann hydrodynamic stability; and total system stabi ty. Rea ivity feedback l l instability of the reactor core ould drive t reactor into l power oscillations. Hydrodynar c channel insta ility could impede heat transfer to the m erator and drive e reactor into power oscillations. The tot system stability co iders control system dynamics combined w' h basic process dynamics 3 A stable system is analytically de.onstrated if no inherent li t cycle or divergent oscillation d elops within the system as a result of calculated step distur ances of any critical variable, such as steam flow, pressure neutron flux, and recirculation flow. The criteria to b considered are stated in terms of two compatible para ters. First is the decay ratio X,/Xo, designated as e ratio of the magnitude of the second overshoot N to the first vershoot resulting from a step perturbation. A plot of the ecay ratio is a graphic representation of the physical r sponsiveness of the system, which is readily evaluated in a tim -domain analysis. Second is the damping coefficient es n the de nition of which corresponds to the pole pair closest to the j axis in the s-plane for the system closed loop transfer fun ton. This parameter also applies to the frequency-domain in rpretation. The damping coefficient is related to the decay tio as shown in Figure 4.4-4. ) 4.4-20 )

LGS FSAR 4 4.4.6.3 Stability Criteria t The ssurance that the total plant is stable and therefore has a signt icant safety margin is demonstrated analytically when the decay atio, X,/Xo, is less than 1.0 or, equivalently, when the damping coefficient e, is greater than zero for each type of n stabilit discussed. Special attention is given to different' te between i erent system limit cycles and small, acceptable l'.it cycles tha are always present, even in the most stable rea tors. The latter e caused by physical nonlinearities (deadband striction, e ) in real control systems and are not repr enta-tive of inhere t hydrodynamic or reactivity instabiliti in the reactor. The u timate performance limit criteria for e three types of dynamic erformance are summarized below in erms of decay ratio and d ping coefficient: Channel hydrodyn ic stability X,/X < 1, en>0 Reactor core (react'vity) stability X,/ o < 1, en>0 Total system stabilit ,/Xo < 1, en>0 These criteria are satisfie for all atta' able conditions of the reactor that may be encounte d in the c utse of plant operation. For stability purposes, the m t severe condition to which these criteria are applied correspon to t highest attainable rod m l line intersection with natural 'rcu tion flow. 4.4.4.6.4 Mathematical Model The mathematical model represen ng e core examines the linearized reactivity response f a re ctor system with density-dependent reactivity feedbac caused by boiling. In addition, the hydrodynamics of variou hydraulical coupled reactor channels, or regions, are xamined separa ly on an axially multinoded basis by grou ng various chann s that are thermodynamically and h draulically similar. This interchannel hydrodynamic interact

• n, or coupling, exist through pressure i

variations in the in et plenum, such as can be caused by disturbances in th flow distribution between r gions or channels. This a roach provides a reasonably a urate, three-dimensional repr sentation of the reactor's hydro namics. The core mode (Refs 4.4-24 through 4.4-29), shown i block diagram for in Figure 4.4-5, solves the dynamic equa 'ons that represent e reactor core in the frequency domain. Fr m the solution f these dynamic equations the reactivity and i ividual channel ydrodynamic stability of the BWR is determined fo a given eactor flow rate, power distribution, and total powe s This ives the most basic understanding of the inherent core'N O be vior (and hence the system behavior) and is the principal c sideration in evaluating the stable performance of the 4.4-21 -~__,._.,_---,-,___-_-.m_.m. ,m., S.

FYP f LGS FSAR 2a re ctor. As new experimental or reactor operating da are obt ined, the model is refined to improve its capabi; /'and accu cy. The pla t model considers the entire reactor system, eutronics, j heat tra fer, hydraulics, and the basic processes, a.well as associate control systems such as the flow controlle pressure regulator, eedwater controller, etc. Although the c:1 trol systems may e stable when analyzed individually, fin control system settin must be made in conjunction with the :perating reactor so tha the entire system is stable. The pla.: model yields results at are essentially equivalent to thc..e achi ed with the core mo 1 and allows the addition of the co.;rol rs, which have adjusta le features permitting the attaintreat the desired performance. The plant model solve the dynamic equations that pr ent the BWR system in the time doma' The variables, such as tam flow and l pressure, are represente as a function of time. he I extensiveness of this mode (Ref 4.4-9) is show in b ack diagram form in Figure 4.4-5. Many f the blocks are xtensi.+ systems j in themselves. The model is eriodically re ned, as 'ew experimental or reactor operat'ng data are tained, :J 1mprove its capability and accuracy. 4.4.4.6.5 Analytical Confirmatio l Figure 4.4-5 demonstrates the compet ce and inherent l conservatism of the core stability o 1. The relati..nship of, the calculated damping coefficien from the reactor c..re dynamic l analytical code is related to me sured r ults from ;4 control rod oscillator tests performed t large op ating BWF plants by GE. The correlated Most Prob le Values (b ed on a. east squares analysis) and the li e representing a 97.5% two sigma) confidence level, below whi h the actual value will idll, are presented in Figure 4.4-6 The results show the a lytical methods to be an e fc.tive and useful design tool, w' h significant conservatism i its application to boili g water reactor core evaluation. Neal and Zivi (Ref 4.4-7) f ther confirm the effective applica lon of essentially the s.e model to channel and core analyr.is. 4.4.4.6.6 An ysis Results The most li iting condition occurs near end-of-cycle, with po er peaking t ard the bottom of the core. Because of tbc decreas, in delay d neutron fraction, thevalueofthedensityreactivity\\r coeffi

ent, AK/K/s ff As, increases.

The most sensi' 1ve reacto / e oper ing condition is that corresponding to natural irculation 3 l flo and a power level corresponding to the rod block setpoint. l T ical values of reactor core stability are as folle. 'S: 1 4.4-22 l

LGS FSAR o DRAprg ra NATURAL CIRCULATION 51.5% POWER R CTOR CORE STABILITY (105% ROD PATTERN) Deca ratio, X,/Xo 0.625 Resona frequency, Hz 0.287 The calculated lues show the reactor to be in complia ce with the ultimate per mance criteria in the most responsi e attainable mode as ited for the reactor core stabil' y evaluation. The calculated value the decay ratio of the r ctor power dynamic response for th high and low ends of t e automatic flow control range (105% and 8% of rated power) a presented as follows: REACTOR CORE 105% RATED 68% RATED STABILITY ONDITIONS (105S. ROD PATTERN) Decay ratio, X,/Xo .001 0.165 ( Resonant frequency, Hz 0. 6 0.391 l Figure 4.4-7 shows the calculate va iation of the decay ratio over the normal power-flow rang near nd-of-cycle conditions. The channel hydrodynamic per rmance is valuated at the most limiting condition that oce rs at the end f-cycle, with power peaked to the bottom of t core because t control rods are fully withdrawn. The ca ulations yield dec ratios as presented below: CHANNEL HYDRODYN IC NATURAL CIRCU TION PERFORMANC (51.5% POWER Decay Ratio, ,/Xo 0.586 l Resonant requency, Hz 0.364 At this mo responsive attainable mode, the most responsive channel nforms with ultimate performance criteria of 1. decay ratio. he channel performance over the entire range of attai ble operation is well below the threshold of instabi'i,ty. The channel hydrodynamic performance calculation at the ends of ~ th automatic flow-control range indicates the following decay \\s tios: N O l 4.4-23

Oh pS IF [ LGS FSAR u1A CHA.'NEL HYDRODYNAMIC 105% RATED 68% RATED -ERFORMANCE CONDITIONS (105% ROD PATTERN) Decay r io, X,/Xo 0.0 0.269 Frequency, z 0.660 0.508 Conformance with he ultimate performance criterion i further tested by assuming hat the reactor is initially op ating at the most sensitive condt ion. The nuclear system is t en subjected to step disturbances com the pressure regulator

etpoint, control rods, and leve controller setpoint.

T ese time responses are shown in 'gures 4.4-8, 4.4-9, d 4.4-10. It is clear that the decay rati is less than 1.0 d in conformance with the ultimate performa ce criterion. Calculated responses of impo ' ant nuclea system variables to step disturbances from the pr sure reg ator setpoint, control rods, turbine load setpoint, an level controll'er setpoint are tested for rated power-flow condi io and at the nomina @ wer_ corresponding to the lower end'of automatic power-flow control range. Results of the analysis for rat powe and flow are shown in Figures 4.4-11, 4.4-12, 4.4-13 and 4.4 14. Figures 4.4-15, 4.4-16, 4.4-17, and 4.4-18 sh w the resu s of anal'ysis at the low limit of the automatic ow-control ra ge. The analyses are perform using typical con ol settings, since the actual responses ar the result of tests rformed during plant startup, to obta'n optimum control settin s. The disturbances an ytically imposed, one at a t me, using the previously descri d time-domain model are: A pre sure setpoint change of at least 5 ps'l a. N b. A ontrol rod position change equivalent to a Jocal wer change of at least 5% of point (of the m5 nitude l of power at the time of the disturbance) A load demand change of at least 5% of point \\ l d. A reactor water level setpoint change of at lea g \\ 4 i ne h-4.4.5 TESTING AND VERIFICATION l ~~ Tne testing and verification techniques to be used to ensure that t l the planned thermal and hydraulic design cheracteristics of the Lcore have been provided and will remain within required limits g) i 4.4-24 i

LGS FSAR a^ g (~) t ghout core lifetime are discussed in Chapter 14. s. A summary of pr ration and initial startup testing is as follows:/ a. P rational testing: Tests are performe uring the preo ational test program to confirm t construction is comp 1 and that all process a afety equipment is operational. seline data a aken to assist in the evaluation of su uent s. Heat-balance instrumentation and ump flow and core temperature instrumentation a cali ted and setpoints verified. b. Initial s up: Hot functiona or between 5% and 10% pow (sts are conducted with the r er Core performance i nitoredcontinuouslytoensurekhatthereactoris perating within allowable limits (e.g.,' peaking factors, LHGR, etc) and is evaluated periodically to verify the expected and actual core-performance siargins.) g 4.4.6 INSTRUMENTATION REQUIREMENTS ~- The reactor vessel instrumentation monitors the key reactor vessel operating parameters during planned operations. This ensures sufficient control of the parameters. The foll;uin;;* ( reactor vessel sensors are discussed in Sections 7.6 and 7.7: a. eactor vessel temperatur 'N b. Reactor water level N Reactor vessel coolant flow-ra es 1lind differentiah c. pressures / d. Reactor vessel internal pressure eutron monitor.ing.Jysttm / 4.4.6.1 Loose Parts Monitorina System (LPMS) 4 4.4.6.1.1 Design Basis The LPMS is designed to detect loose parts in the a. reactor coolant systems. b. The LPMS is designed to reduce the effects of variations in background noise on system capabilities for the detection of loose parts. The LPMS is designed in conformance with Draft 2 to c. Revision 1 (May 1978) of Regulatory Guide 1.133. I Exceptions to Revision 1 (May 1981) of the guide are as noted below: i M h U" N 4.4-25 Rev. 15, 12/82

1 LGS FSAR rnded A I,edbn 4e Recche kI,"n'nc% l"_ M atyp,,h d m & f ' 06 b* O ~2Oll"#- 4 4.

4.7 REFERENCES

Qd NEDE-240st-P-A ( 4.4-1 General Electric SergiayArtMysts ApasM Nt. TAR)i

ata, Garne Mt%tu d D@fh6 AnbligdtXoh/

gjW E Hptr y Eo a4 ydhup, W9M).V 4.4-2 Core Flow Distribution in a Modern Boiling Water Reacto[ as Measured in MonticelTo,~NEDO-10722A (August, 1976).

4. -3 Peach Bottom Atomic Power Station Units 2 and 3 ety 2

l Analysis Report for Plant Modifications fo Eliminate o l Sionificant Irl-Core Vibration, NEDO-20994 (S 'tember, 1975). 4.4-4 R.C. Martinelli, and D.E. Nelson, "Pr ction of ressure Drops During Forced Convec Boiling of l W ter," ASME Transactions, 70,(1948 pp 695-702. f l f 4.4-5 C.J. Baroczy, "A Systematic Cor lation for Two-Phase Pres re Drop," Heat Transfer nference (Los Anoeles), AICLE, Preprint No. 37, (1966. 4.4-6 W.H. Jen and P.A. Lottes, Analysis of Heat Transfer,

Burnout, essure Drop, d Density Data for High Pressure Wa er," USAEC eport 4627 (1972).

4.4-7 L.G. Neal and .M. Z i, "The Stability of Boiling Water Reactors and Lo s, Nuclear Science and Engineering, 30 j (1967) p. 25. l l 4.4-8 S. Levy, et. al, perience with BWR Fuel Rods Operating Abov Criti 1 Flux," Nucleonics (April, 1965). 4.4-9 Analytical Methods of Plan Transient Evaluations for General ectric Boiling Wa yr Reactor, General Electric l Company BWR Systems Departme t, NEDO-10802 l (Febru ry, 1973). o.4-10 N. ber and J.A. Findlay, "Avera Volumetric t Co entration in Two-Phase Flow Sys ems," ASME T nsactions, Journal of, Heat Transfer, (November, / 65). o.4-11 H.S. Isbin, H.A. Rodriguez, H.C. Larson, nd B.D. Pattie, " Void Fractions in Two-Ptase Flow, American Institute of Chemical Engineers, 5, 4 (Dece er, 1959), 427-432. 4.4-12

Isbin, H.S.,

Sher, N.C., Eddy, K.C., " Void Fractions in }

Two-Phase Steam-Water Flow", American Institute of l Chemical Engineers Journal, _3_, 1 (March, 1957), T38-142.! x ( 4.4-31 Rev. 15, 12/82 d--

DRAFT _,,AR .4-13 J.F. March re, "The Effect of Pressure on Boil k Density in Multiple Rectangular Channels," ANL-5522 (February, 1956). } 4.4-1 E. Janssen, and J.A. Kervinen, "Two-Phase Pressure top in Straight Pipes and Channels; Water-Steam Mixtu sa t 00 to 1400 psia," GEAP-4616 (May, 1964). t 4.4-15 W. Cook, " Boiling Density in Vertical Rectan lar Mult channel Section with Natural Circulation, ANL-5611 (Nove ber, 1956). 4.4-16 G. W. M er, "A Method of Predicting Stead State Boiling r Fractions in Reactor Coolan Channels," WAPD-BT-19 (June, 1960). .4-17 S.I. Rouhani, " Void Measurements in t e Region of Subcooled and ow Quality Boiling," vmposium on Two-i Phase Flow, Uni rsity of Exeter, von, Englana i (June, 1965).

4. 4-18 A. Firstenberg, an L.G. Neal, Kinetic Studies of Heterogeneous Water eactors," STL 372-38 (April 15, 1966).

4.4-19 J.K. Ferrel, "A Study o onvection Boiling Inside Channels," North Caroli State University, Raleigh, North Carolina (Septem er 0, 1964). 4.4 -20 S.2. Rouhani, " Void easurem nts in the Region of Subcooled and Low uality Boi ing," Part II, AE-RTL,788, Aktiebolaget, At energi, Stud vik, Sweden (April, 1966). 4.4-21 H. Christens n, " Power-to-Void Tr sfer Functions," ANL-6385 ( ly, 1961). 4.4 -22 R.A. Eg D.A. Dingee, and J.W. Chas ain, " Vapor Forma n and Behavior in Boiling Heat Transfer," BMI-1 3 (February, 1957). 4. -23 "S pplemental Information for Plant Modif ation to iminate Significant In-Core Vibration", DE-21156, lass III (January, 1976).

4. 4-2 A.B. Jones, " Hydrodynamic Stability of a Boili Channel," KAPL-2170 (October 2, 1961).

4 4-25 A.B. Jones, " Hydrodynamic Stability of a Boiling Ch m :1, Pert @ PL-2208 (April 20, 1962). i i Rev. 15, 12/82 4.4-32

l LGS FSAR c -26 A.B. Jones and D.G. Dight, " Hydrodynamic Stability of a / Boiling Channel, Part 3," KAPL-2290 (June 28, 1963). 4.4-27 A.B. es, " Hydrodynamic Stability of a Bo ng

Channel, 4," KAPL-3070 (August 18 64).

4.4-28 A.B. Jones and W.M. bough,

• activity Stability of a l

Boiling Reactor Part 1, -3072 (September 14, 1964). 4.4-29 A.B. Jones, " React y Stabilit a Boiling Reactor Part 2," KAPL-3 (March 1, 1965). 4.4-30 BWR 4/5 el Design, General Electric Licens Topical Re , NEDE-20944-P, as amended by letter R.H.

holz, to O.D. Park, NRC, dated February 25, 1980. (Octo

(, 1976). l l \\ g yg d.h'! Ao C T l 4.4-33 Rev. 15, 12/82 l l - - - - - - - - - - - - - - - - - - - - - - - - ~ - -

DRAFT ~ ~ mE,,,, (ygpus TABLE 4.4-2 AXIAL POWER DISTRIBUTION USE TO CALCULATE \\ MCPR OPERATING L j De/e/e NODE This FACTOR (Bottom o core) 1 TcL4/e, 7 2 .55 3 .64 4 0.-74 x 5 , 0.85 's 6 O.97 's 7 1,10 s 8 1.21 , ~~ 1.29 NS 10 1.34 11 N,.' 1.38 12 y 1,40 13',- q 1,39 ,A4 N 1.36 s ,- 15 N 1.30 16 1.23 17 1.15 s 18 1.08 i 19 N1.01 / 20 0.93 21 0.84 j/ 22 0,74. / 23 0.60 ( P of core) 24 0.43 \\ 7 o B6o 4.4-S 6.uf 4,4-6 t y,c.k

h* LGS FSAR CORE FLO pIST TION l she DESCRIPT Central Q'S eral e ~ Averace Relative assembly 1.40 s \\ 0.95 O Power 's Relative assemb1Y 0.92 1.06 0.58 s flow N l l l I l

DRApr = " - (NOT USED)' TABLE 4.4-7 CALCULATED VS MEASURED CORE PLATE PRESSURE DROPS / col %E PLATE TEST COMION PRES $URE DROP Power X1ow M9as Calc PLANT SIZE (% r g[,[c,Yated) (msid) (psid) 183-368 8 g. gg 25.10 24.82 i: 251-764 95. 94.9 18.14 17.91 9. 96.9 18.69 18.47 p 224-580 7 3 60.8 5.04 5.05 l 99. 99.3 ' 14.74 14.77 218-548 86.7 100.!6 17T -._19.30 90.3 96.0 16.13 17.85 !~ j/ l 218-560 66.4 59.9 7.47 6.73 79.2 94.4 18.24 17.38 251-764 46.9 / 48.0 6.13 7.55 t 51.3 ' 103.3 18.50 18.00 l 46.5 48.'O 3.99 3.52 l 64.9 70.3\\ 9.42 8.90 l ,/75.9 101.0 \\ 18.51 18.50 g / 57.8 46.4 3.79 3.46 70.1 71.0 9.57 9.32 ,/ 96.4 98.9 19.50 19.25 N\\ \\.\\ l l l l

""-DRAFT TABLE 4.4-8 (NOT U.S D TYPICAL RANGE OF TEST DATA EASURED PARAMETER TEST CONDITIONS Adiab c Tests 4 Spacer single-phase loss rec 1) = 0.5x1 s to 3.5x10s coeffic nt T 0 to 5000F = Lower tie late /j g fice single-phas P 800 to 1400 psia 0 886 coefficien %s = 0.5x16* to 1.5x10* Upper ti phase fric IAh/S Ib/h-fta r X = 0 to 40% Spacer two-p lo coefficient Two-phase friction multiplier \\ Diabatic Tests Heated bundle ressure drop \\P = 800 to 1400 psia C\\ = 0.5x10* to 1.5x10* g Ib/h-fta \\ (1 Reynolds number 'N / s l \\ l l l l l

.a-a I4S FSAR ud E s (nrr uno) w' f "** ' *' *] F DESCRIPTION OF UNCERTAINTIES \\ / STANDARD DEVIATION 00 TITY (% OF POINT) COMMENT- / Feedwe.te flow 1.76 This is the largeit component of total react g power uncertainty. Feedwater 0.76 These are e other significant temperature paramete in core power Reactor 0.5 determ ation. pressure s Core inlet OM Affect quali nd boiling temperature N len not measured Core total 2.5 N di ge /of-calculated flow fr The listed h5. ow corresponds ] un to /Ab/c deviation y in vidual pump i /. dif rence. Channel flow ,/ 3.0 This accounts for manufacturing area / and service induced variations f' in the free flow area within the s channel'. l / Friction actor 10.0 Accounts f'Or uncertainty in the multi ier correlation'gepresenting two-phase pressure losses. \\ C nnel 5.0 Represents variN(ion in the friction pressure loss cha Q cteristics factor of individual channels. Flow multiplier area and pressure loss variations affect the core flow B4stribution, influencing the quality and boiling length in indiv Mual i channels. l ,___ _ __ -_. _ _ _._._... _ _ _-- _.. __ ~ _ _ _ _ _. - - -.

D 33gg----,aq.. rv D, ,/ -~ TABLE 4.4-9 (Cont'd) (P' ' 2 of 2) i STANDARD l DEVIATION OUANTI (% OF POINT) C NT TIP readings 6.3 These sety data are the base from whit'h gross power distri-butio 4s determined. The assi ed certainties include al ele d nd geometrical l mpon e/ contribution g,,/e l from extrapola-i tion v er location Ds el assembly to th fj l / segment. e cluded are / certaintie ntributed by / thg LPRM system. LPRM readings / are Nsed to correct the power distribution calculations for chsges that have occurred ./ since the last TIP survey. s The assigned uncertainty affects power distribution in i the same manner as the base j TIP reading uncertainty. R factor 1.5 This is the last [3f the three power distribution r' elated uncertainties. It is a. functionoftheuncertaigy in local fuel rod power Critical power 3.6 Uncertainty in the GEXL correlation expressed in terms i of critical power l D e.le h, O a m 4. h,e c.jsq e~ 0 s;., '%) 4.4-6, Y IO A* ~ \\

4. -

QAs zl 4.4 -Il -hhro"Ok 4.4-17 4.J -l2-ced 4,4 - 10 a 4.4 -2 I. )' ' S

T L49 FSAR TABLE 4.4-l3 d jj i STABILITY ANALYSIS RESULTS 's (Cycle-1 Most Limiting Conditions) Rod Line Analyzed Natural circulation 51.5% mhd cere (lew Rod pattern 105.0% mted %cccut pree? Decay Ratio Total system stability, x /xe = See Figure 4.4-2. z Reactor core stability, x /Xe = 0. 63 (0.29 Hz a resonant frequency) Channel hydrodynamic = 0.K9 (0.34 Hz performance, x /x. resonant frequency) z A 9

LGS FSAR a. Prevent overpressurization of the nuclear system that could lead to the failure of the RCPB b. Provide automatic depressurization for small breaks in the nuclear system occurring with misoperation of the high pressure coolant injection (HPCI) system so that the low pressure coolant injection (LPCI) and the core spray (CS) systems can operate to protect the fuel barrier c. Permit verification of its operability d. Withstand adverse combinations of loadings and forces resulting from normal, upset, emergency, or faulted conditions l 5.2.2.1.2 Power Generation Design Bases The nuclear pressure relief system MSRVs are designed to meet the i following power generation bases: ~_ _, a. Discharge to the containment suppression pool b. Correctly reclose following operation so that maximum l operational continuity can be obtained onMas a.s Mue.d en Sm,w 5.2.2.1.3 Discussion ,2,3 f Ggg g g g,f g,g Q, The ASME B&PV Code requires that each vessel designed to meet Section III be protected from over: pressure under upset + ~ o s. e eq3e a' lows pes Mwabl'a pressure 'tsf 110% ~cf ' onditfonsh The codee that the lowe,st h t es 1 de ign esstare u er et s ty k' gr sh ety alve eci icat ns srequ a e se at og bel w ve el d yig reshure A d that the t h st fet valke be et k thatstot ackmula d pr&sisu're __s _ t e c_eeh110kof t_. _designjessut foc.upse cond Ridy. udTheno setpoints satisfy the ASME Code specifications lor safetiy a x valves, because all valves open at less than the nuclear system design pressure of 1250 psig. The automatic depressurization capability of the nuclear pressure J l l relief system is evaluated in Section 6.3 and 7.3. 1 The following detailed criteria are used in the selection of l MSRVs: l f a. Meet the requirements of ASME Code, Section III b. Qualify for 100% of nameplate capacity credit for the l overpressure protection function l I I 1 5.2-2 i 1

,,..J d i A, e LG5 FSAR Meet other performance requirements such as gesponse c. time, etc, as necessary to provide relief functions ~ The MSRV discharge piping is designed, installed, and tested in accordance with the ASME Code, Section III, Class I. 5.2.2.1.4 Main Steam Safety / Relief Valve Capacity The MSRV capacity is adequate to limit the primary system pressure, including transients, to the requirements of the ASME B&PV Code, Section III, Nuclear Vessels, up to and including the Summer 1969 Addenda for Limerick. The essential ASME requirements that are all met by this analysis are as follows. It is recognized that the protection of vessels in a nuclear power plant is dependent on many protective systems to relieve or l terminate pressure transients. Installation of pressure-j relieving devices may not independently provjde complete i protection. The MSRV sizing evaluation assumes credi4= Lor. operation of the reactor protection system, which may be tr'ipped by either of two sources a direct or flux trip signal. The direct scram trip signal is derived from position switches mounted on the MSIVs or the turbine stop valves or from pressure ( switches mounted on the dump valve of the turbine control valve hydraulic actuation system. The position switches are actuated when the respective valves are closing prior to 10% travel of full stroke. The pressure switches are actuated when a fast closure of the turbine control valves is initiated. Credit is taken for the safety / relief valves in their ASME Code-qualified self-actuating mode. The rated capacity of the pressure-relieving devices is sufficient to prevent a rise in pressure within the protected vessel of more than 110% of the design pressure (1.10 x 1250 psig = 1375 psig) for events defined in Section 15.2. Full account is taken of the pressure drop on both the inlet and discharge sides of the valves. All MSRVs discharge into the suppression pool through a discharge pipe from each valve that is designed to achieve sonic flow conditions through the valve, thus providing flow independence to discharge piping losses. Table 5.2-5 lists the systems that could initiate during the design basis overpressure event. 5.2.2.2 Desian Evaluation g "'O" l* 05m M (* MS* 5.2.2.2.1 Method of Analysis 'Q2.3 OG GESTAR. E MeS 5.2.-s).'N-l' j lT ign pressu rotectio or the nue r notre NsystTm extens ana al mo represen 11 ess al dynamic characteri s of system e simulate son a larg scomputingx 5.2-3 ~

a s LGS FSAR f ility. These m dels include t hydrodynamih of the., flow loo the reactor k etics, the th mal characte 'stics of the fuel nd its transfer f heat to the colant, and 11 the princi I controller fe tures, such a feedwater fl recircul ion flow, reac gr water level pressure, an load demand. se are represeqted with all heir principa nonlinear features in odels that havb evolved throu't;h extensive alysiswithadqualBWRtes(perience nd favorable omparison of data. A tailed descr Rtion of this del is documented in licensing topi al report NEDOzl0802, " Anal ical Methods of Plant Transient ( Evalu ions for the GE-BWR," (Ref .2-1). MSRVs 'are simulated Jn i l a nonli ear representation, and the odel thereby sllows full investiga(ionofthevat(nyetpointsthatareavailablein h ous valve response times, valve s-I apacities and actuatio s a licable rdware systems l The pical va ve characteristic as modeled is shown incontrol valv\\ Figur 5 associated by e, and LMSIVchk.2-2. sarealsosikass,turbin acteris ulated in the model. 's 5.2.2.2.2

System Design

A parametric study was conducted to determine the required steam flow capacity of the MSRVs based on the following assumptions. 5.2.2.2.2.1 Operating Conditions a. Operating power e 3435 MWt (104.3% of nuclear boiler rated power) b. Vessel dome pressure P <1020 psig c. Steamflow = 14.86 x 10* lb/hr (105% of nuclear boiler rated steamflow) These conditions are the most severe because maximum stored i energy exists at these conditions. At lower power conditions the transients would be le'ss severe. - m %dik, d desuded & Sohee.h S.2.3 d 5.2.2.2.2.2 Transients ' G.Esm E @f 6.2-D,7Jw resu/h of nA1%ss.ent ___ ~ a<s shenn a F ars 5.2-6 t 9 The overpressure prot tion system must accommodatV the most severe pressurization Jr4Hsient. ere are two major transTeffts; s mn closure of afi MsIVs Knd a turb e-generator tqip with a l i roi cident\\ closure f the taarbine st m bypahs system valvks, ? that repres' bot the m st sev re abnorma' operahional\\(ransiehts i tbehaiorwi(tnfinal cesul ing in nuclea system ressure ise. The eva uation\\of transi ant con 'guratiqn has hown that t isol tion v ve clos re is ightly m re sevgre wh credit,- is aken ly fo indirec ly deri ed sera thergfore, t is s s th overpt ssure.p tectio basis ent an't shown, in/ %, Qe a 5.2-4

r=merm h2 { LGS FSAR a i Figure 5.2-1. Table 5.2-8 lists the sequence of events for the main steam line isolation closure event with flux sefam and with the installed MSRV capacity. 5.2.2.2.2.3 Scram a. Scram reactivity curve - Figure 5.2-3 b. Control rod drive scram motion Figure 5.2-3 5.2.2.2.2.4 MSRV Transient Analysis Specifications a. Valve groups: 3 b. Pressure setpoint (maximum safety limit): 1. 1142 psig - group 1 2. 1152 psig - group 2 3. 1162 psig - group 3 The setpoints are assumed at a conservatively high level above the nominal setpoints as shown by Table 5.2-2. This is to account for initial setpoint errors and any instrument setpoint drift that might occur during operation. Typically the assumed setpoints in the analysis are 1% to 2% above the actual nominal setpoints. Highly conservative MSRV response characteristics are also assumed. 5.2.2.2.2.5 MSRV Capacity Sizing of MSRV capacity is based on establishing an adequate margin from the peak vessel pressure to the vessel code limit (1375 psig) in response to the reference transients (Section 5.2.2.2.2.2). M GESTAR H (Ref 6.2-0. 'as deumented ln susedan 5.2.3 5.2.2.2.3 Evaluation of Results of $ls aul sss art. Oeun en Es y xv.s 5.2.2.2.3.1 MSRV Capacity

6. 2 -4 W F. 2-6, The required MSRV capacity is determined by analyzing the 7

pressure rise from an MSIV closure with flux scram transient,j st I pne ons .5 , eumeu v uw up auan ac une u coine-netas t l esign onditi s at a aximum ssel dome pre sure o 1020 ig l T anal is hyp hetica ly assum th failure f the ' rect l iso ation alve p ition ram. Th rea tor is s t down y th back ind ect, h h neut n flux ram. For th analysi the safet etpoi ts are sumed be in er ge of 1 2 to 1 2 p 'g. e ana sis in ' cates t at the d ign valve ca acity ilk rap ble o y aint 'ning a adequa margin elo the pea ASME m

ode allowable pressure in ha nuc _ ar syst (1

,5 osia). i 5.2-5 i l l

.~. l Iyf3p

g. -=acrm i

LGS FSAR i gure 5. -1showscurkv produced by t analysis. e sequence o events in TabTe 5.2-8 assumed 4p this analysts was investigateh to meet code rMquirements and haluate the pragsure \\ ) \\s relief system xclusively, s U er the Generalsgequirements for Protection Against 's Ove ressure as given in Section III of the ASME B6PV Code, Ng credi\\can be allowedsfor a scram from the reactor protection system.\\ In addition, bredit is also'taken for the protective j circuits \\that are indire'etly derived when determining the required MSRV capacity. The backup reactor high neutron' flux cram is conservatively applied as a design basis in determining t t e required capacity of the MSRVs. Application of the direct position scram's in the design basis could be'used, since they qual'ify as acceptable pressure protection devices when j s determining the required safety / relief valve capacity of nuclear i vessels under the provisions of the'ASME Code. s 's N + The parametric relationship between peak vessel (bottom) pressure i ynd MSRV capacity for the MSIV transient with high flux and position trip scram is described in Figure 5.2-4. Also shown in Figure 5.2-4 is the parametric relationship between peak vessel (bottom) pressure and MSRV capacity for the turbine trip with a coincident closure of the turbine bypass valves and direct scram, t which is the most' severe transient when direct scram is considered. Pressures shown for flux scram result only with i multiple failure in th,e redundant' direct scram system. 3 (he time response of the, vessel pressure to the MSIV transient \\d wt h flux scram and the turbine trip w'ith a coincident closure of /\\ s th urbine bypass valves and direct scram for 14 valves is l \\ illu ated in Figure 5.2-5.'N This shows that the pressure at the Ch vessel'hottom excee's 1250 ps'lg for less than 5.8 seconds, which I 6 / is not 1 g enough to*xtransfer'any appreciable amount of heat / into the v sel metal hd which 'is at a temperature well below / 5500F at the start of th'e_ transient. s 5.2.2.2.3.2 Pressure Drop in Inlet and Discharga Pressure drop on the piping from the reactor vessel to the MSRV is taken into account in calculating the maximum vessel pressures. Pressure drop in the discharge piping to the suppression pool is limited by proper discharge line sizing to prevent the backpressure on each MSRV from exceeding 40% of the valve inlet pressure, thus ensuring choked flow in the valve orifice and no reduction of valve capacity due to the discharge piping. Each MSRV has its own separate discharge line. '1 5.2-6

g y LGS FSAR r ar ba~ d d b t 5.2.5.12 Conformance to Reculatory Guide 1.45 (May 1973) - Reactor Coolant Pressure Boundary Leakace Detection System The leak detection system design is in conformance with the guidelines of Regulatory Guide 1.45 except that, with reference to Paragraph C.2 of the guide, the containment airborne radiation monitors may not always be capable of detecting a leak rate of 1 gpm in one hour, as explained in Section 5.2.5.5. The procedures and technical specification limits recommended by the guide will be followed during operation. 5.2.5.13 Seismic Capability of Leak Detection System The RCPB leak detection system is designed to seismic Category I criteria to remain functional following a safe shutdown earthgualze (SSE). Instrumentation associated with the plant drainage system is not designed to seismic Category I criteria. 5.

2.6 REFERENCES

~ q/ f.2-1\\R. Linford,, "AnalyticalsMethods ol Plant Transient / 5 Yvaluation for. the General-Electric Boiling Water Reaction," NEDO-10802 (Aoril'1973). 4 m 7 'f 5.2-2 J.M. Skarpelos and J.W. Bagg, " Chloride Control in BWR Coolants," NEDO-10899, (June 1973). 5.2-3 W.L. Williams, Corrosion, Vol 13, p. 539t (1957). 5.2-4 M.B. Reynolds, Failure Bonavior in ASTM A106B Pipes j Containing Axial Through-Wall Flows, GEAP-5620 (April g 1968). ~ \\ 5.2-5 " Investigation and Evaluation of Cracking in Austenitic i Stainless Steel Piping of Boiling Water Reactor Plants," NUREG 76/067/ NRC/PCSG (October 1975), s h. "Genecoj Elselde, Sta.ded App [#b,o/h4 for fea/c.- Lef ^,y,;,,,,y r 6,2 - 1 $s "%hd $4chs S lemef NEDE-Mott-P-A ad l'ets-9A0ll-P-A-lls (la}es} 9?me) w;s;g), 5.2-45

JRAFT' ~ 1 CLOSING PATH OPENING PATH er \\tr W 2 VALVE CLOSURE OVED CAPACITY CHARACTERISTICS 4 -. y +- ] y _ r- - W f i 5 I W I 5 80 1 JL b ~ j 5 i [ y JL o 60 E e Ig JL 0 3 40 s l )L 20 U JL l I i 1 1 l o' 97 98 99 100 101 102 103 104 PERCENT OP PRESSURE SET POINT LIMERICK GENERATING STATION UNITS 1 AND 2 FINAL SAFETY ANALYSIS REPORT SIMULATED SAFETY / RELIEF VALVE CHARACTERISTICS FIGURE 5.2 2

LGS FSAR m en DRAPt 6.3 EMERGENCY CORE COOLING SYSTEMS y 6.3.1 DESIGN BASES AND

SUMMARY

DESCRIPTION Section 6.3.1 provides the design bases for the emergency core I cooling systems (ECCS) and a summary description of these systems, as an introduction to the more detailed design descriptions provided in Section 6.3.2 and to the performance analysis provided in Section 6.3.3. 6.3.1.1 Desion Bases 6.3.1.1.1 Performance and Functional Requirements The ECCS is designed to provide protection against postulated loss-of-coolant accidents (LOCA) caused by ruptures in primary system piping. The functional requirements (e.g., coolant delivery rates) specified in detail in Table 6.3-1, are such that the system performance under all LOCA condilions postulated in the design satisfies the r,equirements of paragraph 50.46, " Acceptance Criteria for Emergency Core Cooling System for Light l Water Cooled Nuclear Power Reactors," of 10 CFR Part 50. These requirements _ the

t impester.t of Jhich 10 th:t-the pert-L^CA 8'~'

8 p s' c!:ddir.g t;;peratoce ' PCT' b: licit;d t: 2200^Ti are summarized in Section 6.3.3.2. In addition, the ECCS is designed to meet the following requirements: a. Protection is provided for any primary system line break up to and. including the double-ended break of the largest line. b. Two independent phenomenological cooling methods (flooding and spraying) are provided to cool the core. I c. One high pressure cooling system is provided, which is capable of maintaining the water level above the top of the core and preventing automatic depressurization system (ADS) actuat' ion for breaks of lines less than 1 inch nominal diameter. d. No operator action is required until.10 minutes after an accident, to allow for operator assessment and decision. e. The ECCS is designed to satisfy all criteria specified in this section for any normal mode of reactor operation. f. A sufficient water source and the necessary piping, pumps, and other hardware are provided so that the containment and reactor core can be flooded for possible core heat removal following a LOCA. I 6.3-1 l., - - - - - -. - - -~ --. - - - --- -

LGS FSAR 3 operational purposes, these changes will be cordance with procedures having similar administrative control.s. The o.nal f/ cal models AN g 6.3.3 ECCS PERFORMANCE EVALUATION documedl m Subsec.bM 6.2.S.2 of GESTAR H Me? & 3-3)~./) The performance of the ECCS is determined through application of the 10 CFR Part 50, Appendix K evaluation models, and by conformance to the acceptance criteria _of 10 CFR Part 50.46.<F ' / TN 205T6 ' Ret 6.3- ), " Gene al Ele tric Co Tdy7natytica M for Lo of Coo nt Anal sis In ccordan e with 40 CFR Part Appen 'x K," p vides a complet descri tion of the thods sed to erform he cale ations. These ethods re su arize herei A su ry des iption the L As is Iso of th{e LOCA e prov ed, eae lete scripti _nts see _ d) s s <Ref 6. -1. J The ECCS performance is evaluated for the entire spectrum of break sizes for postulated LOCAs. The accidents, as listed in Chapter 15, for which ECCS operation is required are located in the following sections: i 15.2.8 Feedwater piping break 15.6.4 Spectrum of BWR steam system piping failures outside of containment 15.6.5 Loss of coolant accidents 6.3.3.1 ECCS Bases for Technical Specifications The maximum average planar linear heat generation rates calculated in this performance analysis provide the bases for Technical Specifications designed to ensure conformance with the acceptance criteria of 10 CFR Part 50.46, " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Cooled Nuclear Power Reactors." Minimum ECCS functional requirements are l specified in Sections 6.3.3.4 and 6.3.3.5; and testing requirements are discussed in Section 6.3.4. Limits on minimum l suppression pool water level are discussed in Section 6.2. l l 6.3-29 Rev. 7, 06/82 i l

1 LGS FSAR wanna 'n g 3 L 6.3.3.2 Acceptance Criteria for ECCS Performance t The applicable acceptance criteria, extracted from 10 CFR Part 50.46, are listed below. For each criterion, applicable parts of Section 6.3.3 (where conformance is demonstrated) are indicated. A detailed desc ription of the methods used to show j compliance is contained indtef 6.3-f(h Sasse% 52.5.2 *F GESTAR li. 1' Criterion 1, Peak Claddina Temperature: "The calculated maximum fuel element cladding temperature shall not exceed 22000F." Conformance to Criterion 1 is shown in Section 6.3.3.7.3 (Break Spectrum); 6.3.3.7.4 (Design Basis Accident); 6.3.3.7.5 (Transition Break); 6.3.3.7.6 (Small Break); and specifically in i Table 6.3-4. Criterion 2, Maximum Claddino Oxidation: "The calculated total oxidation of the cladding shall nowhere exceed 0.17 times the total cladding thickness before oxidation." Conformance to Criterion 2 is shown in Figure 6.3-10 (break spectrum plot), in I Table 6.3-4 (local oxidation versus exposure), and in Table 6.3-5 (break spectrum summary). Criterion 3, Maximum Hydrogen Generation: "The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all the metal in the cladding cylirder surrounding the fuel, excluding the cladding surround'.ng the plenum volume, were to react." Conformance to Criterion 3 is shown in Table 6.3-4. Criterion 4, Coolable Geometry: " Calculated changes in core geometry shall be such that the core remains amenable to cooling." As described in Ref 6.3-1, Section III, conformance to Criterion 4 is demonstrated by conformance to Criteria 1 and 2. Criterion 5, Lono-Term Coolino: "After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core." Conformance to Criterion 5 is demonstrated generically for General Electric BWRs in Ref 6.3-1, Section III.A. Briefly summarized, the core remains covered to at least the jet pump suction elevation, and the uncovered region is cooled by spray cooling and/or by steam generated in the covered part of the core. Rev. 4, 05/82 6.3-30

p7 LGS FSAR DRa 6.3.3.6 Limits on ECCS System Parameters Rake 4. 54sesbu A.G.3.3.6 beh A.6 3 s.7.2 J Appli 4,fGesmr(g,E As.1) The' mits on he ECCS pa meters ar ciscus a in se ions 6.3.3. and 6.3. .7.1. An ber of mponent n a giv

system, to an includi the enti system r most c es, may e out of ervice.

The max' um 11 cwa e time o of serv e is a-netion o the lev 1 of r undan and the pecified st int vals, as discuss in See on 1 . 5. r 6.3.3.7 ECCS Analyses for LOCA 6.3.3.7.1 LOCA Analysis Procedures and Input Variables Rn.kc1b Susash A.G.3.s.7.I a Aprede Aof GESr$5.E(Ref W. fThe procedures approvea fur. i.vCn analysrs conformance ~ ~ ' calculations are described in detail in Ref 6.3-J. These procedures are used in the calculations documented in Section i6.3.3. For convenience, the f'our comput.er codes are brielay i escribed below. The interfaces between the codes are she se matically in Figures II 2a, II 2b, and II 2c in the " Doc ntation of Evaluation Models," Section II.A of ef 6.3-1. The maj interfaces are briefly noted below. Short Term The 1 Hydraulic Model (LAMB): e LAMB code is a model used to ana e the short term ther dynamic and thermal hydraulic be vior of the coola - in the vessel during a l postulated LOCA. In pa icular, LAMB redicts the core flow, core inlet enthalpy, and re press e during the early stages of the reactor vessel blowdown. Fo detailed description of the model and a discussion regardi sources of input to the model refer to the " LAMB Code Docu nt ion," Section II.A.3 of Ref 6.3-1. Transient Critical B r Model (SCAT): Th SCAT code is used to evaluate the shor term thermal hydraulic r onse of the coolant in the core dur g a postulated LOCA. SCAT r eives input from LAMB and anal es the convective heat transfer ocess in the thermally li iting fuel bundle. For a detailed d cription of <the model d a discussion regarding sources of inpu to the model r er_to the " SCAT Code Documentation," Section .A.4 of Ref 6 -1. f 6.3-33 Rev. 4, 05/82

LGS FSAR .i ,,ono Term Thermal Hydraulic Model and Refill /Reflood Model \\ GAFE/REFLOOD): The SAFE /REFLOOD code is a model used'to anal ze I t e long term thermodynamic behavior of the coolant in the ve el. The SAFE /REFLOOD code calculates the uncovering a ref1 ng of the core, and the duration of spray coolin and (for sma reaks) the peak cladding temperatur. For a detailed description the del and a discussion regarding sources of input to the odel refer to the " SAFE Code and REFLOOD Code Documentation," Sections II.A.1 and II.A.2 of Ref 6.3-1. Core Heatup Mode sTE): The CHA E code solves the transient heat transf quations for specific ial planes of each fuel bundle ty , for large breaks. CHASTE ceives input from SCAT, SAFE, a REFLOOD, and calculates claddin temperatures and locali claddi oxidation during the entire LOCA tr 'ent.' For a detai d description of the CHASTE model and a discussi reg ding sources of input,' refer to the " CHASTE Code umentation," Section II.A.5 of Ref 6.3-1. h The significant input variables used by the LOCA codes are listed in Table 6.3-1 and Figure 6.3-11. 6.3.3.7.2 Accident Description of Appsedoc Ask g.n.4 shcfdm A.G.3.3.7.2 GESTAR.h N A detai ef_6.3geddescriptionoftheLOCAcalculationisprovidedin# ^ 7).ffor convenience, a snorr. demaiption of sne majo % s during the DBA LOCA is included here. i Immediat after the postulated double-e ecirculation line break, ves 1 pressure and core flow beg to decrease. The i initial pres re response (Figure 6.3-12' is governed by the closure of the SIVs, and the relative vnlues of energy added to the system by de y heat, and by energy removed from the system by the initial blo own of fluid from th! downcomer. The initial core flow decrease 'gure 6.3-13) is r pid, because the recirculation in e broken loo as lost suction and ceases to pump a ost imme The p in the intact loop coastdown relatively slow is pump coastdown governs the l core flow response for the nex veral seconds. When the jet pump suct ions uncover, calculated flow decreases to near zero. When the recirculation pump suct1 ozzle uncovers, the energy lease rate from the break increases s 'ficantly, and the ssure begins to decrease more rapidly. As a It of the eased rate of vessel pressure loss, the initially subc =v Rev. 4, 05/82 6.3-34 j

LGS FSAR ter in the lower plenum saturates and flashes up through the increasing the core flow. This lower plenum flashing cor contint s at a reduced rate for the next several seconds. Heat transf rates on the fuel cladding (Figure 6.3-14) d ing the early at es of the blowdown are governed primarily the core flow res nse. Nucleate boiling continues in the igh power plane until sh rtly after jet pump uncovery. Boili ransition follows shortly after the core flow loss that re ts from jet pump uncovery. "ilm boiling heat transfer r then apply, with increasing heat <ransfer resulting from t core flow increase during the lower lenum flashing perio. Heat transfer then slowly decreases u til the high pow axial plane uncovers. At that time, convecti heat transf is assumed to cease. a Water level inside the shr Figure 6.3-15) remains high during the early stages of the blo down, ecause of water f' lashing in the core. After a short me, the 1 inside the shroud decreases to uncover the ore. Several conds later, the ECCS is actuated. As a resu the vessel wate evel begins to increase. Some time 1 er, the lower plenum filled, and the core is subsequently pidly recovered. The cladding tempe ature at the high power plane (Fig e 6.3-16) decreases initial y, because nucleate boiling is mainta ed; the 1 heat input dect ses; and the sink temperature decreases. A rapid, short d ation cladding heatup follows the time of oiling transition w n film boiling occurs and the cladding temper ture approaches at of the fuel. The subsequent heatup is slower being go rned by decay heat and CS heat transfer.

Finally, he heatu s terminated when the core is recovered by the

) a ulation of ECCS water. / ( e l l 6.3.3.7.3 Break Spectrum Calculations I A complete spectrum of postulated break sizes and locations is considered in the evaluation of ECCS performance.'"LT.'.c ;cncral # ' 'y ;;;1ytt::: p::::dar:; f;; crnducting b:::h ~'~=ctrum calculati;;; *~~ q;- cre di::a;;;d in C tien !!! B ef==f f : *2 For ease of reference, a summary of all figures and tables presented in Section 6.3.3 is shown in Table 6.3-6. A summary of the results of the break spectrum calculations is shown in tabular form in Table 6.3-5, and graphically in Figure 6.3-10. Conformance to the acceptance criteria (PCT $22000F, local oxidation <17%, and core-wide metal water reaction <1%) is 6.3-35 Rev. 4, 05/82

LGS FSAR The maximum average planar linear heat generation rate, maximum local oxidation, and peak cladding temperature, as functions of exposurefromthe)bC"i:TZ analysis of the DBA LOCA, are shown in Table 6.3-4. Important variables in two other large break calculations (break size = 0.80 x DBA LOCA break size, 0.60 x DBA LOCA break size) are shown in Figures 6.3-22 through 6.3-37. These results (the same variables as the first eight noted above for the DBA LOCA) are taken from the lead plant analysis for this product line. 6.3.3.7.5 Transition Recirculation Line Break Calculations Important variables from the analysis of the transition (1.0 fta) break are shown in Figures 6.3-38 through 6.3-49. These variables are: a. Core average pressure (large break methods) as a function of time - f se; Li"2 ::d;, - e b. Core flow (large break methods) as a function of time,s-- fre-LA"3 cedc 3_ c. Core inlet enthalpy (large break methods) as a function of time,fre; La"" ::d l d. Minimum critical power ratio (large break methods) as a function of timeg free Sc2T cede - - e. Water level (large break methods) as a function of times s f ce S*FE '"EFLOOO ced; " f. Pressure (large break methods) as a function of time,*- fre; CATE/"EFLOOD ::d: g. Fuel rod convective heat transfer coefficient (large break methods) as a function of time; fee-CM^STE ced= S ~ i h. Peak cladding temperature (large break methods) as a function of time # fre-CMASTE rede - I 6.3-37 Rev. 4, 05/82 l

LGS FSAR ORAFmi demonstrated. Details of calculations fod' specific breaks are included in subsequent paragraphs. y i The analysis results presented in this section are obtained from a typical LOCA analysis, which is representative of this plant size and product line. A plant-specific LOCA analysis will be i submitted as an amendment later. t 6.3.3.7.4 Large Recirculation Line Break Calculation \\ Important variables from the analyses of the DBA LOCA are shown in Figures 6.3-12 through 6.3-21. These variables are: a. Core average pressure as a function of time -I6wm L."f _ J ' g b. Core flow as a function of timer fre; L.'Z" rede c. Fuel rod convective heat transfer coefficient as a function of timer fs u s C:' ACTE red? " d. Water level as function of timer frer 1 FE%"#?LC^" - 'e e. Peak cladding temperature as a function of time -frer A-r -OHA0T2 Ewh f. Core inlet enthalpy as a function of time, fts L "."O i g. Minimum critical power ratio as a function of time -frer " J r fgy-0 CAT cede h Pressure as a function of time,--Ice... CAT:L'" TLOOO c;d: i. Average fuel temperature as a function of time -fse; J r gA5IZ ce04 j. PCT rod internal pressure as a function of time -fr:: #' ) en. __2- ,qy-s..-.e., Rev. 4, 05/82 6.3-36

h{ LGS FSAR i. Water level (small break methods) as a function of time,A-

• _- _ ~ ~ ~ '"i F L ^ ^ 2 :: d : :

j. Pressure (small breaks methods) as a function of time r*' c frer S".?"/""FLOOO ;;d; c k. Fuel rod convective heat transfer coefficients (small break methods) as a function of timec-frc REFLOOD red,- 1. Peak cladding temperature (small break methods) as a function of time,frcr REFLOOD : d: 6.3.3.7.6 Small Recirculation Line Break Calculations Important variables from the analysis of the small break yielding the highest cladding temperature are hhown in Figuren 6.3-50 through 6.3-53. These variables are: a. Water level as a function of time frer SAFE /REFLOOS' r code b. Pressure as a function of time,frer SAFE /REFLOOD cod;'~~ c. Convective heat transfer coefficients as a function of timer 'ce; ":FL^^D :;i: : d. Peak cladding temperature as a function of timeg f r - "- REFLOOD-erde,, The same variables resulting from the analysis of a less limiting small break are shown in Figures 6.3-54 through 6.3-57. 6.3.3.7.7 Calculations for Other Break Locations Reactor water level and vessel pressure frer th: C'FE/REFLOOD "- j ~1-ced;, and fuel rod convective heat transfer coefficients and peak cladding temperature '--- - """' ^^" - ' *are shown in Figures 6.3-58 through 6.3-61 for a CS line break and in Figures 6.3-62 through 6.3-65 for a feedwater line break. Figures 6.3-66 and 6.'3-67 show the reactor water level and vessel pressure ire; th:3 Rev. 4, 05/82 6.3-38

[ LGS FSAR g.h7fL - " "" """' ^^" - '- for a main steam line break inside the containment, u i An analysis was also performed for a main steam line break outside the containment. Reactor water level and vessel pressure Q 'l = th: C'.it," T*G00 cede,r-' fuel rod convective heat transferj coefficients and the peak cladding temperature 'rer th: R;"LOOD j

d; are shown in Figures 6.3-68 hrough 6.3-71 cr 6.3 3 7.6 Impmsal Decay hlra.f Corru fJn kFee k shechbn A.4.5.3.78 dheadm A /GESTAR E(ge$ 4*3-2),

6.3.3.8 LOCA Analysis Conclusions Having shown compliance with the applicable acceptance criteria of Section '6.3'.3.2, it is concluded that the ECCS will perform its function in an acceptable manner and meet all of the 10 CFR Part 50.46 acceptance criteria, given operation at or below the maximum average planar linear heat generation-rates in Table 6.3-4. ~ 6.3.4 TESTS AND INSPECTIONS 6.3.4.1 ECCS Performance Tests All systems of the ECCS are tested for their operational ECCS function during the preoperational and/or startup test program. Each component is tested for power source, range, direction of rotation, setpoint, limit switch setting, torque switch setting, l etc., as applicable. Each pump is tested for flow capacity for a comparison with vendor data (this test is also used to verify flow-measuring capability). The flow tests involve the same suction and discharge source; i.e., the suppression pool or condensate storage tank. All logic elements are tested individually and as a system to verify complete system response to emergency signals, including the ability of valves to revert to the ECCS alignment from other positions. Finally, the entire system is tested for response time and flow capacity, taking suction from its normal source and delivering flow into the reactor vessel. This last series of tests is performed with power supplied from both offsite power and onsite emergency power. 6.3-39 Rev. 4, 05/82

n 2 LGS FSAR vessel low water level and drywell high-pressure plus the indication that at least one CS loop or LPCI pump is.pperating. The HPCI, CS, and LPCI systems automatically realign from system flow test modes to the emergency core cooling mode of operation following receipt of an automatic initiation signal. The CS and LPCI systems injection into the RPV begins when reactor pressure decreases to system discharge shutoff pressure. HPCI injection begins as soon as the HPCI turbine pump is up to speed and the injection valve is open, since the HPCI is capable of injecting water at full flow into the RPV at pressures up to the reactor pressure specified in Mode A of Figure 6.3-1. 6.

3.6 REFERENCES

(NEDE-Jof44-F November I176 / 6.3-1.%G_eneral Electric _ Comnany Analytical Model for Loss of O'C661&n't' A'nilysis 1n Acco'_rda'nce' w'ifh' 10 CFk W Ambe6 dix K ~ ( & dr Ud 2E5EE Edu ' c 1974> and~6ensFal Eicy_LG'd "rwiii11 1Re ood Ca kulation, (Supple nt to Sa e Cod 'Dese I tion) \\ransmit d to US RC by 1 .ter, . L. 4ynrey o Victor,Etello. r. (D , ber 2'1L 197 m H. M. H i rs ch, "M.e t h o_ds _ f or _ Ca l cu l_at i n_a S a f_e Te s t_ 6.3-2. _In<:ervals and AIroVa61e Beo~ air' Time's for~ Eridin'ee' red ]3aj:eduard 'Systeins?'NEDO'-1~07'39 IJanuary 1MI). 9 u.edly[' lk Se belu Tul, dc/w 4 3-3 "4ceeW Elecinc. Stuked A a f NEDe. -24ctl-r-A ud MEDE-Mc Il-ss "U.aki Skl%. 5 P-A-us (Ideshymed revisiA). 6.3-43 Rev. 4, 05/82 -- '~-

n cms,AP l LGS FSAR 2 Initial conditions chosen for these analyses are conservative and tend to force the analytical results to be more severe than would otherwise be expected. These analyses, unlike the pump trip series, will be unaffected by deviations in pump / pump motor and driveline inertias because the flow controllers are what cause rapid recirculation decreases. 15.3.2.4 Barrier Performance The barrier performance considerations for these events are the same as those discussed in Section 15.3.1.2.3. 15.3.2.5 Radiolocical Consecuences While the consequence of this transient does not result in fuel failure, it does result in the discharge of normal coolant activity to the suppression pool via MSRV operation. Because this activity is contained in the primary containment, there will be no exposure to operating personnel. This transient does not result in an uncontrolled release to the environment, so the plant operator can choose to leave the activity bottled up in the containment, or discharge it to the environment under controlled release conditions. If purging of the containment is chosen, the release will be in accordance with established technical specifications and, at the worst, would only result in a small i increase in the yearly integrated exposure level. I 15.3.3 RECIRCULATION PUMP SEIZURE 15.3.3.1 Identification of Causes and Frequency Classification The seizure of a recirculation pump is considered as a design basis accident (DBA). It has been evaluated as being a very mild accident in relation to other DBAs such as the loss of coolant accident (LOCA). The analysis has been conducted with consideration to both single-and two-loop operations. A 4 failed d.scassoon as geven m.%buefen S.1.F.6 sf GE57AK.ZC (Ref /S.5-/), l The seizure accident postulated would not be due to a sudden, complete failure of such a device. Safe shutdown components l (e.g., electrical breakers, protective circuits) would preclude an instantaneous seizure accident. 15.3.3.1.1 Identification of Causes The case of recirculation pump seizure represents the extremely i I unlikely event of instantaneous stoppage of the pump motor shaft of one recirculation pump. This accident produces a very rapid decrease of core flow as the result of the large hydraulic resistance introduced by the stopped rotor. l 1 l l 15.3-7

LGS FSAR m I 5 15.3.4.5 Barrier Performance '. a vd The bypass valves, and momentary opening of some of the MSRVs, limit the pressure well within the range allowed by the ASME vessel code. Therefore, the RCPB is not threatened by overpressure. 15.3.4.6 Radiolooical Consecuences While the consequence of this accident does not result in fuel failure, it does result in the discharge of normal coolant activity to the suppression pool via MSRV operation. Because this activity is contained in the primary containment, there will be no exposure to operating personnel. This accident does not result in an uncontrolled release to the environment, so the plant operator can choose to leave the activity bottled up in the containment, or discharge it to the environment under controlled conditions. If purging of the containment is chosen, the release will be in accordance with established technical specifications and, at the worst, would only result in a small increase in the yearly integrated exposure level. I6.3.5 RETERENCES Is.s -t %awalEledrL Sfar,dard,4 lkbh -G %be kI," y suladh Se %}ed Shdes L lenf "riew-zhn-3 g N ud hf0E-2401I-P-A-MS Otduf yd rev'is,x). i 15.3-12 r y

---

.,y,.

,-,---.-_-,.---7. - -,, - - -. - -, -, ~ -..

LGS FSAR also put in an incorrect location. Third, the misplaced bundles would have to be overlooked during the core verification,* performed following initial core loading. 15.4.7.1.2 Frequency Classification This accident occurs when'a fuel bundle is loaded into the wrong location in the core. It is assumed the bundle is misplaced in the worst possible location, and the plant is operated with the mislocated bundle. This accident is categorized as an infrequent incident based upon the following data. Expected Frequency: 0.004 events / operating cycle The above number is based upon past experience. The only misloading accidents that have occurred in the past were in reload cores where only two errors are necessary. Therefore, the frequency of occurrence for initial cores is even lower since three errors must occur concurrently. 15.4.7.2 Sequence of Events and System Operation The postulated sequence of transients for the misplaced bundle accident (MBA) is presented in Table 15.4-5. Fuel loading errors, undetected by incore instrumentation following fueling operations, may result in undetected reductions in thermal margins during powre operations. No detection is assumed and, therefore, no c'.crective operator action or automatic protection system functioning occurs. 15.4.7.2.1 The Effect of Single Failures and Operator Errors This analysis already' represents the worst case (i.e., operation of a misplaced bundle with three SAF or SOE) and there are no further operator errors which can make the accident results any worse. It is felt that this section is not applicable to this accident. Refer to Section 15.9 for further details. 15.4.7.3 Core and System Performance f[ i 5. 4.7. 3.' Math atical del Amm 4 4e W m M us b e46 amolu H, t5 4 w u i T d le (6.4 - 6 A t ree-dime ional WR simul or mod is used t ~ calculaetnef/b"r* core erforma e res ting from the acc ent. This odel i 3 descr ed in d all i Ref 15.4-gid-3.j s. 4 l .4.7. 2 Inpu Para ters and itial C ditions The initia core con ists f bundles ith avera enrichmen that re hi , medium or 1 w with cor esponding different gadoll la con entratio T e fuel bun le loadina rror with e 1S.4-14 I ,,n n.,., ..,,.-.-----~~--,----n-,--..------

(p~ JiWT LGS FSAR I \\ [ sever t consequences occ s at beginning-of-cy e (BOC) when low en 'ched bundle (which ould be loaded at th eriphery) is intercha ed with a high ene hed bundle located ad cent to a LPRM and p edicted to have the highest LHGR and/or 1 est CPR in the core. ter the loading er r is made and has go undetected, i is assumed for pu ses of conservatism hat the operator uses control pattern th t places the limitin

bundle, in the four-bund e array containing the misplaced bundle, on design thermal li ts as recorded by the LPRM.

A a result of loadi the low enriche bundle in an imprope lo tion, the reading f the adjacent L M decreases. Con quently, because t ere are no instr ents in the three mirro images of this fo bundle array, t e operator believes these a rays are operatin at the same powe as the instrumented one, whe in fact they are t (since no loa 'ng error occurred in these adrants). As a r ult of placing e instrumented array on li ts, the three mir r image arrays xceed the design limit. By re lacing the high en ched bundle (w th the greatest power peaking) ith the low enric d bundle, it i assured that the difference i power peaking bet een the instru nted and the % n-instrumented rays is maximum, rather, that e MCPR and M GR is the upper und for this erro. Other ssumed input pa meters are given n Table 15.4-6 and igure .4-8. 15.4.7.3._3 Results O Results of analyzing the worst fuel bundle loading error are reported in Table 15.4-7. As can be seen, MCPR remains well above the point where boiling transition would be expected to occur, and the MLHGR does not exceed the 1% plastic strain limit for the cladding. Therefore, no fuel damage occurs as a result of this accident. 5.4. .3.4 Cons' eration o neertainties In order assure th onservati of this ana is, major nput parameters e taken as worst case, i.e., the b die is pla n the locati with the h hest LHGR d/or the lo t CPR in t core, and t bundle is erating on esign therma limits. This sures that he MCPR and he MLHGR a the upper N nds for the err X 15.4.7.4 Barrier Performance An evaluation of the barrier performance was not made for this accident ince it is a very mild and highly localiced. No percepti le change in the core pressure would be observed. 15.4-15 j

i LGS FSAR DBAFT 15.4.7.5 Radioloolcal Consequences An evaluation of the radiological consequences is not required for this accidente since no radioactive material is released from the fuel. y l 15.4.8 SPECTRUM OF ROD EJECTION ACCIDENTS This is not applicable to BWRs since the BWR has precluded this transient by incorporating into its design mechanical equipment which restricts any movement of the control rod drive (CRD) 1 system assemblies. The CRD housing support assemblies are described in Chapter 4. i c.oiibl ad dep acciCaoe.s u d $$c f 15.4.9 CONTROL ROD DROP ACCIDENT M

• _. %, ecbenhed k S d sscf & S,2.i,1

/ 15.4. dentification of Causes and Frecuency Classification (15 4.9.1.1 Identification or Aauses ~ ~7- ' The co rol rod drop accident (CRDA) is the result of a postulat accident in which a high worth control rod within l constraints f the banked position rod sequence control s (RSCS) dro em 'the core. ps m the full inserted or intermediate po icn in The h1 est worth rod becomes'decoupled m its drive mechanism. The mec isra is withdrawn but the oupled control rod is assumed to be s k in place. At a 1 e moment, the

control rod suddenly falls ee end drop o the CRD position.

This r s in removal o arge ative reactivity fro /the cor nd results in ocalized

ion, more detailed discus given in 15.4-3.

15.4.9.1.2 FrecueDc9 Chrssification The CRDA is egorized as a limiting fault because is not expecte o occur during the lifetime of the plant; but pos ated to occur, it has consequences that include the ential for the release of radioactive material from the fy 15.4.9.2 Secuence of Events and System Ooeration g ) ~ 5.4.9 2. Sequence f Event (54see.honS.2,s.1 of GETARR Bef e the CRD is possi g occur. h t.La _ct g an k:.. g se the s uenQof f II8*3/ in Ta 15.4-8 t-i...: hi-t tMe - u--t. 15.4-16

LGS FSAR fl .4.9.2.2 System Operation The likely set of circumstances, referred to above, makes possi e the rapid removal of control rod. The dropping of the rod res ts in high reactivity in a small region of the e re. For large, loosely coupled core, this would result in a ighly peaked powe distribution and s sequent operation of s tdown mechanisms. gnificant shifts the spatial power g eration would occur dur g the course of he excursion. The RSCS limits the worth of any c ntrol rod that uld be dropped by regulatin the withdrawa sequence. T ts system i prevents the movement an out-of-equence rod n the 100% to ~ 75% rod density range, a d from the 75% rod de sity point to the preset power level the RS will only allow nked position mode rod withdrawals or insertio This syste is described in 'Ref 15.4-1 for a typical BWR. The RSCS i ed to te t o gh t thi accide The RWM ovide the same prot he RSCS i e RS S-was not unctioning and the RWM was,ec The termination of this excursi s complished by the automatic safety features of he ent utdown mechanisms, Therefore, ru) operator acti dur ng the excursion is required. t Although other normal pla instr amentati and controls are assumed to function, no redit for their op ration is taken in the analysis of this cident. 15.4.9.2.3 The fect of Single allures and perator Errors Systems mitiga ng the consequence of this accid nt are RSCS (or RWM) and APR scram. The RSCS and NWM system are esigned as a redundant s stem network and therefore provide sing failure protectio. The APRM scram system Ls designed to th single failure iterion. Therefore, ter ination of this acc' dent within he limiting results discus ed below is assured. 1 No o rator error (in addition the one that initiates is l acc ent) can result in a more imiting case since the rea or l pr ection system will automa cally terminate the accident. l l ( ction 15.9 provides a deta ed discussion of this subject. 15.4.9.3 Core and System Performance ^ ~~~~~:..... 15.4.9.3.1 Mathematical Model S kb56Cbod l CRDA S.2.f.I d WAR l The analytical methods, ass ptions, and conditions for -(R*f if 4-3). evaluating the excursion a ects of the _ rete:1 red dr:p scci4;ni me are described in detail in Ref: 15.? O, ?S A-A_ - - ' " S. 0 ,w-Tr.:y?t= <nnt eca= 4 dar=4 te previd: ; realicti y-* centervati:: screrrr:it EL-15.4-17

LGS FSAR he associated consequences. The data presented in Ref 15.4-show the RSCS Banked Position mode reduces the control worths to - degree that the detailed analyses presente Refs 15.4-3, 1 and 15.4-5 of the bounding an s presented in Ref 15. are not necessar. iance checks are made instead to verify t max o worth does not exceed i 1% AK. If this crite s not met, then the bo ng analyses are perform he rod worths are determined usin e BWR Simulator Mo Ref 15.4-2). Detailed evaluations, if neces aremadej s ng the methods described in Refs 15.4-3, 15.4-4, and f_ 15 Input Parameters and Initial Conditions b .4.9.3.2 = core at the time of CRDA is assumed to be at the point 1 i cycle hich results in the highest incremental control rod wort to cont n no xenon, to.be in a hot startup condition, and to I have the ntrol rods in sequence A at 50% rod density (gr s 1-4 withdra Removing xenon, which competes well f neutron absorption, in eases the fractional absorptions, orth, of the control rods. The 50% control rod densit lack and white" rod pattern), which minally occurs at ot startup condition, ensures tha ithdrawal e next rod results in the maximum increment of reac i Since the maximum i mental ro orth is maintained at very low values, the po ated CRDA cannot r It in peak enthalpies in excess of calories per gram for any t condition. The data p ented in Section 15.4.9.3.3 show th imum control rod wor Other input parameters and initial condit re given Table 15.4-9. 15.4.9.3.3 Results ^ The radiological evaluations are based upon the assumed failure of 770 fuel rods. The number of rods which exceed the damage l l coreexposur(,providedthepeakenthalpyislessthanthe280 threshold is less than 770 for all plant operating conditions or cal /gm design limit. The results of the compliance-check calculation, as shown in Table 15.4-20, indicate that the maximum incremental rod worth is well below the worth required to cause a CRDA which would result in 280 cal /gm peak fuel enthalpy, M:: ncf 55.0 2, ?5_d-4, s.'" ,;;.0 5'. The conclusion is that the 280-cal /gm design limit is l not exceede nd the assumed failure of 770 fuel rods for the j l radiologica evaluation is conservative. La qd & COA an:.gw.n.d cs ud cadLs b danbsd di Sd5'c h u.2.5 i 4 GESTAR.Ti. (Rd 15.#-3)y s 15.4-18

1 LGS FSAR 15.4.9.4 Barrier Performance i t An evaluation of the barrier performance was not made for this accident, since this is a highly localized accident with no significant change in the gross core temperature or pressure. 15.4.9.5 Radiolocical Consecuences Two separate radiological analyses are provided for this accidents i a. The first is based upon conservative assumptions considered to be acceptable to the NRC for the purpose of determining adequacy of the plant design to meet 10 CFR 100 guidelines. This analysis is referred to as the " design basis analysis." b. The second analysis is based upon assumptions considered to provide a realistic conservative estimate of radiological consequences. This analysis is referred to as the " realistic analysis." A schematic of the leakage path is shown in Figure 15.4-9. 15.4.9.5.1 Design Basis Analysis The design basis analysis is based upon the NRC's Standard Review Plan 15.4.9 (Ref 15.4-3). The specific models, assumptions, and the program used for computer evaluation are described in Ref 15.4-#f Specific parametric values used in the evaluation are presented in Table 15.4-11. 15.4.9.5.1.1 Fission Product Release from Fuel The failure of 770 fuel rods is used for this analysis. The mass fraction of the fuel in the damaged rods which reaches or exceeds the initiation temperature of fuel melting (taken as 28040C) is estimated to be 0.0077. Fuel reaching melt conditions is assumed to release 100% of the noble gas inventory and 50% of the iodine inventory. The remaining fuel in the damaged rods is assumed to release 10% of { both the noble gas and iodine inventories. A maximum equilibrium inventory of fission products in the core is based upon 1000 days of continuous operation at 3458 MWt. No delay time is considered between departure from the above power condition and the initiation of the accident. i 15.4-19

1 LGS FSAR 15.4.9.5.1.2 Fission Product Transport to the-Environgent The transport pathway is shown in Figure 15.4-9 and consists of carryover with steam to the turbine condenser prior to MSIV closure, and leakage from the condenser to_the environment. No credit is taken for the turbine enclosure. Of the activity released from the fuel, 100% of the noble gases and 10% of the iodines are assumed to be carried to the condenser before MSIV closure is complete. Of the activity reaching the condenser, 100% of the noble gases and 10% of the iodines (due to partitioning and plateout) remain airborne. The activity airborne in the condenser is assumed to leak directly to the environment a rate of 1% per day., Radioactive decay is accounted for during residence in the condenser, however it is neglected after release to the environment. ) The activity airborne in th condenser is presented in Table 15.4-12. "in Table 15.4-13.The release of activity to the environment is presented 15.4.9.5.1.3 Results The calculated exposures from the design basis analysis are i presented in Table 15.4-16 and are well within the guidelines of 10 CFR 100. 15.4.9.5.2 Realistic Analysis The realistic analysis is based upon a realistic but still conservative assessment of this accident. assumptions, and the pro The specific models, described in Ref 15.4-f'!* gram used for computer evaluation are Specific values of paratieters used in the evaluation are presented in Table 15.4-11. 15.4.9.5.2.1 Fission Ecoduct Release from Fuel t The following assumptions are used in calculating the fission product activity released from the fuel: t The reactor has been operating at design power fct' a. 1000 days until 30 minutes prior to the accident. When translated into actual plant operation, this assumption means that the reactor was shut down from design power, taken critical, and brought to the initial temperature conditions within 30 minutes of the departure from design power. The 30-minute time represents a conservative estimate of the shortest period in which the required plant changes could be accomplished and i ( t l 1 15.4-20 J m ,_,..m. ..e.,-,_-..,y.-._.,m.,_, __-_-______.,_.__,,_.,-.._.__-,-._,_--_____m._,..,____-_g

~ LGS FSAR defines the decay time to be applied to the. fission product inventory calculations. b. An average of 1.8% of the noble gas activity and 0.32% of the halogen activity in a failed fuel rod is assumed to be released. These percentages are consistent with actual measurements made during defective fuel experiments (Ref 15.4-3&). 7 The fission products produced during the nuclear c. excursion are neglected. The excursion is of such short duration that the fission products generated are negligible in comparison with the fission products already present in the fuel. 15.4.9.5.2.2 Fission Product Transport to the Environment The following assumptions are used in calculating the amount of fission product activity transported from the reactor vessel to the main condenser: The recirculation flow rate is 25% of rated, and the a. steam flow to the condenser is 5% of rated. The 25S. recirculation flow and 5% steam flow are the maximum flow rates compatible with the maximum fuel damage. The 5% steam flow rate is greater than that which would be in effect at the reactor power level assumed in the initial conditions for the accident. This assumption is conservative because it results in the transport of more fission products through the steam lines than would be expected. Because of the relatively long fuel-to-coolant heat transfer time constant, steam flow is not significantly affected by the increased core heat generated within the time required for the MSIVs to achieve full closure. b. The main steam isolation valves (MSIVs) are assumed to i receive an automatic closure signal 0.5 see after detection of high radiation in the main steam lines and to be fully closed at 5 s'ec from the receipt of the closure signal. The signal originates from the main steam line radiation monitors. The total amount of fission product activity transported to the condenser before the steam lines are isolated is, therefore, governed by the 5.5 see isolation time and the l conditions in (a) above. I All of the noble gas activity is assumed to be released c. to the steam space of the reactor vessel. i d. The mass ratio of the halogen concentration in steam, to that of the water is assumed to Se 2%. 15.4-21

LGS FSAR Fission product plate out is neglected in the reactor e. vessel, main steam lines, turbine, and condenseb. Of those fission products released from the fuel and transferred to the condenser, it is assumed that 100% of the noble gases are airborne in the condenser. The iodine activity airborne in the condenser is a function of the partition factor, volume of air, and volume of water. The partition factor assumed applicable is 100. Based upon the above conditions, the activity airborne in the condenser is presented in Table 15.4-14. The following assumptions and conditions are used to evaluate the activity released to the environment: a. The leak rate out of the condenser is 0.5% per day. This leakage is mixed with the turbine enclosure free volume of 4,912,000 ft3 b. The activity released from the condenser becomes airborne in the turbine enclosure. The turbine enclosure ventilation rate is 3.22 air changes per hour. Based upon the above assumptions, the integrated fission product release to the environment is presented in Table 15.4 15 ) 15.4.9.5.2.3 Results The calculated offsite exposures for the realistic analysis are presented in Tables 15.4-16 and demonstrate the width margin of conservatism in the design basis analysis. ~ 15.4.10 REFERENCES Ps ition Withdrawal Sequence}," 15.4-1 Paone, C.J. "Bankad January 197 NEDO-21231 Q 15.4-2.

Wooley, J.A.,

"Three Nencional Boiling Water Reactor, Simulator,"gMay 1976, NEDO-20953)7 I.4-3

Stirn, R.C.,

et al., " Rod Drop Accident analysis ivt Large BWRs," March 1976 (NEDO-10527). e k15.4-4

Stirn, et al.,

" Rod Drop Accida nalysis for l Large BWRs, 1972, Sup 1 (NEDO-10527). 15.4-5

Stirn, R.C.,

"Ro Accident Analysis for ' January 1973, Supp t 2 (NEDO-10527). Lar e l 1 "GE BWR Generic Reload Application for 8x8 Supplement 3 to Revision 1 (NEDO-20360). J ) 164-3 "qe cuJ hecdnc stalwa 4pg;J64 g,u/yht,",,,c/d,; 1%,. %Q SMu Svyplemed," pgos -2aan.t-A ud NEDE-2400-P-A-KS l (Ided cypeved esmien)15 4-22

DRAFT LGS FSAR A 15.4-J USNRC Standard Review Plan, NUREG-75/037, Was'hington, D.C., November 24, 1975. ~ 15.4-Stancavage, P.O. and E.J. Morgan, " Conservative Radiological Accident Evaluation-The CONAC01 Code," March 197gnwo-21143 % 4 15.4-f

Nguyen, D., "Reali tic Accident Anal The RELAC Code,"g ctober 197 NEDO-21142 @ ysis.

O

• /

15.4-)6 Horton, N.R., W.A. Williams, and K.W. Holtzclaw, " Analytical Methods for Evaluating the Radiological Aspects oft eneralElectricBoilingWaterReactorsp March 1969 APED-5756 p N ~. = e e l 15.4-23 l

DRAFT

m.,,AR

{~, (NOTll5FP)s TABLE 15.4-8 SEQUENCE OF EVENTS,OR CONTROL ROD DROP ACCIDENT ~ -SEC EVENT Reactor is operating at 50% rod density pattern RWM is not functioning / Highest wor control. rod blade -tiecomes decouple the CRD ,./ / and with.dtaws the bank hat includes the control 0 ~ deco led rod ,fY-ontrol od sticks in a-fully or an ermediate bank position 0 Antrol d becomes unstuck and drops to the drive sition at the nominal (. measured ve ocity plus three standard viations <1 Reactor goes on 4 positive period and the initial power increase is terminated by the Doppler coefficient <1 AFRM 120% power signal. scrams reactor <5 Scram terminates accident \\ (~ l l

L = t o o DRAFT TABLE 15.4-9 MT USE INPUT PARAMETERS AND INITIAL CONDITIONS E d CONTROL ROD WORTH COMPLIANCE CALC ON 1. Reacto wer, % rated o,o 2. Reactor flow, % rat o,o 3. Core average 0.0 NfNgbfC-0.50 (approx) 4. Control r 5. Average)ue ure,

• 286 6.

Average mode tor temperature, ob 286 7 Xenon state Mone A 'N e I t j 0.-.- ,,}}