ML20008D764
| ML20008D764 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 01/13/1969 |
| From: | CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.) |
| To: | |
| References | |
| NUDOCS 8007300645 | |
| Download: ML20008D764 (175) | |
Text
{{#Wiki_filter:1, ko [ ' TABLE OF CO?iTENTS l Secticn Page
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3 REACTOR 3-1 k 3.1 DESIGN EASES 3-1 3.1.1 PERF0FF.AI;CE CEJECTIVES 3-1 b 3.1.2 LIMITS 3-1 l 3.1.2.1 I;uelear Limits
,5-1 j 3.1.2.2 Reactivity Control Li=its 3-2 3.1.2.3 Therral and Hydraulic Limits 3-2 1
3.1.2.h Mechanical Limits 3-3 3.2 EEACTOR DESIGN 3-6 j 3.2.1 GENEPAL
SUMMARY
3-6 3 3.2.2 NUCLFAR DESIGN AliD EVALUATIO!; 3-8 f 3.2.2.1 Nuclear Characteristics of the Design 3-8 t 3.2.2.2 Nuclear Evaluation 3-20 t 3.2.3 THEPP.AL AND HYDRAULIC DESIGN AND EVALUATION 3-29 3.2.3.1 Ther=al and Hydraulic Characteristics 3-29 3.2.3.2 Thermal and Hydraulic Evaluation 3-39 3.2.h MECHANEAL DESIGN LAYOUT 3-65 3.2.4.1 Internal Layout 3-65 3.2.h.2 Fuel Assemblies 3-71 3.2.h.3 Control Rod Drive System 3-86 ! 3.3 TESTS AND INSPECTIONS 3-100 3.3.1 NUCLEAR TESTS AND INSPECTION 3-100 3.3.1.1 Critical Exteri=ents 3-300 3.3.1.2 Zero Power, Antroach to Fever, and Pc er Testing 3-100 400 300
- MS DOCUMENT CONTAINS 3_i POOR QUALITY PAGES
TABLE OF CC:;TE.TS (Contd) i r Section
?are 3.3.2 THE . MAL ED, HYDRAULIC TESTS A :D I;;SPECTIO:; 3-100 3.3.2.1 Reactor Vessel Flev Distribution and Press =e Ercp Test 3-100 3.3.2.2 Fuel Asse:bly Heat Transfer and Fluid Flev Tests 3-101 ! 3.3.3 FUEL ASSEGLY, CC::TEOL EOD ASSEGLY, A?;D COI;TFDL ROD DRIVE lECEAI;ICAL TESTS A!!D II;SPECTIO:;
3-103 3.3.3.1 Pretetvre Testing 3-103 3.3 3.2 Mcdel Testine_ 3-104 3.3.3.3 Cc renent and/cr Material Tecting 3-10k 3.3.3.h Control Rod Drive Tests an:1 Insrection 3-105 3.3.h II;TE:J;ALS TESTS A?;D IIiSFECTICI; 3-103
- 3. L ' EEFERE;CES 3-110 7
I P 00;n 1 3-11
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._ e . . . . . m ' /f
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LIST OF TAELES Table Uc. Title Pare
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3-1 Core Design, Ther:al, and Hydraulic Data 3-7
/ 3-2 g Luclear Design Data .-c 3-3 Excess Eeactivity Conditions 3-10 3-L First Cycle Reactivity Centrol Distributicn 3-10 3-5 Shutdcen Reactivity Analysis 3-13 3-6 Scluble Ecrcn Levels and Worth 3-1L t
I i 3-7 Exterier neutren Levels and Spectra 3-17 3-8 Calculated and Experi= ental Ecd and Ecd Assembly Cceparison 3-22 3-9 Coefficients of Variaticn 3-3L 3-10 DNE Eesults - Maximu: Design Conditi;n 3-36 3-11 D5E Results - Mcst Probable Conditien 3-37 2 l 3-12 Ect Channel Coolant Cenditions 2-~L 3-13 DN3 Eatics in the Puel Asse:bly Channels (W-3) -
}- Nc=inal Case 3-60 , 3-lh DNE Eatics in the Puel Assembly Channels (W-3) -
j Postulated Worst Case (Design) 3-60 ..i i 3-15 Fuel Assembly Cc ponents, Materials, and Di=ensions 3-72 3-16 Clad Circurferential Stresses 3-77 3-17 ELW High Eurnup 1rradiatien Prcgram - Capsule Puel Test 3-83 3-16 E&W Eigh Eurnup Irradiation Prcgram Schedule 3-6k t 3-19 Centrol Ecd Drive Design Data 3-SS 3-20 Centrol Ecd Assembly Design Data 3-97 3-21 Xenon Centrol Ecd Acce bly resign Data 3-99
) 3-22 Eurnable Pcisen Ecd Asse:b1;. Design Data 3-99 00fs2 3-11:
m
. . - . .. _ - _ _ - . . ~ - . . . - . . _ - -. - . _ . - - - . - _ _ - . _ _ - - - . . -
4 i . I ! 1 / j LIST OF FIGURES t (At Rear of Section) I (' Title Boron Concentration Versus Core Life i Axial Peak to Average Power Versus Xen'n Override Rod Insertion 3-3 Axial Power Profile, Xenon Override Rods 5: "ercent Inserted 3-4 .Incation of Fuel Assemblies Containing Burnable rulson Rods 3-5 Per Cent I;eutron Power Versus Ti=e Following Trip 3-6 Effect of Fuel Temperature (Doppler) on Xenon Oscillatiens - Beginning of Life j [ 3-7 Effect of Pael Te=perature (Doppler) on Xenon Oscillations - I; ear End of Life
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l 3-8 Control of Axial Oscillation With Partial Rods i 3-9 Population Protected, P, and 1-P Versus DIG Ratio (W-3) 3-10 Power Shape Reflecting Increased Axial Power Peak for Ikh-Inch Core 3-11 Distribution of Fuel Rod Peaking 3-12 Possible Fuel Rod DIG's for Maxi =u: Design Conditions - . 36,816 - Rod Core 3-13 Possible Fuel Rod DIi3's for Most Probable Conditions - 36,816 - Rod Core i
- 3-14~ Distribution of Population Protected, P, and 1-P Versus Ita=ber of Rods for Most Probable Conditions
, 3-15 DIG Ratios (W-3) in not Unit Cell Versus Reactor Power ! Maxi =u: Hot Channel Exit Quality Versus Reactor Power 3-16 3-17 Ther=al Conductivity of UO2 3-18 Pael Center Te=perature at the Ect Spot Versus Linear Power a 3-19 .tranber of Data Points Versus CE/:C -
- g L
- -003.63
.3-iv ;
e- ,
JQgF LIST OF FIGURES (Contd) Figure Io. Title 3-20 Hot Channel, Ebetors Versus Percent Population Protected 3-21 Burnout Factor (W-3) Versus Population fcr Various Confidence Levels 3-22 Design Hot Channel and 110 ical Channel Exit Qualities Versus Reactor Power (Without Engineering Hot Cnannel Factors) 3-23 Flow Regine Map for the Hot Unit Cell 3-24 Flow Regbre lup for the Ect Control Rod Cell 3-25 Flow Regine Map fer the Hot Vall Cell 3-26 Flow Regine Itp for the Eot Corner Cell 3-27 Hot Channel D:!3 Ratio ('/-3) Versus Power for various Axial Flux Shapes 3-28 Reactor Coolant Syste: Flow Versus Power 3 Hot Channel D:;3 Ratio (W-3) Versus Pcwer With Reactor Syste: 3-29 Flov and Energy Mixing as Para eters 3-30 Thermal Conductivity of 93 5 Per Cent Eense Sintered UO Fellets 2 3-31 Fuel Center Te perature for Beginning-of-Life Conditions 3- 32 Pael Center Temperature for End-of-Life Conditions 3-33 Fuel Temperature Versus Total Fuel Volute Fracticn fcr Equilibriu: Cycle at End of Life 3-3h Typical Reactor Fuel Assenbly Power Distribution at End of Life Equilibrius Cycle Conditions for 1/8 Core 3-35 Fuel Rod Temperature Profiles at o and 10 kW/ft 3-36 Per Cent Fission Gas Released as a Functi:n of the Average Temperature of the UO Fuel 2 3-3'l Axial Ipcal to Average Burnup and Instantanecus Pcwer Cc= pari-sons 3-38 Fission Gas Release for 1 5 and 1 7 rax/avc Axial Fcwer Shapes 3-39 Gas Fressure Inside the Fuel Clad for Varicus Axial Eurnup and Power Shapes for Ideal Therral Expancien Model ()O ki$dl 3-v
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i dR)F LIST OF FIGURES (Contd) Figure no. Title 3-40 Sensitivity Analysis of the Effects of Fuel Cracking on Fuel-to-Clad Gap Conductance 3-41 Sensitivity Analysis of the Effects of Fuel Cracking on Internal Pressure a 3-42 Sensitivity Analysis of the Effects of Reacter Power on Internal Pressure 3-43 no=inal Fuel Rod Power Feaks and Cell Exit Enthalpy Rise Ratics 3-4h .';axinur Fuel Rod Power Peaks and Cell Exit Enthalpy Rise Ratics 3-45 Calculated and Desi6 n Limit Local Heat Flux Versus Enthalpy in the Eot Unit Cell at the nominal Conditica i 3-46 Calculated and Design 11=it incal Heat Flux vs Enthalpy in the Hot Unit Cell at the Design Condition 3-47 DN3 Ratio (W-3) Versus Power for Varicus Inlet to Cutlet
' Core Bypass LeakaEe 3-48 Rea ctor Vessel and Internals - General Arrangenent 3-49 Reacter Vessel and Internals - Cross section 2 3-50 Core Flooding Arrangement t
j 3-51 Internals T en+ valve 3-52 Fuel Asse=bly i 3-53 Orifice Rod Assedbly 3-54 Eurnable Poison Rod Assenbly 3-55 Centrol Rod Drive - ceneral Arrangenent 3-56 Control Rod Drive - Vertical Section 3-57 Control Rod Drive systen and Trip Elcek Diagr1= 3-58 Control Rod Assebbly 3-59 xenon Centrol Rod Assedbly
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00 ws 3-vi a
O l i LIST OF APPENDICES t Um.ber Title l Modal Analysis of Xenon-Induced Oscillation, 1 3A i t t t i I . 3 i
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b 0a ?x,e. 3-vil
V i - 1 i - 1 Q LIST OF FIGURES (At Rear of Section) d , Figure No. Title f i 3A-1 Reference Axial Stability Index Versus Core Enrichment for j Various Power Shapes at BOL and DOL 3A-2 Minitum Axial Stability Index Versus Core Enrichnent for 6 Various Power Shapes at BOL and EOL 4 3A-3 Stability Index Versus Flatness Beginning and End of Life - i 2452 IG:T (A:ituthal) 1 4 4 i e
~ ' 00'67 3-viii
3 REACTOR 31 TESIGN EASES The reactor is designed to meet the performance objectives specified in 31.1 without exceeding the limits of design and operation specified in 31.2. 3 1.1 PERF0EMARCE OILTECTIVES The reactor is designed to operate initially at 2,L52 MWt('} with sufficient design margins to acco==odate transient operation and instrument error with-out damage to the core and without exceeding the pressure at the relief valve settings in the reactor coolant system. The ultimate operating power level of the reactor core is expected to be 2,552 MWt. This section of the report describes only reactor operation at the initial power level. The fuel rod cladding is designed to maintain its integrity for the antici-pated core life. The effects of gas release, fuel dimensional changes, and corrosion- or irradiation-induced changes in the techanical properties of cladding are considered in the design of fuel assemblies. Reactivity is controlled by control rod assemblies (CRA), burnable poison rod assemblies (EPRA), and soluble boron in the coolant. Sufficient CRA vorth is available to shut the reactor down (keff < 0.99) in the hot condi_ tion at any time during the life cycle with the cost reactive CRA stuck in the fully withdrawn position. Equipment is provided to add coluble boron to the reactor coolant to insure a similar shutdown capability when the reactor coolant is cooled to ambient temperatures. The reactivity worth of CRA, and the rate at which reactivity can be added, is limited to insure that credible reactivity accidents cannot cause a tran-
, sient capable of damaging the reactor coolant system or causing significant fuel failure.
3 1.2 LIMITS 3 1.2.1 INelear Limits The core has been designed to the following nuclear limits :
- a. Fuel has been designed for a caximum burnup of 55,000 igd /1/TU.
- b. The power Doppler coefficient is negative, and the control system is capable of compensating for reactivity changes resulting fro nuclear coefficients, either positive or negative.
- c. Control syste=s will be available to handle core xenon instabili-ties should they occur during operation, without jeopardicing the safety conditions of the reactor coolant syster.
' + ) Pull (rated ) ccre thermal Power.
00 n8 3-1
- d. Tne core vill have sufficient excess reactivity to produce T,he de-es3 sign power level and lifetite without exceeding the control capac-ity or shutdown margin.
- e. Controlled reactivity insertion rates have been limited to 9 2 x 10-5 ( Ag/h)/see for a single regulating CRA group withdrawal, and 7 x 10- (Ak/k)/see for soluble boron renoval.
- f. Reactor control and maneuvering procedures vill not produce peak-to-average power distributions greater than those listed in Table 3-1. The lov vorth of CRA groups inserted during power operation limits power peaks to acceptable values.
3 1.2.2- Reactivity Control Limits Tne control system and the operational procedures vill provide adequate con-trol of the core reactivity and power distribution. Tne follovirs control i limits will be met: [ a. Sufficient control vill be available to produce an adequate shut-
} down margin. ,
- b. The shutdown margin vill be maintained with the CRA of highest worth stuck out of the core.
- c. CRA withdrawal limits the reactivity insertion to 9 2 x 10 0 h (Ak/k)/see on a single regulatire group. Eo f- .f limited to a reactivity insertion of 7 x 10 p n dilution is alro (Ak/k)/sec.
]t Ther=al and Hydraulic Limits l 3 1.2 3 Tne reactor core is designed to meet the follovire limiting ther=al and hy-draulic conditions:
- a. No central melting in the fuel at the design overpower (llh percent).
a
- b. ' A 99 percent confidence that at least 99 5 percent of the fuel rods
,in the core are in no jeopardy of experiencire a departure from nu-cleate boiling (DN'B) durire continuous operation at the design over-power.
- c. Essentially 100 percent confidence that at least 99 96 percent of the fuel rods in the core are in no jeopardy of experiencire a E!!3
'during continuous operation at rated power.
.d. The generation of net steam in the hottest core channe's is permis-sible, but steam voids will be lov enough to, prevent flov instabil-
; ities.
I The. design overpower is the highest credible reactor operating power permitted by the safety syste=. Normal overrever to trip is significantly less than v
) the design-overpover. Core rated ;cver is 2,h52 IGt.
i$. 1 1 ~ 00 M;.S l 3-2 t.-
i 1 1
#[g 3 1.2.h Mechanical Limits i
{- 3 1.2.4.1 Reactor Internals s-t= The reactor internal components are designed to withstand the stresses re-l. sulting from start-up; steady state operation with two, three, or four re-actor coolant pumps running; and shutdown conditions. No damage to the reactor internals will occur as a result of loss of pumping power. l
- Reactor internals will be fabricated from SA-2hO (Type 30h) taterial and vill be designed within the allowable stress levels permitted by the ASME Code, Section III, for normal reactor operation and transients. Structural integ-
) rity of all core support assedbly circumferential velds will be assured by 5 cc=pliance with ASME Code, Sections III and IX, radiographic inspection ac-l ceptance standards, and welding qualifications. l The core support structure vill be designed as a Class I structure, as defined in. Appendix 5A of this report, to resist the effects of seismic disturbances. The basic design guide for the seistic analysis will be AEC publication TID- ! 702k, " Nuclear Reactors and Earthquakes." Lateral deflection and torsional rotation of the lower end of the core sup-
. port assedbly vill be limited to prevent excessive movements resulting from I '
seismic disturbance and thus prevent' interference with control rod assen-blies (CRA). Core drop in the event- of failure of the normal supports will gdqg be limited so that the CRA do not disengage from the fuel assedbly guide tubes.
- The structural internals vill be designed to maintain their functional integ-rity .in the event of a major loss-of-coolant accident as described in 3 2.L.1.
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The dynamic loading resulting from the pressure oscillations because of a loss-of-coolant accident vill not prevent CRA insertion. Internals vent valves are prov'ided to relieve pressure generated by steaming in the core, following a postulated reactor coolant inlet pipe rupture, so that the core vill remain sufficiently cooled.
^
3 1.2.h.2 Fue1' Assemblies The' fuel assemblies are designed to optrate satisfactorily to design burnup and to retain' adequate integrity at the end of life to permit safe removal from the core. The-assemblies are designed to operate safely during steady state and tran-sient conditions under the codbined effects of. flow-induced vibration, clad-
. ding strain _ caused by reactor _ pressure, fission gas pressure, fuel growth, and differential thermal expansion. The cold-worked Zirealoy L cladding is designed to be freestanding. Fbel rods are held in place by mechanical spacer grids 'that are designed to maintain dimensional 'esntrol of the fuel rod spacing throughout the design life without 1. utr;ng cladding integrity.
Contact loads are limited to minimi e fret *.n;- I. OM73 3-3
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l l Tne spacer grids are al.so designed to permit differential ther=al expansion
- 7. J :
of the fuel rods without restraint that would cause distortion of the rods.
)' The fuel c.ssembly upper end fitting and the control red guide tube in the in-ternals structure are both indexed to the grid plate above the fuel assemblies, ! thus insuring continuous align =ent of the guide channels for the CRA. Tne control rod travel is designed so that the rods are always engaged in the fuel asse=bir side tubes, thus insuring that CRA can always be inserted.
The assembly structure is also designed to withstand handling loads, shipping loads, and earthquake loads. Stress and strain for all anticipated normal and abnor=al w rating conditions will be li=ited as follows:
- a. Stresses that are not relieved by small deformations of the caterial vill be prevented frc= leading to failure by not permitting these
, stresses to exceed the yield strength of the material nor to exceed l 1evels that would use in excess of 75 percent of the stress rup-
.r' ture life of the caterial. An example of this type of stress is the circumferential =e=brane stress in the clad due to internal or external pressure.
, b. St'resses that are relieved by s=all deformations of the caterial, and the sirgle occurrence.of which will not make a significant con-tribution to the possibility of a failure, vill be permitted to ex-ceed the yield strength of the =aterial. 7nere such stresses ex-l-g ceed the =aterial yield strength, strain limits vill be set, br. sed
#-V on low-cycle fatigue techniques, using no more than 90 percent of the material fatigue life. Evaluations of cyclic loadings vill be based on conservative esti=ates of the number of cycles to be ex-perienced. An exa=ple of this type of stress is the ther=al stress resulting from the thermal gradient across the clad thickness.
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- c. Co=binations of these two types of stresses, in addition to the in-dividual treat =ent outlined above, vill be evaluated on the low-cycle fatigue basis of Ite= b. Also, clad plastic strain due to
.l diameter increases resulting fro = thermal ratcheting and/or creep, l including the~ effects of internal gas pressure and fuel swelling,
, vill.be li=1ted to about 1 percent.
i i ~d. Minimum clad collapse pressure margins vill be required as follows: h (1) 10 percent =argin over syste= design pressure, on short-time (' collapse, at end void. i H. (2) End void cust not collapse (=ast be either freestanding or have adequate support) on a long-time basis. Y D (3) 10 percent =argin~over syste= operating pressure, on short-
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L. time collapse, at hot spot average te=perature through the ( ' .- clad vall, E f) (h) Clad must be freestanding at design pressure on a short-time basis at <T25 ? hot spot average temperature through the clad
.vall.
4 00 ' 7 ', 3 L- ,w
't 3
3 1.2.h.3 Control Rod Assembly (CRA) Tne control rod clad is designed to the same criteria as the fuel clad, as
!) applicable. Adequate clearance vill be provided between the control rods and if: the guide tubes, which position them within the fuel assembly, so that con-i trol rod overheating vill be avoided and unacceptable techanical interference e
between the control rod and the guide tube vill not occur under any operating condition, including earthquake. Overstressire of the CRA components during a trip will be prevented by mini-mizing the shock loads by snubbing and by providing adequate strength. i; 3 1.2.h.h Control Rod Drive o $ Tne control rod drives provide control rod assembly (CRA) insertion and with-4 drawal rates consistent with the required reactivity changes for reacter op-erational load changes. Tnis rate is based on the worths of the various rod groups, which have been established to limit power-peaking flux patterns to design values. Tne taxitum reactivity addition rate is specified to licit the
- magnitude of a pcssible nuclear excursion resulting fro = a control system or operator =alfunction. Tne normal insertion and withdrawal velocity has been established'as30in./ min.
Tne control rod drives provide a " trip" of the CRA which results in a rapid Q shutdown of the reactor for conditions that cannot be handled by the reactor j -g control syste=. Tne trip is based on the results cf various reactor emergency
- ~
F analyses, including instrument and control delay times and the arount of re-activity that =ust be inserted before deceleration of the CFA occurs. Tne maximum travel time for a 2/3 insertion on a trip co :and of a CFA has been established as 1.h sec.
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The control rod drives can be coupled and uncoupled to their respective CRA vithout any withdrawal movement of the CRA.
! All pressure-containing co:ponents are designed to meet the requirements of >
the ASME Code, Section III, Nuclear Vessels, for Class A vessels. 4 Materials' selected for the control rod drive are capable of operating within
- j ..
!j the specified reactor environment for the life of the techanist withcut any
- j- deleterious effect.s. Adequate clearance vill be provided between the sta-tionary and moving parts of the control rod drives so that the CRA trip time
.* -to-full insertion vill not be adversely affected by techanical interference under all operating conditions and seis=ic disturbances.
! Structural integrity and adherence to allovable stress limits of the control rod drive and related parts during a trip vill be achieved by establishing a
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1i -limit on impact loads through snubbir4 3! 3 1.2.h.5 Methods of Icad Analysis To ",e Employed for Reactor
-j ; laternals and Core
- t. .
4
} Static or dynamic analyses will be used as appropriate.
analysis vill be used- for earthqu2kes and the subeccMd pcrtien cf the In general, dynamic [ 00!!n - -
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- t. '
3-5 v ~J --_,._s- -_ , . . . _ , - ~ , _ . . , _ _ _ _. . ,m _ . . _ _ , _ - . .- _ . ,
g loss-of-coolant accident (IOCA). For the relatively steady state portion of the IOCA, a static analysis will be used. Vnere it is indicated that substantial coupling, ie, interrelationship, exists
/'
between major components of the Raclear Steam System (NSS) such as between the steam generator, the piping, and the vessel, the dynamic analysis will include the response of the entire coupled syste=. Ecwever, where coupling is found to be small, the component or Groups of components will be treated independently of the overall system. The dynamic analysis for LOCA will use predicted pressure-time histories as input t- a lumped-cass model. For earthquakes, actual earthquake records, norma' ei to appropriate ground motion, will be used as input to the model. Tne c_ cut from the analysis will be in the form of internal motions (dis-place ents, velocities, and accelerations ), motions of individual fuel asse:- blies, impact loads between adjacent fuel assemblies, and impact loads be-tween peripheral fuel assemblies and the core shroud. In addition, seismic analysis will also be perforced usire a todal superposi-tion and response spectra approach. I For the simultaneous occurrence of LOCA and the maxirar earthquake (E' ), both j time-history excitations will be input to the system cicaltaneously. Rela-
; tive starting times will be changed until maximum structural motions, indi- ! cating maxican stresses, are obtained. Output will be those mentioned above.
5
.' M Using the output from the lumped-cass model and additional information such as pressure-time histories on separate internals and core co ponents (includ- ; .ing control rods), stresses and deflections will be calculated. Tnese stres-ces and deflections will be compared to the allowable litits for the varfous loading combinations as established in Appendix 5A to insure that they are less than these allowables. ! 32 REACTOR DESIGN 3 2.1 GENERAL SU:G!ARY The important core design, thermal, and hydraulic characteristics are tabu-lated in Table 3-1.
1
)
013 79 3-6 t
1
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- i Table 3-1 Core Design, Thermal, and Hydraulic Ihta
't J Reactor I . 1 Type Pressurized Water ] Rated Heet Output, MWt Vessel Coolant Inlet Temperature, F 2,452 555 Vessel Coolant Outlet Temperature, F 602.8 Core Outlet Temperature, F 60h.3 Operating Pressure,.psig 2,185 Core and Fuel Assemblies Total Number of Fuel Assemblies in Core 177 j Uumber of Fuel Rods per Fuel Assembly 208 1 Number of Control Rods per Control Rod Assembly 16 Number of In-Core Instrumentation Positions per Fuel Assembly , 1 - Fuel Rod Outside Diameter, in. 0.430 Clad Thickness, in. 0.0265 1
' Fuel Rod Pitch, in'. 0 568 Fuel Assembly Fitch Spacing, in. 8.587 Unit Cell Metal / Water Ratio (Volume Basis) 0.82 4 ii}- -
Clad Material Zircaloy h (Cold-Worked) Fuel , Material UOp ll Form Dished-End, Cylindrical Pellets Diameter, in. O.370 y] Active! Length, in. 1hh 4 - Density, % of Theoretical , 93 5
- , Heat Transfer and Fluid Fiov at Rated Power r- Total Heat Transfer Surface in Core, ft h9,73h Average Heat Flux, Btu /hr-ftj s 163,725 Maximum-Heat Flux,. Btu /hr-ft 510,300 Average Power Density in Core, kW/l 79 60 1 .-
Average Thermal Output, kW/ft of Fuel Rod 5.h j Maximum Thermal Output,.kW/ft of Fuel Rod 16.83 R Maximum Clad Surface Temperature, F 65h Average Core Fuel Temperature, F. 1,3h5 jb ' Maximum Fuel Central Temperature at Hot Spot, F Total Reactor. Coolant Flow,,lb/br h,150 131 32 x 106 Core Flov Area (Erfective for Heat Transfer), ft h9 19 b Core Coolant Average Velocity, fps 15 2 Coolant' Outlet-Temperature at Hot Channel,'F- 6!.2.6 h 00 n.2 1 3-7
a 1 l l i t m Ihble 3-1 (Contd) Pcwer Distribution i Maximut/ Average Pcver Eatio, Radial x Local (F;3 Nuclear) - 1.7c Maxi =ur/ Average Power Ratic r Axial (i, Nuclear)
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1.70 Overall Power Estio (Fq Nuclear) 3 03 Pcver Generated in Fbel and Cladding, % 97 3 Ect Channel Factors Power Peaking Factor (Fq) 1.011 i l Flcv Area Eeduction Factor (F.n ) - t Interior Bundle Cells 0 9c Feripheral Eundle Cells 0.97 Local Eeat Flux Factor (F;") 1.Ol' j Ect Spot Maximur/ Average Heat Flux Eatic a (:q nuc an rech) 3 1c m ta n va.. a Design OverpcVer Ratio 1.1h IN3 Eatio at resign Overpower (W-3) 1.71 ENE Eatic at Eated Pcwer (W-3) 2.21 j
) 3 2.2 NUCLEAR DESIGN AND EVALUATION l -
- The basic design of the core satisfies the fellowing require ents:
- t. .
; a. Sufficient excess reactivity is provided to achieve the design power level over the specified fuel cycle. ; b. Sufficient reactivity control is provided to permit safe reactor j operation and shutdown at all times during core life' ite.
1 3 2.2.1 Nuclear Characteristics of the resign 3 2.2.1.1 Excess Reactivity The nuclear design characteristics are given in Table 3-2. The excess reac-tivities associated with varicus cere conditions are tabulated in Table 3-3 The core vill operate for L60 full-pcVer days fcr the first cycle and vill 3 i have a 310 full-power day equilibriu cycle. Lesign limits vill be held with respect to reactivity control and ,'ver distributicn. In-core instruzents-
! tien vill be used to indicate power peaking levels. Single fuel asse1bly re-
! activity information is also included in Table 3-3 go . _
t t 4 =
- v[
.1.}
Table 3-2 Nuclear Design D1ta Frel Assembly Volume Fractions Fuel O.303 Moderator 0 580 Zircaloy 0.102 Stainless Steel 0.003 Void 0.012 1.000 Total UO2 (BOL, First Core), Metric Tons 9h.5 Core Dimensions, in. .- Equivalent Diameter 128.9 Active Hei 6ht . 1hh.0 Unit Cell H O2 to U Atomic Ratio (Fuel Assembly) Cold- 2.85 k,_} ~~~ Hot 2.0L Full-Power Lifetime, Days First Cycle h60 Each Succeeding Cycle 310 i Fuel Irradiation, mwd /IEU f First Cycle Average , 13,5ho
- j. -Succeeding Cycle Average 9,125 Feed Enrichments. v/o U-235 First Cycle 2 30/2 30/2.6h (by Zone)
Control Ihta i Control Rod Material' Ag-In-Cd Number of Full-Length Control Rod Assemblies h9
. Number of Xenon (Part-Iength) Control Rod Assemblies 8 8.0 l} Total Full-Length Control Rod Worth (tak/k), 5 Control Rod Claddin6 Material ,
Type 30k SS 1 oo 7s ,p. i 3-9
1 lI ,, , j _. .. Table 3-3
- } Excess Reactitity Conditions
,i Effective Maltiplication, ke rt (" Cold, Zero Power, No Burnable Poison 1.271 4 i Ect, Zero Power, No Burnable Poison 1.216
} Hot, Rated Power, No Burnable Poison ' 200 ) Hot, Rated Power, With Barnable Poison 1.115 1 Ect, Equilibriu= Xenon, Rated Power, With 3
Barnable Poison 1.0Sh 1 L- Sirgle Fuel Assembly (b) 0 77 Hot Cold(C O.87 (* First Cycle at Beginning of Life (BOL) (b) Based on Highest Probable Enrichment of 3 5 Weight Percent i . A Center-to-Center Asse=bly Pitch of 21 Inches Is Required for This 1: keff in Cold, Nonborated Water With No Xenon or Samarium ll }. ; The mini =u.= critical = ass, with and without xenon and samarium poisoning, 1
=ay be specified as a single assembly or as multiple assemblies in various geometric arrays. The unit fuel asse=bly has been investigated for com-
, parative purposes.- A single cold, clean asse=bly containing a maxitu prob-i able enrichment of 3 5 weight percent is suberitical. Tao asse=blies side-by-side are supercritical except when both equilibrium xenon and samariu are present. Three assemblies side-by-side are supercritical with both equilibrium xenon and samariu= present.
3 2.2.1.2 _ Reactivity Control Distribution Control of excess reactivity is shown in Table 3-L. Table 3 h - t First Cycle Reactivity Control Distribution s-
- f. Ak/k 1
- 1. Controlled by Soluble Boron
- f. a. Moderator Te=perature Deficit (70 to 520 F) 3.h
/ b. F4 uilibriu= Xenon and Samarium 35
. c. Fuel Eurnup and Fission Product Buildup 7.2 00f.77
^
i p I[. 3-10
-r-7 e,ae g c T T--k"y fr yrw* v-ur+- swS' y-g P-y-F w *T-7^* z v' ++maY M@--T- ' ' ' " "
7 0
'~'
i ( LJ
/
Table 3 h (Contd) 9 I ik /k
- 2. Controlled by Burnable Poison Rod Assemblies (EPRA)
Fuel 32rnup and Fission Product Buildup 6.0 d 3 Controlled by Inserted Control Rod Assemblies 1 Transient Xenon (Normally Inserted) 0.8
- h. Controlled by Movable Control Rod Assemblies (CRA)
- a. Ecypler reficit (0 to 100% Rated Power) 1.2
- b. Moderator Temperature Deficit (0 to 15% Power at End of Life, 520 to 579 F) 0.8
- c. Dilution Control 0.2
- d. Shutdown Margin 1.0 Tbtal Movable Control Worth Required 32 5 Available Control Rod Assembly Worths i
- a. Ibtal CRA Worth 8.0
- b. Stuck Rod Worth (Rod of Highest Reactivity Value) (-) 25 i
- c. Pdni=um Available CRA Worth 55
- d. Mini =u Movable CRA Worth Available h.5 Explanation of Items Above
; 1. Soluble Scron
[! Baron in solution is used to control the following relatively slow-covin6 i reactivity changes :
?
I
- a. The moderator deficit in going from anbient to operating terperatures.
The value shown is for the maxirum change which would occur toward the end of the cycle.
- b. Equilibrium xenon and samarium.
1
- c. The excess reactivity required for fuel burnup and fission product
^
buildup throughout cycle life. B Figure 3-1 shows the typical variation ir b:r:n concentratior with life fcr Cycle 1 and the equilibrium cycle. J 0098 3 3-11 1,
y i
; 4g 2. Burnable Poisen The 16 control rod holes Jr 72 of the fuel assemblies not equipped with control rod assemblies vi?.1 be utilized as locations for burnable poison rods. . The 72 element locations are shcvn in Figure 3 h.
i
- 3. Inserted Control .
, Sufficient rod worth remains inserted in the core during normal operation ! to overcome the peak xenon transient following a pcver reduction of 50 l percent of rated power for 90 percent of the fuel cycle. This override
,) capability facilitates the return to normal operating conditions without 3, . extended delays. The presence of these rods in the core during operation j)- does not produce power peaks above the design value, and the shutdcun =ar-
;{. . gin of the core is not adversely affected. Axial power peak variation, resulting frcs partial or full insertion of xenon everride rods, is de-scribed fully in Figures 3-2 and 3-3 The loss of scvable reactivity control due to the insertion of this group produces no shutdevn diffi-culties and is reflected in Table 3-5 .
- h. . Movable Control S
7'- a. Power level changes (Doppler) and regulation, g
- j b. Between 0 and 15 percent of rated power, reactivity cc:pensation by 5,. CRA =ay be required as a result of the linear increase of reactor coolant temperature fres 520 F to the normal cperating value.
'l c. Additional reactivity is held by a group of partially inserted CRA 11 (25 percent insertion maximu=) to allow periodic rather than centin-
] .uous soluble boron dilution. The CRA are inserted to the 25 percent M li=it as the boron is - diluted. Automatic vithdrawal of these CRA
! during operation is allowed to a 5 percent insertion limit where the ij dilution procedure-is again initiated and this group of CRA is
$f reinserted.
. 4 li- d. A shutdown margin of 1-percent Ak/k below the hot critical conditien is also considered as part of the reactivity centrolled by CRA. .
- 5 Rod Worth 1.
A total of 3.2 percent'ah/k("I is required in novable control. Analysis of the h9 full-length CRA under the reference fuel arrangement predicts a total CRA vorth of at least. 8.0 percent ak/h. The stuck-ou RA vorth
~
a vas also evaluated at a value no larger than 2.5. percent Ak/k . This
- evaluation included selection of the highest worth CRA under the first j
CRA-u} condition. -The minimum available CRA verth of h.5 percent
- ak/k .is sufficient to ceet covable centrol requirements. The eight
" Lee - ct include transient centrol. See Table 3-L.
D)First' cycle. See. Table 3 L. 4 (}() f. /.h 3-12 E
1 i 4 xenon control rods (XCRA) are worth from 0.2 to 0.h percent Ak/k. This value { is not included in the 8.0 percent Ak/k rod worth reported above. , f 3 2.2.1 3 Reactivity Enutdown Analysis i
- The ability to shut dove the core under both hot and cold conditions is 11-lustrated in Table 3-5 In this tabulation both the first and equilibrium cycles are evaluated at the beginning of life (BOL) and the end of life (IDL) for shutdown capability.
Table 3-5 Shutdown Reactivity Analysis First Cycle Ecuilibrium Reactivity Effects, 4 ok/k BOL EOL BOL EDL i 1. Maxi =um Enutdown CRA Requirement l Doppler (100 to 0% Power) 1.2 15 1.2 15 0.8 Moderator Deficit (15 to 0% Power) 0.8 0.0 0.0 i Total 1.2 23 1.2 23
- 2. Maximum Available CRA Worth (" -8.0 -8.0 -8.0 -8.0 '
p}; Transient Xe Insertion Worth 0.8 0.0 0.8 0.0 Possible Dilution Insertion 0.2 0.2 0.2 0.2 i 3 Minimum Available CRA Worth i g All CRA In -7 0 -7.8 -7 0 ~7.8 One CRA Stuck-Out(b) -4 5 -5 3 -4 5 -5 3
. L. Minimum Hot Shutdown Margin +
1 All Ch5 In -5.8 -5 5 -5.8 -5 5 One CRA Stuck-Out -3 3 -3 0 -3 3 -3 0
- Total Worth of L9 CRA s (b)CRA of Highest Reactivity Value t
l- Examination of Table 3-5 for Minimum Hot Shutdown Margin (Ite: 4) shows that, j- vith the' highest worth CRA stuck out, the core can be maintained in a suberiti-
! cal condition. Normal conditions indicate a minimum hot shutdown targin of
'5 5 percent'Ak/k at the end of life.
Under conditions where a cooldown to reactor building a:bient temperature is
. required, concentrated soluble boron vill.be added to the reactor coclant to produce a shutdevn =argin of at least 1 percent Ak/k. Beginning-of-life teren levels' for several ccre cendit. tens are listed 'in Tatle 3-6 alcng with teren verth values. The conditions shcvn with no CRA illustrate the highest rc-
, quirements.
i 003B0 1.- 3-13
Table 3-6
}
Soluble Boron Levels and Worth (First Cycle) BOL Baron Levels, Core Conditions ppm
- 1. Cold, k ,, = 0 99 e..
No CRA In 1,290
! All CRA In 810 i One Stuck CRA 960
- 2. Hot, Zero Power, k ,, = 0 99 eu Uo CRA In 1,250 All CRA In ,
h50 One Stuck CRA 700
]
3 Hot, Pated Power, k eff = 1.00 No CEA In 1,030
- 4. Hot, Equilibriue Xe and Sm, Bated Power, k eff = 1.00 No CRA In 680 Core Condition Boron Worth. (% Ak/k)/ ppm i
; Hot 1/100 Cold 1/75 ?
3 2.2.1.L Reactivity Coefficients Beactivity coefficients for= the basis for digital studies involving nor.21 and abnor:21 reactor operatire conditions. Tnese coefficients have been in-vestigated as part of the analysis of this core and are described below as to function and everall rarse of values.
- a. Icrrler Coefficient
! Tne Doppler coefficient reflects the change in reactivity as a l function of fuel temperature. A rise in fuel temperature results in an increase in the effective absorption cross section of the i fuel (the Doppler broadening of the resonance peaks) and a cor-responding reduction in neutron production. Tne range for the Icppler coefficient under operating conditions is expected to be
-1.1 x 10~5 to -1.7 x 10-5 (ak/h)/F.
Ui 00'S* 3-ll
y - - - - - ,. Moderator Void Coefficient
; b.
The moderator void coefficient relates the change in neutrcn multi-She expected plicationtothepresenceofvoidsinthemogerato;.to range for the void coefficient is h.0 x 10- -3 0 x 10-3 (ak/k)/ percent void. I
- c. Moderator Pressure Coefficient f.
I The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the correspondin6 effect on neutron production. This coefficient is opposite in sign and considerably smaller when compared to the rod-erator temperature coefficient. A typical range of pressure coef-ficients over a life cycle would be +h x 10-8 to +3 x 10-0 (ak/k)/ psi. .' d. Moderator Temperature Coefficient The moderator temperature coefficient relates a change in neutron cultiplication to the change in reactor coolant temperature. Re-actors using soluble boron as a reactivity control have fever neg-ative moderator temperature ~ coefficients than do cores controlled solely by movable ' or fixed CRA. The major temperature effect on the
- coolant is a change in density. An increasing coolant temperatu.*e
- ] produces a decrease in water density and an equal percentage re-duction in boron' concentration. The concentration change results
~
h i in a-positive reactivity component by reducing the absorption in the coolant. The magnitude of this component is proportional to the
., total' reactivity held by soluble boron.
The moderator. temperature coefficient has been calculated for three conditions of the hot, clean (no xenon) core with 1030 ppm boron in p the' moderator. '
-Zero Power a =0 ; m 15 Fercent Power a = - 0.32 x lo- (ak/k)/F 100 Fercent.Pcver n = - 0.h2 x 10- (ak/k)/F Since equili'rium xenon.is covered by soluble poison, it follevs that at the full-power condition with xenon, the moderator tem-perature coefficient-vill be even more negative than shown above.
The: moderator temperature coefficient vill be -3 0 x 10- (ak/k)/F
et the end of the equilibrium fuel cycle.
? -e. pH Coetficient Currently,'there is no.cefinite correlaticn to predict 72 rese-
.tivity effects 'c'etween various operatir.6 reactors. ;E efftetr versus'
- 00 iS?.
~
Ta
+
.3-15
reactor operatire time at power, and changes in effects with vari-ous clad, temperature, and water chemistry. Yankee (Ecue, Mass. ), .' Saxton, and Con Edison Indian Point Station No.1 have experienced reactivity changes at the time of pH changes, but there is no clear-cut evidence that pH is the direct influencing variable with-out considerire other items such as clad materials, fuel assembly I crud deposition, systes average temperature, and prior system water 1 chemistry. Saxton experiments have indicatet a pH reactivity effect of 0.16 percent reactivity per pH unit chs.nge with and without local boiling in the core. Operating reactor data and the results of applyirg Saxton observations to the reference reactor are as follows : 1 (1) Tne proposed system pH vill vary from a cold .rmured value of approximately 5 5 to a hot calculated value of 7.8 with 1, LOO pp boron and 3 ppm KOH in solution at the beginning or life. Lifetime bleed cilution to 20 ppm boron vill reduce pH by approximately 0.8 pH units to a hot calculated pH value of 7.0.
' (2) Considering the system make-up rate. of 70 gps, the correspon-ding changes in pH are 0.071 pH units per hour for boron di-lution and 0.231 pH units per hour for KOH dilution. Apply-ing pH vorth values of 0.16 percent ak/k per pH unit as n
observed at Saxton, insertion rates are 3 16 x 10-6 (, percent
~
' ak/k)/see and 1.03 x 10-5(percentak/k)/sec,respectively.
Tnese insertion rates correspond to 1.03 percent power / hour and 3.h percent power / hour, respectively, which are easily
] compensated by the operator or the automatic control system.
3 2.2.1,5 Reactivity Insertion Rates Figure T-7 displays the. integrated rod vorth of three overlapping rod banks
~
as a function of. distance withdrawn. The indicated 3r ups are those used in the core durire power operation. Using approximately 1.2 percent ak/k CPA groups and a 25 in./ min driv ' meed in conjunction with the reactivity re-
-sponse given in Figure 7-7 y? 3 a tax 1=ut reactivity insertion rate cf 9 2 x lo-5.(ak/k)/sec. Tne es 22. reactivity insertion rate for soluble.
l: boren removal'is 7 x 10-6 (ak/k)/sec. t b Power Decay Curves
- 3.2.2.1.6' t
, Figure 3-5 displays the beginning-of.-life power decay curves for the CPA vorths
;- . corresponding to the 1 percent hot shutdown margin with and without a stuck rod. The power decay is-initiated by the trip release of the CPA with a 300
=sec delay from initiation to start of CPA motion. Tnetimerequiredfor2/3 rod insertion is 1.4 sec.
3 2.2.1 7' Neutron Flux Distribution and Spectrum
. _ The neutron flux levels at the-core edge and the pressure vessel vall are given in Table 3-7. - At both locations, the values shown include an axial i
peakirg factor of..l.3, a' scalire factor of 2 and a safety cargin cf : .9 3 3 r 00f33
^
t '3-16 ].
Y
>I 1
I
- Table 3-7 Exterior Neutron Levels and Snectra
+
j Neutron Flux Levels neut/cr2-s Interior Wall of i Flux Core Edge Pressure Vessel {- Groun- (x 10 -0) (x 10 ) i^ l 0.821 MeV to 10 MeV 6.0 3.h 2 1.230 kev to 0.821 MeV 90 l- 3 0.klh eV to 1.230 kev 6.2 75 ] h Less Than 0.41h eV 7.1 5.7 2.1 The. calculations were perferred using The Babcock & Wilecx Cc pany's LIFE
,; code (
TOPIC. g '-293, 3 6 3)-to generate input data for the transport code, A h-group edit is obtained from the LIFE output which includes diffusion coefficients, absorption, re= oval, and fissien cross secticns, and the zeroth and first moments of the scattering cross section. TOPIC is an S ncode designed to solve the 1-dimensional transport equatien in cylindrical coordinates for up to six groups of neutrons. For the radial j1 and azi=uthal variables, a linear approxication to tha transpcrt equation is used; for the polar angle, Gauss quadrature is used. Scattering fune-
'"g tions are represented by a Legendre series. The atituthal angle ca*. be . y partitioned into k to 10 intervals on the half-space between 0 and . The nu=ber of mesh points in the radial direction is restricted by the number of these intervals. For the core exterior flux calculations,.four inter-N vals en the azimuthal vere used. This allows the maxi =ur nunber of resh 4 points (2h0) in the "r" direction to describe the nield cc plex. An ' option is available to use either equal intervals en the a ituthal angle or equal intervals on the cosine of the angle. Equal intervals on the cosine vere chosen since this provides more detail in the forward direction of the flux (toward the vessel). Five Gauss quadrature points were used on the cosine of the polar angle in the half-space between 0 and n.
'i. Results frem the above method of calculation have been ec pared with ther-
=al flux seasgg=ents throu6h an array of iron and water slabs in the LIDO pool reactor. i Althou6h this is 'not a direct co=parisen with fast neu-tron reasure=ents, it does provide a degree of confidence in the nethed since the. =agnitude of the thermal flux in shield regions is scverned by ftst neutren penetration.
t Results of the co=parison shoved that fluxes predicted by the LIFE-TOPIC j calculation were lover, in general, by cbout a factor of 2. Results of i.
. the fast flux calculations are, consequently, increased by a factor cf 2 to predict the nyt in the reactor vessel.
The folleving censervatisms were ,also incorporated in the calculaticns:
'J
- a. Neutron fluxes outside the core are based en a taximum p:ver density of kl vatts/cc at the cuter edge of the ccre rather thsn 00 10 3-17
an estimated average of 28 vatts/cc over life, resulting in a safety g 1F =argin of about h5 percent.
- b. A =axi=u= axial power peaking factor of 1 7 was used. This is about 30 percent greater than the 1.3 expected over life.
Uncertainties in the. calculations include the folleving:
- 1. The use of only four neutron groups to describe the neutron energy spectru=. .
^2. Use of the LIFE code to generate the h-group cress sections. In the LIFE progra=, the b-group data in all regions are ec=puted fro = a fission spectru= rather than a leakage spectru=.
3 Having only four intervals, ie, n = h in the S calculation, to describe the angular segnentation of the flux." It is expected that the cc=bineticn of 1 and 2 above vill conservatively predict a high fast neutron flux at the vessel vall because it underesti-
=ates the effectiveness of the ther=al shield in reducing the fast flux.
In penetration through water, the average energy of the neutrons in the group above 1 MeV increases above that of a fission spectru=, ie, the spec- - tru= in this group hardens. For neutrens above.1 MeV, the nonelastic cress
- section of iron increases rapidly with energy. Therefore, the assu=ption c
of a-fission spectru to ec=pute cross secticns in the ther=al shield, and i y the use of a-fev-group model to cover the neutron energy spectru=, would [
~~
undere'sti=' ate the neutron energy loss in the ther=al shield and the sub- ! sequent attenuation by the water between the vessel and ther=al shield. !. The results- frc= 3k-group P3MGl(3) calculations show that reduction of [ the. flux above 1 MeV by the thermal . shield is about a factor of k greater j , than that computed fro = the h-group calculations. 1 j> The .effect= of 3 above is expected to underestt:ste the flux at the vessel E vall. In calculations at ORNL using the S technique, n a ec=parisen be-3 tween an Sh and an S1 2 calculation was made in a penetration through hydre-- gen. The result's for a variety of energies over a penetratien range of ILO ] !. c= showed the--Sh calculation to be 1cuer than the S12 by about a factor of j, 2 at maxi =u=. Good agreement was obtained between the S12 and =c=ents
~
=ethod calculations. ,
The above uncertainties indicate that the calculation technique shculd cver-esti= ate the fast flux at the reactor vessel vall. Ecvever, the cc=parisen
~
vith thermal flux data indicates a possible underesti= ate. Until a better co=parison with data can be made, we have assured that the underesttrate is ccrrect and accordingly have increased the flux calculations by a facter of 2 to' predict the nyt' in the reactor. vessel. The. reactor:utilices a larger water gap and thinner ther=al shield between uthe core and the reactor vessel vall when ec= pared to currently licensed
.[ The effect of this steel-vater configuratien en (a) the neutren 7 . plants.
r gh irradiation, and (b)~the ther=al stresses-in the reacter vessel vall, were evaluated as follevs: h c-18 00385'-
.I
' g.
-mm-,------ --.n, _ , _ . _ _ , _ , , _ . , , , ,
= -
- a. Neutron Irradiation Calculations were performed in connection with the reactor ves-sel design to deter =ine the relativt effects of varying the baffle and ther=al shield thicknesses on the neutron flux (>l MeV) at the vessel vall. These a culations were perfer=ed with the F1 option of the F3MG1 code 3 usins 3h fast neutron groups. The results shoved that the neutron flux level at the vessel vall is dependent, for the cost part, en the total metal and water thick-ness between the core and the vessel. Ecvever, there was scre variation in fluxes depending upon the particular configuratien of steel-vater laminations. Also, the Eain in neutron attenua-tien by replacing vater with steel dininishes screvhat with in-creasing steel thickness.
In general, hcVever, the result.m shoved that for total steel thicknesses in the ran6e of 3 to 6 in., 1 in. of steel in place c f l in of water vould reduce the neutrcn flux abcve 1 MeV by about 30 percent. In pure water, the calculations shewed that the neutron flux would be reduced, en the averaEe, by a facter of 6 in-6 in. of water. i Based on the above analysis, a ec=parison has been sade of the neutron attenuation in this reacter vessel with these in San 4 .. . Oncfre, Turkey Point 3 and h, Indian Point 2, and Ginna. The total distance between this core and the reactor vessel is 21
~~
k 5' in. This provides- frc= 1.5 to as =uch as 5.75 in. =cre distance between the core and the vessel than in the other reactors. Fcr i i neutrons above 1 MeV, it was found that this additional distance {_ vould provide additional attenuation ranging frc= a factor of
; 1.1 to 5 ti=es greater than that in the cther FW3 considered.
I
- b. Thermal Stresses.
The ga--a heating in the reactor vessel is produced by pri=ary
~
{ [ ga--as frc= the core and by secondary Sa==as originating in the core baffle plate, tarrel, ther=al shield, and the vessel itself.
- In this reactor design, the major portion cf the heat is gener .
< ated by ga- a rays frc= the core and by recendary ga==a rays frc=
the core baffle plate and barrel. Since' the ga- as -frc= each of these sources =ust penetrate the thernal- shield to reach the . vessel, the vessel heating rate is
-f dependent on the thernal shield. thickness.
I E ,
.For designs which employLthicker therral shields,.cr in which
,i: internals are to be exposed 1to higher neutron fluxes, gn--e rays t originating in the ther=al shield or in .the ~ vessel itself =ay govern the vessel heating rates. Since Ez= a rays frc= these
~-
sources would have to penetrate only portic=s or ncne ofLthe ther=al shield to . reach the vessel, the' vessel heating in such 5 . b') cases' vould be less dependent en ther=al shield thickne:s than-in this reacter cesign.
=
1 i _
'3-19
y ' l 1 4 A comparison was made between the garma attenuation provided by the vater and metal in this reactor vessel and that in other FWR by as-suming that, in each design, the vessel heating was dependent on the ga==a ray attenuation provided by the ther=al shield. This approach j 1vould be conservative since, as noted above for se=e designs, ganza j sources other than those attenuated by the ther=al shield nay con-1 tribute appreciably to the vessel heating. The results of the ec=- parison showed that the difference in ga :a attenuation between this reactor and other FWR ranged from a negligible difference to a factor of 5.3 less for this reactor design. The maximum steady-state stress resulting frcs garra heating in the vessel has been calculated to be 3,190 psi (tension). This is a rel-atively lov value, and no proble=s are anticipated frcm ther=al stresses in the reactor vessel vall. 1 j 3.2.2.1 Nuclear Evaluation Analytical models and the application of these models are discussed in this section. Core instabilities associated with xenen oscillation are also de-scribed, with threshold data evaluated under reference conditions. 3.2.2.2.1 Analytical Models
.[ Reactor design calculations are cade with a large number cf ec=puter codes.
"* fj The choice of which code set or sets to use depends on which phase of the
, {J design is being analyzed. A list of codes used in core analysis with a brief i
discussion follows in 3.2.2.2.2.
< a. Reactivity Calculations Calculation of the reactivity of a pressurized water reactor core is performed in ene, two, or three dimensions. The gec=etric choice depends on the type of calculations to be made. In a clean type of calculation'where there are no strong, localized absorbers of a type. differing from the rest of the lattice,1-di=ensional analysis is satisfactory. This type of proble= is handled quite vell by the B&W 1-di=ensional depletion package code LIFE. LIFE j is a composite of MUPT -(Ref h), KATE (Ref 5), RIP, WANDA (Ref 6),
i and a depletion routine. Nor: ally, the KUF portion is used with 1 - 3L-energy groups, an exact treatment of hydrogen, the Greuling-I Goertzel approximation for ele =ents of mass less than 10, and i Fer=1 age for all heavier elements. The KATE portion nor: ally uses a Wigner-Wilkins spectrum. In WANDA, h-energy groups are utilized. Disadvantage factors for input to the ther=al group are calculated with the THERICS (Ref 7) code. This code set has been shewn to give reliable results for a reactivity calculation of this type. Recent check calculations on critical experizents have a standard deviation of less-than 0.5 percent Ak/h. r A 1-dimensicnal analysis of a'geo=etric arrangement, where ther? are
, -( ) . localized strong absorbers such as CRA, requires a preliminary i-
.p dimensicnal analysis. The required prcperties of the 1-dirensicnal 00ES7
,f 3-20 ji
)=j j V'
syste= are then =atched to the 2-dimensional analysis. In this =an-ner, it is possible to analyze the simpler 1-dirensional syste: in a depletion survey proble= with only a c=all loss in accuracy. The 1-dimensional calculations are used as preli=inary guides for the more detailed 2-dimensional analysis that follows. Values of reactivity coefficients, fuel cycle enrichments, lifetimes, and soluble poisen concentrations can be found to improve the intial conditions specified for 2-dimensional analysis. Two-di=ensional reactivity calculations are done with either the PDQ (Ref 8) or TURBO (Ref 9) diffusion and/or depletion cedes. These codes have =esh limitations en the size of a ecnfiguration which can be shown explicitly and an often studied with quarter core sy:=etry. S;. zetry is desirable in the design, and no less in generality occurs. The geometric description includes esch fuel assembly and as =uch detail ac is possible, ie, usually each unit in the fue' ===a-bly. Analysis of this type per.its detailed pcver distribution studies as well as reactivity analysis. The power distribution in a large.PWR core which has cne loading cannot be predicted reliably with 1 .1:ensional calculations. This is par-ticularly true when local po"er peaking as a function of power history is of interest. It~is necessary to study this type of proble= with at least a 2-dimensional code, and in some cases, <, 3-dimensicnal calculations are necessary. Use of the 2-dinensional programs requires the generation of grcup constants as a function of =aterial conpocition, pcuer history, and gec=etry. For regions where diffusion theory is valid, MUPI',and KATE vith TiiERMOS disadvantage factors are used to generate epi-ther:al and ther:al coefficients. This vould apply at a distance of a few mean paths fro: boundaries or discontinuities in the fuel rod lattice. Discentinuities refer to water channels, instrumenta-I tion ports, and CEA guide tubes. The interfaces between regions of different enrichment are considered to be beundaries as well as the outer ~ limit of the core. j To generate coefficients for regions where diffusien theory is i inappropriate several methods are utilized. Tne arrange =_nt of
.j structural- caterial, water chtnnels and adjacent fuel rod rows can
; be represented well,in slab geometry. Tne coefficients so gen-i erated are utill:ei in the epitnereal energy ran6e. Coefficients
; for the thermal energy range are generated by a slab TiiEmiOS cal-culation. The regions adjacent to an interface of =aterial of different enrich:ent are also well represented Vith the P3'O code.
-Tne arrangement' of instrumentation ports and control rod guide tubes lends itself to cylindrical geonetry. DTF-IV (Ref 10) is quite effective in the analysis of this arrangement. Input to
'" Iteration is ITIF-IV is fro:- GdM (Ref 11) and TiiE:MOS or F. ATE. [.. required betteen the codes. Tne ' lux shape is~ calculated by LTF-
-IV and cross ~ sections by the others. .Tne cuter boundary of the r 1. - 2
~
core where there is a tra-e"'~ '- '"el ic reflector ard tsff'e
.; is also - represented by the l'~E-IV c:de. Tne 3-dirensicnsi anslysis 1
j
+
- i. 00 %
21-t
T \ t . , i - 1 4 l is acccmplished by extending the techniques of 2-dimensional rep-I resentation. 1
- b. Centrol Red Analysis _
B&W has developed a procedure for analyzing the reactivity worth of i small Ag-In-Cd rods in fuel lattices. Verification of this procedure vascadebytheccparativeanalysisoflhc-itic{}expericentswith varying rod and rod assembly configurations.ll3,1 8 Critical lattice secretries were similar to those of the reference core design. Ecrcn concentration ranged frc= 1,000 to 1,500 pps. The Ag-In-Cd rods were arranged in various gecretrical configuraticns which bracket the ref-l erence design. Water holes, simulating withdrawn rods , vere included as part of the lattice study. The resulting compariscn of the ana-lytical and experimental verths are shown in Table 3-8. Details of
- i. the critical configurations are given in References 13 and lb.
( Table 3-8 f Calculated and Exne-izental Rod and Rod Assembly Cc narison I Ag-In-Cd b." ole ~e Rod Assen? s -
- Ecd Assembly -
Core Asse blies Rods Per F'20 Calculated Experimental Uc. Per Core Asse:bly Per Core Werth, 7 ok/k Worth, 5 dk/k t 5-3 L h 252 2.00 1 9S
- L-F h 9 0 3.35 3.3h j 5' 2 12 276 2.33 2.35 h-D 1 16 0 1.L3 1.k2
. 5-D 2 16 26h 2.60 2.82
- , L-E 1 20 0 1 5L 1.52
.' 2t 5-E 2 20 292 3.05 3.01 The mean error in calculating these configurations is shown to be less (} than 1 percent. Cc parison of the power shape associated with the 16-Il red reference asse blies showed good similarity. Pcint-to-average power had a maxinun variation of less than 2 percent with experimental data. The analytical methed used for this analysis is based en straight dif-fusien theory. Thernal coefficients fcr a centrol red are citained frc THERMOS by flux-veighting. Epithermal ccefficients fcr the u,;per energy grcups are generated by the ELW LIFE prcgram. The resulting coeffi '.ents are used in the 2-dirensional cede FDQ to cbtain the re-quired eigenvalues.
~
'i GAKER and LISPM are used to prepare data for THERMOS. GAKI? generates scattering cross sections for hydrogen by the Selkin technique. LIEFM uses the Erown and S- Jchn free gas redel for generating the remain-ing scattering crcss sections. A c;/. THERMOS is urec 'r ivo steps. First, the critical fuel cell is ana-lyzed to cttain a velccity-veighted disaivant;ge fac cr. :.i r ic :rd 00!.S9
. a me w
i i I .i
# .g in the ho=ogenization of fuel cells and gives a first order cor-zy rection for spatial and spectral variation. Tne ratio of flux in the moderator to faux fc the fuel was analyzed to vithin 2 percent of experimental values using the velocity-veighting technique. The j second step is to use THERMOS in a calculation where the Ag-In-Cd rod is surrcunded by fuel. This is used to generate the flux- 'j veighted co,ntrol rod cell coefficients as a function of aron con-centration
- As a check on the validity of the Th? "0E approach,
=ethod.\l5extrapo}at on distances were cc= pared to those given by the Spinks The agreement was within 2.2 percent for a set of cases wherein the pu=ber densities of Ag-In-Cd were varied in a range up to 250 percent. All other coefficients are generated by LIFE in =uch the same canner as with THERMOS. The data are used in a 2-dimensional PDQ 1ayout where each fuel rod cell is shown separately.
I
- c. Deter =ination of Reactivity Coefficients This type of calevlation is different fro = the reactivity analysis only in applicati sn, ie, a series of reactivity calculations being required.
Coefficients are determined for moderator te=perature, voiding, and pressure, and for fuel te=perature. These are cal-g j l culated frc= s=all perturbat uns in the required parameter over the range of possible values of t..: para =eter. The roderator te=perature coefficient is deter =ined as a function of soluble poison concentration and =oderater te=perature, and i fuel te=perature or Doppler coefficient as a function of fuel te=perature. U The coefficient for voiding is calculated by vary-ing the moderator concentration or percent void. 1 3.2.2.2.2 Codes for Reactor Calculations . l This section contains a brief description of codes centioned in the preceding sections. j THERMOS (Ref 7) - This code solves the integral for= of the Eolta inn Transport Equation for the neutron spectrum as a functicn of position. A diagonalized connection to the isotropic transfer
= atrix has been incorporated allowing a degree of anisotropic scattering.
FUF2 .(Ref h) - This program solves the P1 or El multigroup equation for the first two Legend e coefficients of the directional neutron I flux, and for the isotropic and anisotropic co=ponents of the
- j. sloving down densities due to a cosine-shaped neutron source.
Coefficients are generated with MUPT for the epither=al energy l range. YJ.TE (Ref 5) - The code solves the Wigner-Wilkins differential equatien for a hc=cgeneous mediu= moderated by chemically unbound hydro-7 - gen ato=s in thermal equilibriu=. . Ccefficients for the ther=al energy range are generated by KATE. c 3-23 00190
t
, RIP - This pregra= averages cross sections over an arbitrary group i structure, calculates resonance integrals for a set of re-sol >ed peaks, and cc=putes L-factors for ing to KJFT, Pl!G, and FN:G.
;. ..WANDA (Ref 6) - This code provides nunerical solutions of the 1-dicensional
, fev-Grcup neutron diffusien equations.
1 LIFE - This is a 1-dimensional depletien package code which is a ec - bination of MUFT, EATE, RIP, and WANDA. The ec bination tech- {g. anizes the procedures for using the codes separately, i
.- GAM (Ref 11) -This code is a =ultigrcup coefficient generaticn pro-I Gra= that solves the F1 equaticns and includes anisctrepic 4
{ scattering. Inelastic scattering and resenance parameters ! are also treated'by GAM. 1, P3MG (Ref 3) - The code solves the =ultienergy transport equatien in ' i various gaccetries. The cod 2 is primarily used for epithermal , 11 coefficient generations. r ,;{ ~ DTF .(Ref 10) - Th' i s code solves the =ultigroup,1 -dimensional Ecltnrann p{ . . transport-equatien.by the method of discrete crdinates. DIF ! allevs cultigrcup anisctrepic scattering as well as up and dern _ s :,
. scattering.
- , PDQ (Ref 6) - This pregra solves the 2-dimensicnal neutren diffusicn-depletion proble
- with up to five groups. It has a flexible ,
, representation of ti=e-dependent cross sections by means ,cf fit optiens. 1 TURBO (Ref 9) - This code is. si=ilar in application to the PDQ deple-j tion progrs=. It, hcvever, lacks the great flexibility cf the f FDQ fit ~cptiens. I
- 1. ! CANDLE (Ref 9) - This code is similar to: TURBO, but selves' the diffusien C-
- equations in ene di=ension.
TET iRef 9) - This code is si=ilar in application to TURSO, but . is a ; , 3-di=ensional cede extended frc: ERAC 3.2.2.2.3 -Xenon Stability Analysis Initial studies of the reference ccre indicate that underda ped xenen escil-1; .lation shculd be considered. Features have been provided in the design to - [j allev centrol of axial oscillations and to take the ccre stable against azi-
'!- '=uthal oscillaticns. _(Radial oscillations are.unlikely.) These features are.
!I . discussed belev. o Asito axial' oscillaficns,- c'ertain "cf the centrol red assenb1' es vill centain u-
"~
poisen' enly in a portien of their lengths. They vill be pcsitiened.during the cperaticn cf the unit-to =aintain an a'eceptable distributien cf ';cver fcr any ,L 1 00?.9_1-
, .-c,
~g particular operating condition in the core, thereby reducing the tendency F for axial osciliations. Instructions will be =ade available to the operator concerning positioning of the partially poisoned centrol reds to achieve de-sirable axial power shapes.
In regard to azi=uthal oscillations, fixed poisen in the for= of burnable poison vill be used ir, fuel asse=blies (see Figu; e 3 h) as necessary to in-sure azi=uthal stability in the core for the design power density. A brief description of the studies performed to date follows:
- a. Method (1) Modal Analysis The details of the =ethods and init.a1 results of the
=cdal analysis are described in Appendix 3-A. ?
s
'(2) One- and P.o-Grout Methods and Modal Counling i One- and two-group treat =ents have been ec= pared. A one-group model is satisfactory for large, vnter-=oderated, lov 1 leakage cores such as the reference design. The effects of
- d. =odal coupling have been examined and shown to be of no ecn-sequence for cores similar to the reference design. Values 4 _ of critical. parameters varied no =cre than 1 to 2.8 percent
;s_h for the sane core with and without =cdal coupling. The lover i value was co=puted with a zero-power coefficient and was not f conservative without =odal coupling. The higher value was
- 4. . computed with the reference power coefficient and was con-j servative without =odal coupling.
3 i (3) Digital Anelysis G Xenon stability studies are continuing with codes in various d geometries which have the=nal-nuclear iteratien capability for both fuel. and_ :oderator te=perature feedback. The ene-F di=ensional feedback code LIFE handles the iteration in the i following =anner: , i i n
- " cut LTg = (T out - in i I (}Z ( }
a ,, i' in 4 Af - H where' AT = tenperature change in regien "i" l-1 PD(Z) = power density in the Z direction !- - . N;/
.N Zin, 2 cut = region "i" boundaries 0042 sr., f3Cp
's A i 3-25
? , and AT core c=H (_)
2 J PD(Z)df o I j vhere E is the active fuel height. Equation (A) is solved to Tout of region "i." Since Tin is kncvn frc= core inlet conditions, the average fluid temperature j is defined as fcllev;:
- m. + m.
out in . 2 T,..,,, 2...
=
2 1 (c) -1 1 The newly ec puted, regicn-averaged fluid terperatures are used to cc pute new fluid densities. These fluid densities are then used to adjust the number densities for water and soluble poison. Local or bulk boiling is nct permitted in the codel, but incorporation vould increase stability. The j , average fuel temperature for each axial ccre region is then cc puted frcs the average fluid terperature and power den-i sities as follevs: i i . ,3 i 'a - T fuel.
= f(PDi ) + T fluid.
1 1 s
! where FD. is the average pcVer density of region
.i i a
, ni,n an-i f(FEI ) is a tatular function relating fuel ter-4
^
i perature increase and pcVer density ob-tained frc auxiliary calculations. After the new fluid terperatures, nederater densities, and fuel temperatures are obtained, these quantities are used as new LIFE l input to obtain a new power distributica until a convergence
- i. criterion is ret.
4 This analysis used an exact solutien in that the spectru vas recalculated for each scne (11 axial zcnes described the reac-tor) for each iteration at every tire step. This included the effects of the coderater terperature coefficient. v.b This Lir pschage was used to deter:ine the effects cf the un-certainty in the pcVer Dc;pler cr thu st 't i- . cf tne cere-()() 9I:'
"[ 1 21 o The uncertainty in the Doppler was more than commensated with 1 a reduction in fuel te=perature of 500 degrees. 1 ,e reference core was analyzed with core average fuel temperatures of 1,h00 jf F and 900 F. Figure 3-6 compt s the cyclic response of these !i two cases following the 3-ft insertion and removal (after two hours) of a 1.2 percent Ak/k rod bank near the beginning of life. These studies were made at beginning-of-life boron lev-1 3 els of approxi=ately 1,900 pps. This level is apprcxi=ately l} 800 ppm above the predicted beginning-of-life level and, con-sequently, reflects a positive moderator temperature coefficient which is not expected.
- 4 Case 1 on Figure 3-6 depicts the behavior of the core if the f heat transfer equations were not included in the calculation.
, Figure 3-7 shows the effect of fuel temperature toward the end
- of life. It is easily verified thr.t the 900 F fuel terperature
. case approach'.i the threshold condition for axial oscillation I in this core. On the basis of the infor:ation presented, it
'I can be said that for a realistic fuel temperature this core does not exhibit axial in:tability at any tire during the ini-tial cycle. .
~! The one-ditensional model was used to deter =ine a rethod of con-
~
3 {' trolling the ccre without taking into account the stabilizing i [. effect of the power Doppler. I!or= ally, this vould produce a divergent oscillatien as shcun in Figure 3-8. A study was ec - pleted wherein a 1 percent Ak/k red bank with a 3-ft-long see-
.i tion of regular control rod material was successfully =aneu-vered to control the core after a perturbation of the power shape at a point about 3/k of the vay through Cycle 1. The
{' controlled results are also shown in Figure 3-8. The =ini=u
- j. rod motion was 1 foot, and the time step employed was L.8 hours.
l . More precise rod movement over shorter time periods would pro- , , duce a much smoother power ratio curve. This control =echanism l- appears to be quite adequate. I cr "X-Y" geometry is studied with the Stability)in HARMOIiY(6"R-Z" code, which in either case can be used with fuel and =oderator te=perature feedback. This code is used with fitted coefficients to obtain a more ecsplete solutien to the perturbed behavior of the reference design. The-digital results are processed by first fitting pcVer dis-y tribution results to an equation of the following for= by a least square technique: 1 2 where P, = Ae:t sin 2rt T P' = excess power
- = stability iniex i T = cscillation period
, 00l%
1 a 3. i . . f I , The calculated stability index is then extrapolated to zero-
. n/
.5 ~
length time steps by the procedures of References 20 and 21
, as follovs:
l l ' T - T
,=Z (1 - e #) -
i 1+e r
- 2 (1 -T e #)
o T 2 7 _ 4 T = <A> T t.. r
<A> = A +c4xo x
. i. j. Where Z = extrapolated stability index
. T = time step length r
& = average thermal flux 4
Y 7 Reasonable agreement has been found between the initial =odal g_f . analysis and one-dimensional results. 11
. , , (h)~-Control of an-Axial Oscillation The one-di=ensional model was used to determine a method of con-trolling the axial power distribution without considering the stabilizing : effects lof thenral-nuclear feedback. Ucr= ally,.this i . calculation'vould show a divergent esci11ation. A study was co=pleted'in which a 3-ft section of control red poison =aterial was successfully maneuvered to control the core after introduc-tion of a perturbation. The minimum rod motion was 12 in.-vith a h.8-hour mini =u= period for rod move =ent. More p2ecise =cve-cent over a shorter mini =um period'vould improve centrol. .The procedure vill be extended to two di=ensions.
- b. Conclusions j: The following conclusions have been made as a result of the studies J to date:
i
, (1) ' Instability. in the radial direction .will 'not occur.
(2)l'The core design under exa=ination vill not be susceptible to '
~
diverging azi=uthal oscillaticns. m-E-)'.- , j 00495 , 3-at
(3) Potential axial oscillations will be thwarted or centro 11ed by the part length control rods. The nominal and minits: (acccunting for the ur. certainties ccrree-tion listed above) stability indices are described in the Appendix 3-A. 3.2.3 THE RAL A'D HYDRAULIC DESIGN A'O EVALUATIO:; 3.2.3.1 Thermal and Hydraulic Characteristics 3.2.3.1.1 Puel Assembly Heat Transfer Design
- a. Design Criteria The criterion for the heat transfer design is to be safely telev Ocparture From :,ucleate Eciling (D::E) at the design cverpcver (11-percent of rated pcver). The analysis is described in detail in 3.2.3.2.2, Statistical Core Design Technique.
)
e I e s-
' e.
3-29 L
s
. The input information for the statistical core design technique and h for the evaluaticn of individual het channels is as fellcvs:
(1) Heat transfer critical heat flux equaticns and data correlations. (2) Nuclear pcVer facters. (3) Engineering het channel facters. (h) Core flov distribution hot channel facters. (5) Maxicus reactor overpower. These inputs have been derived frc: test data, physical reasurerents, and calculatienc as outlined belev.
- b. Heat Transfer Ecuaticn and Data Correlation The heat transfer relationship used to predict limiting heat trans-fer conditiens is presented in Eeferences 22 and 23 The equations are as follows:
1 (1) W-3 unifor flux DNE ccrrelation for single channel with all valls heated: 4 NB eu
/ g = { (2.022 - 0.000h302 P) i 10
+ (0.1722 - 0.000098h P) exp [(18.177 - 0.00L129 F)x]}
.i U
t x[0.lh8h-1596x+0.1729xlxl c + 1.037) 4 10 I x [1.157 - 0.869 x) x [0.266L + 0.8357 exo (-3.151 Le )] x [0.8256 + 0.00078h(E sat - E. in)] where Q" = flux, Etu/h-ft
! P = pressure, psia G = cass velocity, lb/h-ft X = quality, expressed as fracticn D = ecuivalent diareter, in.
., e -
U H = enthalpy, Etu/lb 0 0 ~ ~ "'/
,n
- ~ J
s a (2) W-3 nenunifor: flux D:iB correlatien fer cincle channel with all ' valls heated:(22) iI nn _ c ., p*
%e * %e s. ,
&. 9.g.,e*
a &a ,eu I Where = L:;3 heat flux fcr the ncnunifcr:1v Q".,.,,:'? u ~ heated channel Q",, Dam ,eu = equivalent unifer: ":;5 flux r = C (
\" Q loc a.,
il - exp l- i_.._ )J/ c..:,eu I
, s ,
L ,.-
&s:
,, n x- ,
I , (n) exp [
& n(E~7 ",,,.
u
- n. ) J' 6c. - )
0 ~ w (, - x, u..
)I.o, C = 0.w,, in.
t G 1 1.72 l t
,10 1
D (3) W-3 unifo= flux D:rs correlation fer single channel with un-l heated valls:(23) IIe, with unheated vall
= ( 1. . ,:t + 0.lc_ e9 x)
Q,iB, usire n to replace D n e i
-1.Cc ?y l x (1.2 - 1.6 e ")
i c 4A x (1.33 - C.237 e' *) where De = ecuivalent dic=eter tased en all the vetted periteter, in. D = equivalent dic.neter tased cn only the b' heated perimeter, in. i Individual channels are analyzed te dete=ine a :;3 ratic, it , the j ratio of the heat flux at which a 2:;3 is predicted to cccu- to the
, heat flux in the channel teing investigated. 3is ":;E ratic is re-lated to the dnta errrelttirn :: rhrw. ir Figure 3 c. A ccrfid: ee a=d pcpm1auc= vel =e 1c amc1=ed v1:h eve.., 35 rule u d m .-ihed l
0 9t ; ' W.. s h S *4
i l e (*E b
*k in the Statistical Core Design Technique (3.2.3.2.2). The plet of DN3 versus P is for a confidence of 99 percent. The criterien for evaluating the thermal design targin for individual channels or the +
total core is the confidence-population relaticnship. The DN3 ra-l tics required to =eet the basic criteria or limits are a function ! of the experimental data and heat transfer correlation used, and j vary with the quantity and quality of data. The recentended minitur design DN3 ratio for the W-3 correlation is 130. The DN3 and population relationship fer a design limit of 1.30 in the hot channel corresponds to a 99 percent confidence that at least 9h.3 percent of the population of all such hot channels is in no jeopardy of experiencing a DN3. The DN3 rstics and the frac-tien of the core in no jeopardy of experiencing a LN3 at design conditions are censiderably higher than those given in the design limits outlined in 3 1.2.3 I
- c. Nuclear Pcver Factors The heated surfaces in every flow channel in the core are examined for heat flux limits. The heat input to the fuel rods in a coolant channel is determined frc= a nuclear analysis of the core and fuel assemblies. The results of this analysis are as follevs:
y
'(1) The nerinal nuclear peaking factors for the vorst time in ccre g life are Fan = 1 77 9
F: = 1.70 { Fq = 3.01 (2) The design nuclear peaking factors for the worst time in core j life are 1 Fah = 1.78 F: = 1.70 i l Fq = 3.03 i where Fah = cax/ avg total pcver ratio (radial x local nuclear) e F: = =ax/ avg axial power ratio (nuclear)
" Fq = Fah x F: (nuclear :ctal)
J
}
QQ* M
i l
! o . . Q,. The nc=inal values are the maxi =u= values calculated with nc=inal !
F spacing of fuel asse=blies. The design values are obtained by ex-a=ining taxi =u=, nc=inal, and =ini=u= fuel asse=bly spacing and de-
, termining the vorst values for the co=bined effect of flov and rod ;
peaking. f I t The axial nuclear factor, F:, is illustrated in Figure 3-10. The distribution of power expressed as P/F is shcvn for two conditions of reactor operation. The first condition is an inlet peak with a
=ax/ avg value of 170 resulting frc= partial insertion of a CRA group for transient control following a power level change. This condition results in the =axi=u= local heat flux and =axi=u linear heat rate. . The second power shape is a sy==etrical ecsine which is indicative of the pcVer distribution with xenen everride rods with-I dravn. The flux peak =ax/ avg value is 1 50 in the center of the active core. Ecth of these flux shapes have been evaluated for i ther=al DN3 lt=itations. The li=iting conditien is the 1.50 cosine power distribution. The inlet peak shape has a larger =axi=u= value.
EcVever, the position of the 150 cosine peak farther up the channel results in a less favorable flux to enthalpy relationship. This effect shapes.gg)beende=enstratedinDN3testsofnenunifer= The 1.50 cosine axial shape has been used toflux deter =ine individual channel DNB li=its and to make the associated statistical analysis. u a The nuclear factor for total radial x local roi pcver, Fdh, is calcu-u lated for each rod in the core. A distributicn curve of the fraction ,
). -
of the core fuel rods operating above varicus peaking factors is shown !
]* in Figure 3-11 for a typical. fuel cycle condition with the =axi=u:
fuel rod peaking factor of 1 78.
- d. Engineering. Hot Channel Factors t Fover peaking factors obtained frc= the nuclear analysis are based on jj mechanically perfect fuel asse=blies. Engineering hot channel facters
!!' are used to describe variations in fuel loading, fuel and clad di=en- {i siens,'and flow channel gec=etry frc= perfect physical quantities and
- j di=ensions.
The application of hot channel factors is described in detail in 3.2 3.2.2, Statistical Core Design Technique. The factors are de- !: termined statistically frc= fuel asse=bly as-built or specified data lj vhere Fq is a heat input factor Fqn is a local hect flux factor at a
, hot' spot, and.FA is a flov' area reduction factor describing the varia- ,
. tion in coolant chhnnel-flow area. Several subfactors are. cc=bined l lL .
l'
- statistically to obtain the final values for FQ, F ",Q and F .AThese
- , subfactors are shown in Tsble 3-9. 'The factor, the coefficient of i variatien, the standard deviation, and the =ean value are tabulated.
L 4fp
' 00NO 3 L- . - - -- -- -..- -..
Table 3-9 Coefficients of Variatien CV lio. Descrirtion c i CV 1 Flev Area Interior Eundle Ce12s 0.00190 0.177ko 0.01072 Peripheral Bundle Cells 0.003L6 0.215L6 0.01608 2 Local Rod Dia:eter 0.0006L7 0.L30 0.00151 3 Average Red Dia eter 0.0006L7 0.L30 0.C0151 (Die-Drawn, Lecal and Average Sate) k Lccal Fuel Leading 0.00695 Subdensity 0.006L7 0 935 0.C0692 Subfuel Area 0.00009L 0.1075 0.0005S (Dia eter Effect) 5 Average Fuel Leading 0.00557 Subdensity 0.00L85 0.935 Sublength 0.00519 0.2629L ILL 0.00183 Subfuel Area 0.00009k 0.1075 0.00:55
) (Dia:eter Effect) 6 Lc a1 Enrich =ent 0.00h21 2.30 0.00183 7 Average Enrichment 0.00421 2.30 0.00183 CV Coefficient of Variation c/E (Enrichment values are for verst case o Standard Eeviation of Variable ncrtal assay batch; caxi==n variation i Mean Value of Variable occurs for minicus enrich:ent.)
- e. Cere Flav Distributien Het Channel Factors The physical arrangement of the reactor vessel internals and no :les results in a nenunifor= distributien of coolant f1cv to the varicus fuel asse blies. Reactor internal structures abcVe and telcv the active core are designed to tinimize unfavorable ficv distribution.
A 1/6 scale model test of the reactor and internals is being per-fctned to de=onstrate the adequacy of the internal arrange ents. The final variations in flow vill be dete nined when the tests are cct-pleted. Interi facters for flov distribution effects have been calculated frc test data en reacter vessel =cdels fer previcus pres-surized water reactor designs. A f1cv distribution facter is determined for each fuel assently loca-tien in the ccre. The factor is expressed at the ratic cf fuel as-j/ senti. flew to average fuel assenbly ficw. The finite cal'_t: cf the 00?.0?. 3-3L
' ,7 O ratio =ay be greater er less than 1.0 depending en the position cf V the asse=bly being evaluated. The f1cv in the central fuel asser-blies is in general larger than the flow in the cuternost assentlier due to the inherent flow characteristics of the reactor vessel.
, The flev distribution factor is related to a particular fuel asser-bly location and the quantity of heat being prcduced in the asse:bly.
A ficv-to-pcker comparison is made fer all of the fuel assemblies. The worst condition in the hottest fuel assembly is determined by applying model test isothermal flow distribution data and heat in-t put effects at pcver as outlined in 3.2.3.2.3. Two assumptfens for i flev distribution have been =ade in the therral analysis of the core
?
as fo11cvs: (1) For the taximu design condition and for the analysis cf the bettest channel, all fuel asse blies receive mini =un flev fer the verst conditien, regardless of assently pcVer er lecaticn. (2) Fcr the Ost probable design ccnditions predicted, average f1cvs have been assigned for each fuel assembly censistent with lo-
! cation and power. The flew facter assured for the taximu de-sign condition is censervative. Application of vessel f1cv test data and individual asse bly flew factors in the detailed core design vill result in i=preved statistical statements for the taxi =u: design ecndition.
h f. Maximum Reacter Design Overrever Core performance is assessed at the maximu= design overpower. The selection of the design overpower is based on an analysis of the re-actor protection syster as described in Section 7 The reactor trip l point is 107 5 percent rated power, and the taxinun overpower, which is 11L percent, vill not be exceeded under any conditions,
- g. Maxi =u: Design Cendtions Analysis Su =ary The Statistical Core Design Technicue described in 3.2.3.2.2 was used to analyze the reactor at the maxi =u: design ecnditiens de-scribed previcusly. The total number of fuel reds in the ccre that ha.ve a possibility of reaching DN3 is shown in Figure 3-12 for 100 to 130 percent overpever for the maximum design conditions. Foint 3 cn Line 1 is the taxi =u: design point for 11L percent pcver with the design Fah nuclear of 1.78 and minimu= flcw to every channel in the core. This Point B forms the basis for this statistical state-sent:
, There is a 99 percent confidence' that at least 99 98 percent of the fuel rods in the core are in no jeepardy cf experi-encing a departure frc: nucleate boiling (DN3) during ecn-tinuous cperation at the design everpower cf 11L percent.
j At 100 percent pcuer (2,L52 MWt) as shcvn by Point A, the ctatisti-cal number of fuel rods in jeopardy is less than ene, reru' ting in a 0020,?, 3-35
~3 O
population prctected in excess of 99 997 percent. The limit i:- populaticn prctected of 99.610 percent.peced by a V-3 EN3 rat An additional design anal.ysis conditions except cf fuel rods in jec r ardy fer all the taxi nus Line 2 in Figure 3-12. Each fuel assembly flev distributien is shewn as assembly was ace"
- flev for the assembly pcVer conditions. - *eceive average flev distributien vill result in values within the beunds lines.
Statistical results fer the taxi =u: vc of th lation,3-10. Table as shown by Figure 3-12, may be su= arized as fellevs indesign con Table 3-10 DN3 Results - Maxinu: Desien Cenditien (995 Ccnfidence Level) Fever, ?cpulation Het Channel _Foint y of 2,k52 MWt Fessible Prctected. Fah DN3 DN3 Eatic (W-3) A 100 3 1 7c <1 11L 1 78 >99 997 2.21 C 126 5.0 99 956 1.78 70.0 1.71 99.610 1 30 h. Mest Frebable Desien Cendition Analysis Surrary The previous maximu: 3-12, indicates the total number of rods that ray be in jecpa when it is conservatively assu=ed that every red in the cere has the
=echanical scribed in 3.2 and3.2.2.heat transfer characteristics of a het channel as de F
t$an('flev area factor) less than 1.0, FFcr ext =ple, all channels are a 1.0, and with rinicu: fuelasse:bby(heatinputfacter)creater flev. It is phyrically in-possible for all channels to have het channel characteristics. A tere realistic indicatien of the nu:ber of fuel rods in jecpard y may be cbtained by the application cf the statistical heat transfer data to average red pcver and techanical ecnditiens. An analysis for the rest pretable conditiens has been made based en the average conditiens described in 3.2.3.2.2. analysis are shcvn in Figure 3-13. The analysisThe results of thic may te sr--ericed as follevs in Table 3-11. e 00?.03
..,e.
a
, Table 3-11 DSB Eesults - Most Frobable Ccnditien Fepulatien Het Cnannel Fcver, ?cssible Frctected, DNE Estic Foint j.cf 2,L52 MWt Fdh D53 0 (W-3)
D 100 1.78 <1 >99.997 2.LS t E llh 1.76 1.7 99.995 1 97 F 130 1.78 9.9 99 973 1 51 The analysis was cade frc: Point D at 100 percent pcVer to Fcint F at 130 percent power to shcw the sensitivity cf the analysis with pcver. The verst ecnditien expected is indicated by Pcint I at 11L rercent mover
. r where it is shewn statistically that there is a snall possibility that 1.7 fuel rods cay te subject : a departure frc nucleate bciling (EUS). This result fcr:s the basis fer the fc11cv-ing statistical state ent for the test probable design ecnditions:
i. There is at least a 99 percent cenfidence that at least 99.995 percent of the reds in the core are in no jecpardy cf experiencing a LNE, even with centinucus cperatien at the design everpcver of 11L percent.
]) 1. Distributien cf the Fraction of Fuel Reds Frctecced ! The distributien of the fraction (?) of fuel rods that have been i
shown statistically to be in no jecpardy of a 253 has been calcu-lated for the taximu: design and = cst probable design ccnditiens. f The ec puter programs used previde an cutput cf (5) nu ber cf reds and (F) fraction of rods that vill net experience a D53 grcuped
' for ranges of (?). The results for the ecst pretable design con-diticn are shcvn in Figure 3-1L. .I The pcpulatien protected, (P), and the pcpulatien in jeopardy, (1-?),
are both "rletted. The integral of (1-?) and the nunter cf fuel rods oives the number of reds that are in jeepard.v fer given ccnditiens as shewn in Figures 3-12 and 3-13 The nuste- ' - 's is ettained frc: the product of the percentage times the total nutter of reds being censidered (36,616). Tvc typical distributiens shcvn in Figure 3-1L are for the rest prciable ccnditicn analysis cf ?cints E and F cn Figure 3-13 The lever line of Figure 3-1L shcws P and (1-F) at the j 11L percent pcrer condition represented by Point I of Figure 3-13. The upper curve shcvs P and (1-P) at the 130 percent power cenditien j . represented by Point F cf Figure 3-13 The irtegral of n and (1-P) cf the Icver curve forms the basis for the statistical statcnent at the ecst pretable design ecndition described in (h) ateve.
- j. Ect Channel Performance Sn--*rv The hotte"* " '* cell with all curfaces h s*ti a: been to rined fcr hct c'rarnel facters, E55 ratics, and quality f;r a ranic cf r c cr e
p ., t l)e. s e'at e*q. ,
- si
powers. The cell has teen examined fer the taximum value cf Fth
's ~ ,.v . . e ,, ...e ;- . . . a n.- . e. ., e a-e.n r ..sv.
v. e..+.,. v. . c . c .. e ... . e ., . .o .ae _
,,.e e . , . n_ ,a .n :
c...., t .e e.. s. .:
, . . v.4.>... p, ,. ,
. ., - ... . . e. c .>, a_. ,e.,. . em o ,. c ,. e , 5. , _; c. y ,~. .
The heat generated in the fuel is 97.3 percent cf the tctal nuclear
*en*. *i..e .'e . .a'.*..e-
.. . . 2.7 3.-ce*.. a. .
Loe a _e e " a. ^- *o '.a. e e. ,, ..a. s*.e'. * ...
* *...a.
- e col ar.*. a.e. .". --ocee f. d. u p *.*.. a. k.. a~.. . a. .' . . . . . . . .**e e, . ".#**** . . .... a..d. d. . e .- a. .m #'a-ted as an increase in dT cf the ecolant.
Errcr hands cf 65 psi cperating pressure and :2 7 e.re reflected in
*
- e
- o*.al cm. e ar.e- . .k. c '.. e ~.. . a. '.
.k..a. . a l a e '. . c a .' "" ' a*. 4 c . s #. .. 9. .e d *..
- e c '....* - . ,. . w ^ " a- .'. e.e *. *.. e a' ". e s t .". ". a*.
- c s c .- k.. .' e-*. . e * *. ," = ' 4 1 e s .
.h e "., e..
- a*. . 4 .o ". e . .c "as f -. . e.-.4 s .o..~.~.. .d .. 7.'3,~ e ., 's . .
.c. m. .. a. -- " . . a*. 4 e .
d e.. +
. . v.. e k.. w" *. ^ k.. a".. ..e l
. a*. *. b..e " aX .
4
""" cV e "y ~.~ *. .* C.# .'.' h *ye .'"' *. .*. 4 e. '. 7.'
..-w.d .e. . k. . a.r. . .* e..e - ~. 2 s +o c a , c,0 - a.- r c. a -.'. C C .. #.4. da. .. . a. *. k.. a*. E*.
.' a. a .e . c,c.E2 ,
y g. r
- e. e. . .. . e . . .v... . c.,..,.. . . g w. a"...a. l s C .# *. b..#. s *.*.*f
. a*e .
.. .^,.,
. ,da^,7 -* .- '"j ^#
- v. eX-e g ,. e . w.
y . .a. . .r. 3. e. e g e.*. .;. . ". W. . e. .g. g .. 4 e..a. e. v.e. e. . e. b. . e . e %.
. . . . . . e , e. . a. , .eg . e. .
. e. uea.A 4,. . . v. . o.
d s . "g . 1, . " . "g a...'";*.*..'*.*...-
.A n .e .g.t-. g e.
n..g . e
, u c . e. g... A e .e. C. ' b. e . .#. . . 4 . . .
- 7. , - -
, 0, ,
- r. . n = ..o.L x
- r. , = o.ccs (....e 4-.
... c......,,e) a r.n
= c.ca ("a., , C e'..' e )
s
- The het channel exit quality fer varicus pcvers is shewn in Figure 3-16. The cerbined results may be sunnarized fer 2,120 psig as follevs:
F.eacter Fever, 5 DI3 F at i c (k'-3 ) Exit Ouality, %
.i 1
2.4.,
., . ov (a) lo0 v
.. .e .*.#...e,) 1.o'. '.5
',07
. v5 (. < , c 1. 71
. L (1._x 4 ...
- r. eve ) 2. . c, 125 1.30 9.1
,: ,, .c,
.. 1. C.3 .
a p Subeccled 3.2.3.1.2 Fuel and Cladding Ther=al Ccnditions
- a. Fuel e
' A di-ital t ec puter code is used to calculate the fuel terperature.
- n. . . a 3..z.
. ..ve- ...ea.e . .e. n. ..- .. . t e. c . emn 1.g. e. . , 2. . . e .yi e a v. _
- e. c . . e. . g..4. -v.e.
g_ n .e g e. . e... %. e
- r ft.n.,
e..:, e.e , n..a- e n. ..... - . e_ . , co,. e ... e c .. a.a
- . .; m- .. .e .a
- . . . . n. e .. - .-e..c.o ecefficients deterrined fer therral-hydrs"lic channel sclutiens.
g e.w
...e
. . ,. . u _, , soc w.. 2..
.. .. 4..:.... .- 4e
.s ". n .i e * '... a . w' ' '. a '.d'..<~~.#.~..
- - . e.
.e c......:-.
...... g.e ...s. n. *. ;. g_ , v. a .. . ... .:...~c. ;..Z.e.
. . . . . y. 2 . ,., , .eC
..,.,2..- . . . .
- e. g e 7: ..,.y ,_9'*
g v'..*.. pP f a." ..l.....,,.:. . . . ... W- , .4
..,p .'r. ... .e.*
., Z .b . :- ..
q
... n.
g,
= 6 . o r. . - g y V. e O n 4 .g*-
%.lyh.. -.
...,e .8*.
gu
versus te=perature. The heat transfer frc the fuel to the clad f I is calculated with a fuel and clad expansion model proportional to te=peratures. The te=perature drop is calculated using gas cen-ductivity at the beginning-of-life conditions when the gas cen-2 ductivity is 0.09 Etu-ft/h-F-ft . The Eas conduction model is used in the calculation until the fuel thernal expansion relative to the clad closes the gap to a dimension equivalent to a centact coefficient. The contact coefficient is dependent upon pressure and gas conductivity. A plot of fuel center te=perature versus linear heat rate in kW/ft is shown in Figure 3-18 for beginning-of-life conditions. The linear heat rate at the taximum overpover of 11b percent is 19.2 kW/ft. The corresponding center fuel te=perature shown in Table 1-2 is h,k50 F. The center and average te=peratures at 100 percent pcver are k,150 and 1,3h5 F as shown in Table 3-1. The peaking factors used in the calculatien are Fah = 1.78 F: = 1.70 F ., = 1.03 q F (nue and mech) = 3.12 A conservative value of 1.03 was assu=ed for the heat flux peaking i ~~ factor, Fqu. The assigned value corresponds to a 99 percent popu- ! 1ation-protected relationship as described in~the statistical tech-nique. l
- b. Clad The assu=ptions in the preceding paragraph vere applied in the cal-culation of the clad surface te perature at the maximu everpever.
Sciling conditions prevail at the hot spot, and the Jens and Lottes relationship (25) for the coolant-to-clad ST for boiling vas used to deter =ine the clad temperature. The resulting taximum calculated clad surface te=perature is 654 F at a system operating pressure cf 2.185 psig.
'3.2.3.2- Thermal and liydraulie Evaluation 4
c3.2.3.2.1 Introduction
' Sunnary results for the ' characteristics of 'the : reactor design are precented in 3.2.3.1. Tne Statistical Core Design Technique e=plcyed in the design repre-sents a refine ent in the cethods-for evaluating pressurized water reacters.
. Corresponding single het channel DNB data were presented to relate the nev 1), nethed with. previous criteria. A cc=prehensive description of the.r.ev tech-nique 'is included in 'this 'section to permit a rapid evalustien of the tc.hodt used.
002:06 3-39
7 k . i . A detailed evaluation and sensitivity analysis of the design has been =ade by ext =ining the hottest channel in the reactor for DNE ratio, quality, and 1 fuel te=peratures. The W-3 correlation has been used in th4c analysis. 1 1 3.2.3.2.2 Statistical Core Design Technique { The ccre ther=al design is based on a Statistical Cere Design Technique de-
; veloped by B&V. The technique offers =any substantial h:prevements over older =ethods, particularly in design approach, reliability of the result, and =athe=atical treatment af the calculation. The =ethod reflects the per-for:ance of the entire core in the resultant power rating and provides in-sight into the reliability of the calculation. This section discusses the technique in order to provide an understanding of its engineering = erit.
The statistical core design technique censiders all parameters that affect the safe and reliable operation of the reacter core. By considering each fuel rod, the =ethod rates the reactor en the basis of the perfor=ance of l the entire core. The result then vill provide a goed neasure of the core safety and reliability since the =ethod provides a statistical state =ent for the total core. This state =ent also reflects the conservatis: cr design =ar-gin in the calculation. A reactor safe operating power has always been deter =ined by the ability of the coolant to re=cve heat frc= the fuel =aterial. The criterion that best l =easures this ability is the DNB, which involves the individual parameters of heat flux, coolant te=perature rise, and f1cv area, and their inter-effects. The DN3 criterion is cen=enly applied thrcugh the use of the de- ,] parture fro: nucleate boiling ratio (DN3R). This is the =ini=u= ratio of j the DN3 heat flux (as ec=puted by the DN3 correlation) to the surface heat
'. - flux. The ratio is a =easure of the =argin between the operating power and l the power at which a DN3 =ight be expected to occur in that channel. The
- I DR3R varies over the channel length, and it is the =inimum value of the ratio in the channel of interest that,is used.
j The calculation of DNE heat flux involves the coolant enthalpy rise and cool-t ant flev rate. The coolant enthalpy rise is a function of both the heat in-put and the flev rate. It is possible to separate these two effects; the statistical het channel factors required are a heat input facter, 7;, and a flew area factor, F A. In addition, a statistical heat flux facter, FQ" is
- 4 required; the heat flux facter statistically describes the variatien in sur-face heat flux. The DR3R is = cst limiting when the burnout heat flux is based en =ini=u= f1cv area (s=all FA) and =ax4--- kaat input (large FQ), and when the surface heat flux is large (lan e F Q"). The DNS ccrrelation is pro-Pf vided in a best-fit for=, ie, a forn thm; st fits all of the data on which
- i. the ccrrelation is based. To afford protection against DNE, the DN3 heat
-f flux cc=guted by the best-fit correlation is divided by a DN3 factor (BF) l - greater 'than 1.0 to yield the design DNS surface heat flux. The basic re-i.
'laticnship
+
EbNB '
;g . 'DNER = x f(F , F ) x g q ..
BF surface Fo
;;j. Q L
00207 3 hD
'j involves as parameters statistical hot channel and DNE factors. The DNE fac-ter (EF) above is usually assigned a value of unity when repcrting DNE ra-tics co that the margin at a given condition is shown direct 1;. by a DNEE greater than 1.0, ie,1.30 in the hot channel. Selected heat transfer data are analyzed to ebtain a ccrreacticn. Since ther-
=al and hydraulic data generally are vell represented with a Gaussian (ncreal) distribution (Figure 3-19), rathematical paraceters that quantitatively rate the correlaticn can be easily cbtained for the histegram. These same rathe-catical parameters are the basis for the statistical burnout facter (EF).
In analyzing a reacter core, the statistical inferratien recuired to describe the het channel subfacters may be obtained free data en the as-built core, free data en similar cores that have been ecnstructed , er frc: the specified tclerances fer the preposed ccre. Eegardless of the source cf data, the sub-facters can be shewn graphically (Figures 3-20 and 3-21). All the plcts have the same characteristic shape whether they are fer subfac-ters, het channel factors, or turncut facter. The facter increases with either increasing population er ecnfidence. The value used for the statisti-cal het channel, and burnout facter is a function of the percentage of confi-dence desired in the result, and the pertion of all possibilities desired, as well as the accu =t of data used in determining the statistical facter. A frequently used assumption in statistical analyses is that the data avail-able represent an infinite sample of that data. The inplicaticns of this as-
) su ption should be noted. Fcr instance, if litited data are available, such an assumption leads to a screwhat cptimistic result. The assumpticn also implies that more information exists for a given sample than is indicated by the data; it i: plies 100 percent confidence in the end result. The Era' cal-culational procedure does not make this assumpticn, but rather uses the spec-Ified sa=ple size to yield a result that is such =cre meaningful and sta-tistically rigorous. The influence of the amount of data for instance can be illustrated easily as fo11cvs: Ccnsider the heat flux facter which has the fcr Fn = 1 + Ec Q Fn
%n where Fn r is the statistical het channel factor for heat flux K is a statistical cultiplying factor o is the standard deviation of the heat flux fac-F" tor, including the effects of all the subfacters Q
If cFqn = 0.05 fer 300 data points, then a F. factor of 2.605 is required to protect 99 percent of the pcpulation. The value of the hot channel factcr then is F. G Fn = 1 + (2.t05 x C.C50) = 1.120L c 00 m>c8 2u s a
u s I
~j.
and vill provide 99 percent confidence for the calculation. If, instead of
)
using the .300 data points, it is assu=ed that the data represent an infinite sa=ple, then the K factor for 99 percent of the population is 2.326. The value of the hot channel factor in this case is i 4 l F ,, = 1 + (2.326 x 0.050) = 1.1163 i A which i= plies 100 percent confidence in tha The values of the K factor used above are taken from SCR-607{2g iculation.
. The same basic techniques can be used to handle any situation involvin5 variable confidence, pcpulation, and nu=ber of points.
i Havin5 established statistical hot channel factors and statistical DNE fac-
.h tors, ve can proceed with the calculatien in the classical =anner. The sta-tistical'facters are used to deter =ine t*a '-d-"-
'"acticn of rede protected, or that are in no jeepardy of experiencing a DH3 at each nuclear power peak-ing factor. Since this. fraction is known, the =axi=um fraction in jeepardy is also.known. It should be reco5nized that every rod in the core has an associative DNB ratio that is substantially greater than 1.0, eyen at the de-si 5n overpower, and that theoretically no rod can have a statistical popu-
} , :.ation factor of 100 percent, no =atter hov large its DN3 ratio.
i Since both the fraction of rods in jeopardy at any particular nuclear pcwer
, :.Q peaking factor and the nu=ber-of rods operating at that peating factor are F known,' the total nu=ber of rods in jeopardy in the whole core can be obtained by si=ple sum-ation. The calculation is =ade as a function of power, and ,l the plot of rods in jeopardy versus reactor overoever is obtained (Figure 2
3-12). The su==ation of ae fraction of rods in jeopardy at each peaking factor su==ed over all peaking factors can be =ade in a statistically riger- [ ous =anner only if the confidence for all populations is identical. If an { infinite sa=ple is not assu=ed, the confidence varies with population. To for= this su==ation then, a conservative assumption is required. E&W total core =odel assu=es that the confidence for all rods is equal to that for the least-protected rod, ie, the =inimum possible confidence factor is associated with the entire calculation.
- The result of the foregoing technique, based on the =aximu= design conditions (llh percent pover), is this statistical state =ent:
There is at -least a. 99 percent confidence that at least 99.98 percent of the rods in the core are in no jeopardy of experiencing a INE, even with continuous operation at the design overpower.
- , The maxinun desi5 n conditions are represented by these assumptions
- a. The maxi =u desire. values :of FLh (nuclear =ax/ avg total fuel rod heat input) are obtained by exa=ining the =aximu=, nccir.al, and
=inimum fuel ast e=bly spacing and deter =ining the vorst value for red. peaking.
00209 3-42
f
^ b. The Lxi=u= value of F (nuclear =ax/ avg axial fuel rod heat in-put) is deter =ined for the li=iting transient or steady-state cen-j' dition.
- c. Every coolant channel in t'he core is assu=ed to have less than the no=inal flow area represented by engineering hot channel factors, F , less than 1.0.
A
- d. Every channel is assured to receive the minimum flov associated 11 vith core flev =aldistribution.
- e. Every fuel rod in the core is assumed to have a heat input greater than the =aximu= calculated value. This value is represented by
' engineering hot channel heat input facters, Fq and FQ", which are greater than 1.0
- f. Every channel and associated fuel rod has a heat transfer margin above the experi= ental best-fit li=its reflected in DNB ratics greater than 1.0 at =aximum overpower ecnditiens.
The statistical core design technique may also be used in a similar =anner to evaluate the entire core at the =ost n'robable =echanical and nuclear con- .; , ditions to give an indication of the most probable degree of fuel ele =ent jeopardy. The result of the technique based en the = cst rrobable design conditions leads to a statistical state =ent which is a corollary to the =ax-i=um design state =ent: i) " There- is at least a 99 percent confidence that at least 99.995 percent of the rods in the core are in ne jeopardy of experiencing a DN3,,even vith continuous operatien at the' design overpower. The most probable design conditions are assu=ed to be the same as the =axi-
=u: design conditions with these exceptions:
- a. Every coolant channel is assumed to have the nc=inal flov area (FA = 1.0).
- b. Every fuel rod is assu~ have (1) the =axi== calculated value
,; cf heat input, and (r' 4 and Fqn are assigned values of 1.0. ( c. The flow in each coolant channel is based on a power analysis with- 'l out flov =aldistribution factors. I
- d. Every fuel rod.is assu=ed to have-a nominal value for Fah nuclear.
The full =eaning of the maxi =u: and =ost probable design statements requires
- additional ec==ent. As to the 0.02 percent or 0.005 percent of the rods net included in the state =ents,' statistically, it can be said that no = ore than 0.02 percent or 0.005 percent of the rods vill be in jeopardy, and that in
-). general the n=ber. in jeopardy will be fewer than 0.002 percent or 0.005 per-cent. The state =ents do not =ean to specify a given nu=ber of DNE, but only
.N - acknculedge the-possibility that a given nu=ber eculd occur for the conditicns
'ID. as s =ed . 1
-li 0021.0 3 13 r- r w , - * - ,, ,,. ,i,. .- ,.-% ,, - --
l i
, .: h In su==ary, the calculational procedure outlined here represents a substan-1 i
tially i= proved design technique in two vays: 4
- a. It reflects the performance and safety of the entire core in the
, resultant _ power rating by considering the effect of each rod on the
! power rating, a
- b. It provides infor=ation on the reliability of the calculation and, therefore, the core through the statistical statement.
3.2.3.2.- Evaluation of the Thermal and Hydraulic Design
- a. Hot Channel Coolant Quality and Void Fraction ,
An evaluation of the hot channel coolant conditions provides ad-ditional confidence in the thermal design. Sufficient coolant flow has been provided to insure low quality and void fracticns. The quality in the hot channel versus reactor pover is shown in Figure 3-16. The sensitivity of channel outlet quality with pres-sure and power level is shown by the 2,185 and 2,120 psig system ' tressure conditions exa=ined. These calculations were =ade for an FAh of 1.78. Additional calculations for a 10 percent increase
, in Fah to 1 96 vere =ade at 114 percent power. The significant
) results. of both calculations are su==arized in Table 3-12. Table 3-12 Hot Channel Coolant Conditions Exit Exit Void Operating
- j. Pover, 5 Fah Quality, 5 Fraction, 5 Pressure, usic J
100 1 78 (-)3.0 5D) 0.h(*) 2,185 11h 1 78 15 79 2,185 128 1 78 6.6 28.7 2,165
- l. llh 1 96 7.l 30 5(,) 2,185 100 1 78 -1.0(b) 17 2,120
, 11h 1 78 39 20.0 2,120 , 128 -1 78 91- 37.2 2,120
'llh 1 96 95 38.9 2,120
" Subcooled voids.
(b) Negative indication of quality denotes subcooling-
~ The conditions of Table 3-12 were deter =ined with all of the hot channel factors applied. Additional calculations were =ade for
. unit cell channels.vithout engineering-hot channel factors to
.show the coolant conditions = ore likely to occur in the reacter core. A value for Fah of 1.78 was exa=ined with and withcut fuel asse=bly flow distribution hot channel factors at 2,185 psig as U =shown on Figure 3-22. These results show that the exit qualities 00Mt
-3 bh
frc the hottest cells should in general be censiderably lover h than the taxi =um design conditions.
- b. Core Void Praction The core void fractions were calculated at 100 percent rated pcuer for the nor:al cperating pressure of 2,165 psig and for the =ini=u:
operating pressure of 2,120 psig. The influence of cc: e fuel asse:- bly flev distribution was checked by deter =ining the tctal voids for both 100 and 95 percent total core flow for the two pressure condi-tiens. The results are as follows: Flev, % Fressure, tsig Core Veid Fraction, % 100 2,165 0.023 100 2,120 0.050 95 2,165 0.053 95 2,120 0.186 The = cst conservative condition pf 95 percent flow at 2,120 psiE results in no =cre than 0.19 percent void volute in the core. Con-servative maxi =un design values were used to take the calculation. The void *rogra uses a ec:binaticn of Eovring's(2I) radal with .I Zuber's(2b) ccrrelation between void fracticn and quality. The Eevring model considers three different regions of fcreed convec-tien boiling. They are: (1) Highly Subcooled Boiling In this region, the bubbles adhere to the vall while moving upward through the channel. This region is terminated when the subcooling decreases to a point where the bubbles break through the laminar sublayer and depart frc= the surface. The highly subcooled region starts when the surface te=perature of the clad reaches the surface tempe rature predicted by the Jens and Lottes equation. The highly subcooled regicn ends when m = E1
' sat _T culk V (A) where 4 = local heat flux, Etu/h-ft n = 1.863 x 10-5 (1L + 0.006Sp)
V = velocity of coolant, ft/s p = pressure, psia CH3'$'!" 3 L5
a
'j .
j. The void fraction)in this region is computed in the same man-ner as Maurer,t29 except that the end of the region is deter-mined by Equation (A) rather than by a vapor layer thickness. The nonequilibriu: quality at the end of the region is con-puted frc the void frr.ction as f%11ovt: 1 x 1 d=y,gt_1_ _ 1 (3)
#8
,"d j where x d = nonequilibrius quality at end of Region 1 a
d = v id fraction at T -T = , cf = liquid cc:ponent density, lb/ft3 4
. og = vapor component density, lb/ft .v '
(2) Slichtly Subcooled Boiline
'f ~g- In this regicn, the bubbles depart frc the vall and are
.D transported along the channel (condensation of the bubbles is neglected).
This region transcends to a point where the
- l ther=odynamic quality is zero. In general, this is the re-gion of major concern in the design of pressurized water.
reactors. j', The nonequilibrium quality in this region is co::puted fro: the folleving formula:
, , P
- x =x h d+ihfg(1 + c) z #~# SP d
where x = nonequilibrium quality in Region 2 1 hf = latent heat of vaporization, Stu/lb i - - . i y, = fraction of the heat flux above the single phase heat flux that actually goes to producing voids gp = single phase heat flux, Etu/h-ft c f,= = ass flev rate, lb/h 002t3 .I m L -_ _
f
}' .
l g' P = heated perimeter, ft h
- . y
- = channel distance, ft The void fraction in this regicn is cc puted frcs e
a= (D)
. . , 38.3 Afg o ~cgg (o - e )"1/h c f g C
o x + og/pf(1 - x ) '
+ .
C p 2 p
! 2 5 where. g = acceleration due to gravity, ft/s 1b '~ '**"
b g = constant in Nevten's Second Law = 32.17 lb f s C = Zuber's distribution parameter o , A = flow area, in 2 f
' ; :. o = surface tension l
. Equatien (D) results frc rearranging equations fcund in Ref-erence (33) and assuming bubbly turbulent flev in deter =ining
- the relative velocity between the vapor and the fluid. Zuber has shown that ' Equation (D) results in a' better prediction of the void fraction than earlier =odels based on empirical slip
^
ratios. (3) Eulk Boiling ,
- j. In this region, the bulk te=perature is equal to the satura-t tien temperature,.and all the energy transferred to the fluid
?
results in net vapor generation. Eulk boiling begins when the i thermodynamic-(heat balance) quality, x, is greater than the
- nonequilibrium quality, x*. The void fraction in this region
[. is'cc puted using Equation (D) with the thermodynazic quality, 'p x,. replacing x*. . -c. Coolant Channel Eydraulic Stability : ' Flow regine maps of = ass flow rate and quality were constructed in i order to evaluate channel hydraulic stability. The confidence in , the, design is based en a review of both analytical. evaluations (30, 31,32,331 and experimental results obtained in cultiple rod bundle
'~~G' y burnout tests. Eubb
- proposed by Baker (30}e-to-annular are censistent withand the bubble-to-slug ED' experirentalflov limits v.
4 ec 00?,:4 '} 3 L7
I
~
data in the range of interest. The analytical limits and experi- '
- h. = ental data points have been plotted to obtain the =aps for the four different types of cells in the reacter core. These are shown in Figures 3-23, 3-2k, 3-25, and 3-26. The experi= ental data points represent the exit conditions in the various types of channels just previous to the burnout cenditicn for a representa- I tive sa:ple of the data points obtained at design operating con-ditions in the nine rod burnout test asse=blies. In all of the bundle tests, the pressure drop, flev rate, and rod te=perature traces were repeatable and steady, and did not exhibit any of the characteristics associated with flov instability.
Values of hot channel mass velocity and quality at 11b percent and 130 percent pcver for both nc=inal and design conditions are shown on the =aps. The potential operating points are within the , bounds suggested by Baker. Experimental data points for the re-actor sec=etry with such higher qualities than the cperating con-diticas have not exhibited unstable characteristics. I d. Ect Channel D:iB Cc=rarisens I - DNE ratios for the hottest channel have been deter =ined fer the Ij' ' W-3 correlation, and the results are shown in Figure 3-15 ratics are shcvn for the design 1 50 axial cax/ avg sy==etrical DN3 cosine flux shape frc= 100 to.150 pertent power. The W-3 DN3 i ratio at the =axi=u= design power of 11L percent is 1.71. This
.cc= pares with the suggested V-3 design value of 1.3 A ratio of !- 1.3 is reached at 128 percent power at an exit quality of 9.1 l percent, which is within the prescribed quality limits of tP, cor- . relation. -
- m .
q I~ The sensitivity of the DNB ratio with Fn nuclear was exa=ined frc= i fj 100 to 150 percent power. The detailed results are labeled in Fig-il ure 3-27 A cosine flux shape with an F: of 1.80 and an Fdh of E 1.78 results in a W-3 DNB ratio of 1.hb at 11h percent power. Si=- ilar results are shewn for a value of Fz of 1.65 and for the de-g sign value of 1 5 The W-3 values are vell above suggested design
'f values.
i. l The influence of a change-in Fah was deter =ined by analv:ing the
-het channel for an Fah of 1.96. This value is 10 percent above
! the maximu= design value of 1.78. The resulting W-3 DN3 ratio is 1.2S at 11h percent power. This value is well above the ccrrels-tion best-fit values of 1.0 for the severe conditiens assu=ed.
- i
!F
; l '.
- e. fseactor Flav Effects
- r Another significant variable to be considered in evaluating the
- i. design is the total syste=;flov. Conservative values for syste= t and reacter pressure drop have been deter =ined to insure that the required -syste= flow is obtained in the as-built plant. The ex-
.q peri = ental programs previously outlined in 1.5 2 vill ccnfir: the
;y- pressure drop'and related pump head require =ents. It is sntici-
, pated tha- the as-tuilt reactor flev vill exceed the desicn vslue. s and vill lead to increased pcuer capability. 00ms 3 h8
i i l
. \
Tne reacter core flev and power capability were evaluated by de-h ternir.ing the steady-state pcVer DN3 rates cerru ficv. Analyre vere Ede for (a) variations of pcVer capability with total re-acter flew for a constant LNB ratio cf 1.30, (t) 253 ratics fcr decign flov vith variations in het channel mixing coefficients and (c) ONE ratics for gross flow variatiens of i 1C percent. The reru;ts are thcun in Figures 3-26 and 3-29 Fcr the analysis shown ir. Figure 3-28 fer design hot channel ecnditien, the flev was deter =ined that vculd give a DNE ratio cf 1.3; for a range of reacter pcVers. This analysis shows, for example, that a L53 ra-tic cf 1.30 can te raintained in the hot channel at 11L percent power with a tctal reacter ficv cf 109 x 106 lb/hr as cc pared with the available design flov of 131.3 x 106 lb/nr. The regn;t: c.ycur ey Line e. in z.rgure 2-c9 are ,ye
. .. E..s ratics fcr ratec fiev .
of 131.3 x 106 lb/hr versus pcVer. The limiting ccnditien it 125 percent pcVer fer a EHE ratio of 1 30. Lines 1 and 3 ch v the :53 ratics vertus pcver where the total rytter ficv har teen varied by 210 percent. Adequate DNS ratics can be cair.tained with a substantial reduction in reacter cociant systen flev. The foregoing sensitivity analyses were made using a fuel ascently desi~n 6 =ixing coefficient of 0.62. The final design value ray be as high as 0.06. A s cc+,h tube value for mixing withcut r;accr planes is 0.03 A sensitivity analysis fcr this range of ccerfi-cients was made i for the rated flev cendition. The results are
. shevr by Lines e and 5 c.,rigure -c--y an.2 ciscusrec ar. .
ncre cet al;
.n 3.c. 2 .2.,.j. s
- f. Eeacter Inlet Terrerature Effects The influence of reactor inlet temperature en power carability r .
at design flev was evaluated. A variatien c' ' 7 in reacter in-let terperature vill result in a power capability change of 0.6 percent at a given EN3 ratio.
- g. Fuel Terrerature
'(1) '<ethed of Calculatien A fuel temperature and eas pressure ec=puter ecce was devel-
,o--a .. o c e ., -
- - 4 ,
- e .'ua. l +e e_-, e . c--* ". e s , e x , - . "u r.
--e .. . . , ."a..e'.s='.'-.,
. .^- . .
equiaxed and eclumnar grain revth e center r.itir.c _ _ cf fuel .re - lets, fission gas release, and ficsicn gar precc.re. Fr: gram and data cceparisons vere made en the tasis cf the fracticn of the fuel diameter within these structural regicns: 1 (a) Cuter lirit of equiaxed grain greuth - 2,7CO F. (t) Outer limit of eclumnar grain grcvtr - 5,000 F. (c) Outer limit of =citen fuel (U?-)c - 5,000 F. 003 M Lc
Data frc: F.eferences 3h thrcugh 37 vere used to ec pare cal-culated and experimental fractions of the rod in grain growth g and central melting. The radial expansion of the fuel pellet is ccnputed frc: the mean fuel temperature and the average coefficient cf linear expansion for the fuel over the temperature range considered. This codel ec=bined with the model fcr calculating the heat transfer coefficient was cc pared with the model develcped by Notley et al(38) of AECL. The difference in fuel growth for the two calculation =cdels was less than the experi: ental scatter of data. The fuel may be divided into as many as 30 radial and 70 axial increments for the analysis. An iterative sclutien for the temperature distributien is obtained, and the the=al cenduc-tivity of the fuel is input as a functicn cf terperatu e. The j relative the = al expansion of the fuel and cladding is taken into account when dete=ining the terperature drop acrcss the gap between the fuel and cladding surfaces. The temperature i drop across the gap is a function of vidth, mean temperature, and gas conductivity. The' conductivity of the gas in the gap is determined as a function of burnup and subsecuent release of fission product gases. In the event of fuel clad centact, a
! contact coefficients are determined en the basis of methods suggested by Ross and Stoute(hl). The contact coefficient is
_) determined as a functicn of the mean ccnductivity of the in-terface caterials, the contact pressure, the sean surface roughness , the caterial hardness, and the ccnductivity of, the gas in the gap. 1 The analytical model cc:putes the a=ount of central void ex-pected whenever the temperature approaches the threshold ter-perature for fuel migration, and readjusts the density acccrd-ing to the new secretry. The prcgra: uses a polyncrinal fit relationship for fuel ther-tal conductivity. Three relaticnships were used to evaluate
; the effects of conductivity en fuel temperatures. The Ka' reference design curve which yields an integrated ther:s1 cen-ductivity of 93 v/ce, ie a presented in GEAP h62h.139) modification cf the relatiensiiip The conductivity relationships in GEAF-L62k and CVNA-246 j are cc pared with the reference design in Figure 3-30. McGrath i
A (LO) concludes that the CVNA-2h6 values are lever limits for the high temperature ecnditions. Fuel center temperatu es for
! all three of the conductivity relatienships at the peaking factors given in 3.2.3.1.2 have been calculated to evaluste the cargin to central melting at the ac:itu: Overicver and te show the sensitivity of the calculaticn regarding ther El con-m ductivity. Since the pcver peaks vill be turned off vit? ir-d radiatien, the peaking factorr ured are c;nr :rvative it -he -
end of life. t 00?,:7
=. . v-
, (2) Fuel Center Te=rerature Results The results of the analysis for center te=peratures with the
=ethods described above are shown in Figures 3-31 and 3-32 for beginning- and end-af-life conditions. The beginning-and end-of-life gas conductivity values are 0.09 and 0.01 Btu /h-ft 2 -F. The calculated end-of-life center fuel te=per-atures n're higher than the beginning-of-life values because of the reduction in the conductivity of the gas in the gap.
The effect is apparent even though a contact condition pre-vails. The calculation includes the effect of fuel swelling due to irradiation, but does not censider the effect of fuel cracking and expansion due to ther=al gradients. Credit is also taken for the flux depression in the center of the rod because of the self-shielding effect of UO2 (nonunifer= power generation). The calculated contact pressures are conserva-tively lever than those expected at end-of-life cenditiens in the hottest fuel rods, and the fuel te=peratures shown in the figures above are censervatively high. The B&W :odel gives very good results when compared to the results of others in the field as is shown in Figure 3-32. In the linear heat range of = cst interest, ie, approximately 20 kW/ft, there is only about 200 F difference between the
=aximu= and =ini=u values calculated. Also the s=all dif-ferences between the B&W curve and the other curves indicate ) the_ relative insensitivity of the results to the shape of the conductivity at the elevated te=peratures.
The =ost conservative assu:ptions, using GEAP h62h data with . relatively little increase in ther=al conductivity above 3,000 F, result in central fuel =elting at about 22 kW/ft, which is about 3 kW/ft higher than the =axi=u design value of 19.2 kk/ft at 11b percent power. t3-32) Further evaluation of the two figures shows that central fuel =elting is predicted to occur between 22 and 26 kW/ft depending on the ti=e-in-life and conductivity assu=ptions. The transient analyses at accident and nor=al conditions have been =ade using the design fuel ther=al conductivity curve (Figure 3-17) to reflect a conservative value for the maxi =u= average te=perature_and stored energy in the fuel. Use of this curve results in a higheggte=perature and, therefore, a lower Doppler _ coefficient, since it decreases with te=perature. Thus,-the resultant Doppler effect is also conservative. (3)i Ecuilibriu= Cycle Average Fuel Te=peratures L An analysis has been made te show equilibriu: average fuel
.cenditions in the core. A' typical fuel cycle, end-of-life, ccndition was used to deter =ine the fraction of fuel at a
. given average condition. The results are shcvn in Figure 3-33
-()()::tt3 3-51;
where the average te:perature varies frc 1,300 to 3,300 F,
) and the entire ccre average temperature is about 1,600 F.
The bundle average powers as shown in Figure 3-3h were used to obtain the fuel rod heat rates. A s3 setrical ccrine axial power distributicn with a 1.50 r u / avg value as shevn in Figure 3-10 was used to predict the axial heat rate dis-tribution. It was assumed that 97.3 percent of the pcVer is generated in the fuel. The fuel rods were divided into Ib axial and 10 radial segments to obtain the temperature dis-tribution for this analysis. The heat rate for every fuel
- rod in the core was increased by a local peaking facter of 1.05 to account fer uncertainties in the calculation of local peaks. This has the bulk effect of raising reactor power to 105 percent.
The fuel cenductivity values frc GEAF h62hP "9) were used to provide conservative values fer fuel ceniuctivity . The cui-mu pcVers occurred in fuel assemblies with one and two cycles
' of operaticn as shown in Figure 3-3h, and the asse=blies with the highest burnup did not exceed 1.0h3 tires the average power for the typical case analyzed. The results shown in Figure 3-33 vere cade by grouping all segments of fuel by
; temperature. Typical 6 and 10 kW/ft red radial temperature i profiles are shown in Figure 3-35 Typical fuel-to-clad heat
; transfer coefficients used were 2hD and L60 Etu/h-ft'-F for 6 and 10-kW/ft heat rates, respectively. The corresncnding
~
beginning-of-life coefficients are atcut 500 and 700 Etu/h-ft2.7 l e for 6 and 10-kW/ft heat rates. t The temperature calculations are conservatively based on ncr-l' inal fuel dimensicns without fuel irradiation swelling er crack-
; ing. Fuel swelling and cracking vill result in higher fuel-i to-clad heat transfer coefficients than those used in the anal-I ysis. The volumetric heat generation rate is higher at the i outer periphery of the fuel than in the center regien, and the l cperating temperatures vill be lover than these predicted I by the unifor: radial power condition assured in the analysis.
- h. Fissicn Gas Release (h2)
The fission gas releace is base;l n results ra~ rted in GFE h59o-(b5) Additional data frc GEAp h31hb3 , AECL-603l , and CF-60-12-lh have been ec pared with the suggested release rate curve. The re-
)
lease rate curve (k2) is representative cf the upper limit of release j data in the temperature regicn of =ost i pertance. A design release I rate of h3 percent and an internal gas pressure of 3,300 psi are used j to deter =ine the fuel clad internal design conditicns repcrted in
! 3.2.h.2, Fuel Asse:blies.
The design values fer fission gas release frc the fuel and fcr the
=axi=um clad internal pressure were dete=ined by analyzing varicus p cperating ecnditiens and assigning suitable targins fer p::sible in-V creases in local cr average bu nup in the fuel. Adeque'^ m rgins are 00.3:3 3-52
provided without utilizing the initial porcsity voids present in
') the UO2 fuel. A detailed analysis of the design assunptiens fer fission gas release, and the relationship of turnup, fuel grcvth, and initial diacetral clearance between the fuel and clad, are suscariced in the following paragraphs. An evaluaticn of the ef-feet of having the fuel pellet internal voids available as gas holders is also included.
(1) Design Assumptions (a) Fissien Gas Release Eates The fissicn gas release rate is calculated as a functicn cf fuel temperature at 11h percent cf rated pcver. The procedures fer calculating fuel terperatures are dis-cussed in 3.2 3.2.3.g. The fission gas release curve and the supporting data are shern in Figure 3-36. Most of the data are on or belev the design release rate curve. A release rate of 51 percent is used fer the pcrtien of the fuel above 3,500 F. The fuel terperatures were cal-culated using the ELW design fuel thermal conductivity curve which yields conservatively high values fer fuel terreratures. i - (b) Axial Pcwer and Burnun Assuertiens
) The temperature conditiens in the fuel are determined for the = cst severe axial pcwer peaking - ated to cccur.
Two axial power shapes have been evaluas deter =ine the taxi =u release rates. These are 1 50 and 1.70 tax / avg shapes as shcvn in Figure 3-10 and repeated as part of Figure 3-37 of this analysis. The quantity of
.< gas released is found by applying the temperature-related i release rates to the quantities of fission gas produced l alcng the length of the hot fuel rod. , The quantity of fission gas prcduced in a given axial Ic-cation is obtained frc: reacter core axial regien equilib-riu: burnup studies. Three curves shcving the axial distri-bution cf burnup as a 1ccal-tc-average ratic alcng the fuel rci are shown in Figure 3-37. Values cf 100, 300, and 930 days of cperaticn are shczn.
The 930-day, or end-of-life conditien, is the ecndition
, with the =aximu: fission gas inventcry. The average burn-1 up at the end of life in the het fuel red is 38,150 i mwd /MTU which has been det&rnined as follows:
Calculated Ect Eundle Average Eurnup, igd /:CU 33,000 Hot Fuel Ecd Burnup Factor 1.05 A s) m trgin fcr Dalculatien Accuracy 1.10
, Ec- Ecd Maxinus Average Eurnup, mwd /MTU 38,150 3-53 00220 ~
O The local burnup alcng the length of the fuel red is the ) product of the hot rod maximum average value ateve and the local-to-average ratio shc"n in Figure 3-37. The re-sulting het rod Ic a1 naximum turnup fcr the 930-day, end-of-life conditicn is about L2,000 mwd /MTU. This is the caximum calculated value. Ecvever, local values to 55,000 K4d/MTU have been evaluated to insure adequate local fuel cladding strength fer possible increases in average cr local burnup over the life of the fuel for various fuel =anage ent procedures. (c) Ect Rod Pcver Assuretiens The maximum hot red total pcVer cecurring at any time in the life of the fuel has been used to calculate the over-pcuer terperature conditions. A het red pcVer cf 1.78 times the average red power has teen applied. This re-suits in a maximun linear heat rate cf 19.2 kW/ft which ccrresponds to 11L percent of the taxinun thernal cutput (16.83 kW/ft) shown in Table 3-1. This is a conservative assumption when coupled with the end-of-life fission gas inventory since bundle, and individual fuel red pcVer is expected to decrease with fuel burnup. A study of the power histories of all of the fuel assemblies te equilib-riu: ccnditions shovs that the p wer: in the tundles dur-ing the last 300 days of cperatien are not =cre than 1.3 ) times the average bu=dle power. The peak bundle ratio of 1.69 (1 78 + het red ratio) vill enly occur during the first two fuel cycles when the fission gas inv-n cry is less than the taximum value. (d) Puel Grevth Assu=ttions The fuel growth was calculated as a function of burnup as indicated in 3.2.h.2.1. Puel pellet dimensiens in the thermal temperature and gas release =cdels were in-creased to the end-of-life conditiens as determined above. (e) Gas ccnductivity and Centact Heat Transfer Assu::tions The cutntity of fission gas released is a functicn of fuel terperature. The temperatures are influenced by three facters: (a) the ccnductivity of the fisti:n gas in the gap between the fuel and clad, (b) the distetral clearance between fuel and clad, and (c) the heat trans-fer conditions when the fuel expands enough to centact the clad. A gas conductivity of 0.01 Etu/h-ft -F based cn L3 per-cent release of fissien gas at the end-of-life ccndition was used in the analysis. Diametral clearances cf 0.0025 to 0.0075 inch reflecting ~4 "- and taximum clearances after fuel growth vere analy ci. Tne certact hent transfer ccefficients were esicult:cd is cuicest e; ir Reference L2. 00 m., . 3-5L
F (2) Su==ary of Results t
=
' The fission gas release rates vere determined in the first evaluation. Rates were found for varicus cold diametral clearances and axial power peaking and burnup shapes. The results are shown in Figure 3-38. The IcVest curve is the expected condition for a 1.70 axial pcVer shape with a 930-day axlal burnup distribution as shown in Figure 3-37. The increase in release rate with diametral clearance results frc: the fact that the fuel temperature =ust be raised to
- higher values before contact with the fuel clad is made.
The release rate at the mini =um clearance of 0.00h5 inch is 15 percent. This condition is equivalent to a 0.0025 inch gap after irradiation grovth and produces the maxi =u clad stress (=axi=u: sized pellets with mininum internal diameter cladding). The release rate of 3L percent for the cximu= diametral clearance of 0.0095 inch vill nct occur with the =axinu: stress ecnditien due to fuel grcvth, since the fuel has =cre roo: to grov into the clearance, j Two additional cases were examined to check the sensitivity of the calculations to axial power and burnur shapes. The results are shown by the upper two curves in Figure 3-38. The top curve is a plot of the release rates when it is as-suced that both the axial power and burnup inventory of fis-sion gas are distributed with a 1.50 max / avg ratio as shown en Figure 3-55 Si=ilar results are shevn for the 1 70 nax/ avg power ratio'with a 1 50 max /avs burnup ratio. These curves show the release rates expected are not strongly in-fluenced by the various power and burnup shapes. The second evaluation shevs the resulting internal pressures due to the release of fission product gases. Plots cf in-4 ternal clad pressures for the expected 930-day axial burnup
- l. distribution and a 170 =ax/ avg axial pcVer shape are shown in Fi Eure 3-39 The lover curve is a p2ct of internal gas I pressure assuming that 6.3 percent of the fuel volume is j
avc'.lable to hold the released gas (open perosity). The present design condition being used in clad-stress calcu-i lations assu=es a closed pore condition with all released i gas contained outside the fuel pellets in spaces between
' the expanded dished ends of the _ pellets, the radial gaps (if
!i any), and the void spaces at the ends of the fuel rods. The effects of fuel densification and grain grevth described in 3.2.3.2.3.g.are-included in the analysis. The calcula-tien of =aximu: pressure is also relatively insensitive to the axial burnup distribution as shown by the line in Figure
-j 3-39 for-a 1 50 =ax/ avg axial power and burnup shape. (This corresponds to a local burnup peak of 57,000 igd /:CU. )
Tnere is evidence that the UO y L7,LB) are usually cracked. 3pelletsinanuclearreacter(L6' sensitivity analysis en fuel-
-v. -tc-clad gap conductance was perfor:ed to illustrate the effe: s 00~5.9 s I
3-55
3' of fuel cracking. Two independent fuel =odels were developed
~$# to simulate this phenc=enon. In one =odel, the cold =inimu=
fuel dia=eter was increased by 0.0025 inch in order to de-crease the dieretral gap and increase the gap conductance (2.5-=i1 cracking f et). In the other =edel, the =ethods i of Ross and Stoute vere used to evaluate gap ccnductance with the fuel-to-clad centact pressure equal te zerc ("0-contact" cracking effect). This =odel assu=es that the fuel cracks enough to ec=e into centact with the clad, but not I enough to exert pressure against the clad. As the higher heat rates are reached, the ther= n Growth of the fuel causes a contact pressure which is usec in the calculatiens. Fuel-to-clad gap conductance as a function of linear heat rate is
, plotted on Figure 3 ho te shov the =arked increase in gap ! heat transfer due to fuel cracking. To further illustrate the effects of fuel cracking, a sensitivity analysis en rod internal pressure was perferned using the foregoing fuel =cd-els. Figure 3 kl shows the results of this study fcr the two j cracked fuel =odels and the ideal ther=al expansion =edel at t the =ost conservative fuel-to-clad heat transfer conditic7s } (largest gap, end-of-life Eas conductivity, etc). The fuel l ' cracking =odels show a marked reduction in internal gas pres-
- i
- sure fro = the desi En conditions.
s 4
- A para =etric study on gas pressure versus reactor power was
, perfor=ed to further illustrate the conservatis=s involved in l
- 7) the Eas press e analysis. Figure 3 h2 shows the results for the two cracked models and the ideal ther=al expansicn =odel at the verst fuel-to-clad heat transfer conditions. A conser-
'i vative steady-state reactor power level of 2,795 !Gt (llL
- percent of rated pover) was used to obtain fission Eas release.
j This analysis shows that power increases frc= rated pove.. .
. 2,h52 to 3,000 IGt result in very s=all increases in fuel rod I internal Eas pressure. Final design consideration of power conditions and fuel cracking vill result in lower fuel te=-
! perature and associated internal 6as pressures.
} ,
The allovable design internal pressure cf 3,300 psi is well
~a bove the =axi=u= values of internal pressures calculated for open or closed pellet pores, and the =axi=u= internal pres-sure should only occur with the =aximu= dia=etral clearance condition. An increase.in average fuel burnup can be toler-ated within the prescribed internal preccure design li=its.
j' It has .been . indicated in Reference 38 and in AECL-1598 that 1 the UO fuel is plastic enough .to flow under low stresses 2 j - vhen tne te=perature is above 1,800 F. That fraction of the fuel belev this te=perature may retain a large portion of the original'percsity and act as a fission gas holder. The hot-test axial locatiens producing the highest clad stresses' vill have little if any fuel belev 1,800 F. Ecvever, the ends of the fuel rods vill ~ have sc=e fuel telev this te=perature. 1 aow.- . z . . 3-56
'l 4 : The approxi= ate fraction of the fuel belov 1,800 F at over-
/ h power for a 1 70 axial pcVer shape is as follows for various
, cold diametral clearances.
i 3; Clearance, Percent of Fuel
- j)- in. Belov 1,800 F, %
1!
- 0.00k5 LO 0.0070 20 4
0.0095 5 iI The retention of fuel perosity in the low te=perature and lov
- l burnup regions vill result in =edest reductions in internal
- )
gas pressure.
- 1. Ect Channel Factors Evaluation
{' (1) Rod Fitch and Spvine ' A flow area reducticn factpr is deter =ined for th6 as-built fuel asse=bly by taking caannel flow area =easure=ents and statistically deter =ining an equivalent hot channel flev area reduction factor. A ftel assembly has been =easured, and the ,l._ i results are shcvn in Table 3-9 Interier channel =easure-
,"7 l- =ents and =easure=ents of the channels forned by the cuter: cst fuel rods with adjacent asse=blies have been analyzed. Co-s '- efficients of varl? tion for each type of channel have been
- f '
determined. In the analytal solution for a channel flev, '{, each channel flow area is reduced over its entire length by the Fgfactors shown in Figure 3-20 for the desired population l- protected at a 99 percent confidence. The hot channels have j - been analyzed using values for 95 percent population pro-j tected, or FA in the interior cells cf 0 98 and FA in the
,p vall cells of 0 97 as . listed in 3.2.3.1.1.j .
Special attentien is given to the influence of water gap vari-ation between fuel asse=blies when deter =ining rod powers. Nuclear analyses have been =ade for the nc=inal, =axi=u=, and
.=ini=u= spacing between adjacent fuel asse=blies. The nc=inal and =aximum het asse=bly fuel rod powers are shown in Figptres 3 43 and 3 kh. The hot channel nuclear pcver factor (Fah nu-clear) of 1.78 shcun in 3.2.3.1.1 is based on Figure 3 LL fer the vorst water gap between fuel asse=blies. The facter of
- j. 1.783 is a product of the hot asse=bly factor of 1.68 ti=es j.
the 1.061 hot rod factor. This power facter.is assigned to i the hottest unit cell rod which is analyzed for turnout. Feak-ing factors for cther channels are sbtained in a si=ilar =an-ner. . In all cases, the ec=bined flev spacing and pcwer peak-ing producing .the lowest.DNB ratio is used. nn a 003.M 3-57
'2) Fuel Fellet Diameter, Density, and Enrichnent Facters Variatiene in the pellet size, density, and enrichment are re-flected in ecefficients of variatien :.urbers 2 thrcugh 7 cf Table 3-9 These variaticns have been obtained frc: the rea-sured or specified tolerances and conbined statistically as described in 3.2.3.2.2 to give a pcver factor en the het rod.
Ter 99 percent confidence and 95 percent population condi-tiens, this factor , Fq , is 1.011 and is applied as a power increase over the full length of the hot channel fuel red. The local heat flux factor, Fqn, for similar conditions is 1.01h. These hot channel values are shewn in Table 3-1. The ccrrespending values of Fq and FQ " with 99.99 percent popu-lation protected are 1.025 and 1.03, respectively. A cen-servative value of Fqn of 1.03 for 99 percent confidence and C9 99 percent population is used for finding the taximum fuel linear heat rates as chevn in 3 2.3.1.2. These factors are used in the direct solution for channel en-thalpies and are not expressed as factors en enthalpy rise as is often dene. The coefficients cf variation vill be under centinuous review during the final design and development of the fuel assembly. (3) Flev Distribution Effects
, ) ~n let Flenur Effects The final inlet plenu: effects vill be determined frem the 1/6 scale codel flow test now in progress. The initial runs indi-I cate satisfactory flow distribution. Although the final nu- ; clear analysis and flow test data cay show that the hot bundle positicns receive average or better flow, it has been assu=ed that the flow in the het bundle position is 5 percent less than average bundle flow under isothermal conditions corresponding ! to the rodel flow test ecnditions. An additional reduction of . flow due to hot assembly power is described below.
F.edistributien in Adfacent Channels of Dissisilar Ccelant Cenditiens The het fuel assembly flev ;- less than the f1cv thrcuch an aver-age asse:bly at the sane core pressure drcp because cf the in-creased pressure drop associated with a higher enthalpy ar.6
; cuality ecndition. This effect is allowed for by making a di-rect calculation for the hot assembly f1cv. The ec bined ef-
! fects of upper and lever plenu: flev conditions and heat input to the het assenblies have been used to determine het assembly ficvs. The verst f1cv saldistribution effect has been assured in the initial design, and the sininu: hot asse b2y flev has teen calculated to be 89 percent of the average assently f1cv
'g at 11L percent everpcVer. Actual het asse:bly f1cvs are calcu-
\ s~ lated rather than applying an equivalent het ch,znel erthal;;.
rise fucccr. <> q n.,
' - r 'e p 3-5S
tr J O Physical Mixing of Cool' ant Between Channels The flow distribution within the hot asse=bly is calculated with a mixing code that allows an interchange of heat be-g tween channele. Mixing coefficients have been deter =ined i from multirca mixing tests. The fuel asse=bly, consisting ,t . of a 15,x 15 array of fuel rods, is divided into unit, vall, centrol red, and corner cells as shown by the heavy lines in j Figure 3-k3 The mixed enthalpy for every cell is determined simultaneously so that the ratio of cell to average assebbly enthalpy rise (Enthalpy Rise Factor) and the corresponding Iceal enthalpy are obtained for each cell. Typical enthalpy rise factors are shown in Figures 3 h3 and 3 kh for the hot
- and surrounding cells. The a su=ptions used te describe the d channels for the peaking and enthalpy rise facters shown are 3,.
given in 3.2.3.2.3.j, which follows. , , 4 i j_' j . Evaluation of tF a DNS Eatios in the Unit, Wall, Centrol Eod, and Corner Cel3c DNB Results at Rated Flow i The DNB ratios in the hot unit ' cell at the =axi=um design condition described in 3.2.3.1 are shown in Figure 3-15 The relatienship shewn is based on the application of the W-3 ccrrelation. An ad-6 ditional' sensitivity analysis of the assembly corner, vall, and
]l - c'entrol rod cells has been made for the worst ecmbination of fuel assembly. spacing and~ power peaking.
-The sensitivity of.the asse=bly design with respect to variations
. . of mass velocity (G), channel spacing, mixing intensity, and local peaking on the DNB ratios in the fuel. asse=bly channels has been
{' evaluated by analyzing the nominal conditions and a postulated
} -verst case condition. The s"--ary results are shown belev in-i ; Table 3-13 for the nc=inal case and 3-lh for the maxitun design or postulated vorst case. .The unit cell DNB ratios are repeated for t
cc=parison. All of.the DNB ratios are for 114 percent overpcVer.
?
4-1 . 4 1 1 g 00226 s
' , 3-50
Table 3-13 h DN3 Eatics in the Fuel Assembly Channelc ('a'-3 ) Nc=inal Case Cell Tyre G , lb/h-ft x 10~ INER (W-3) (llL0 Pcver) Unit 2 55 1.97 Corner 2.59 1.97 Wall 2.59 2.03 Control Rod 2.L5 2.15 Table 3-lh DN3 Eatics in the Fuel Assembly Channels (W-3) Postulated Worst Case (Design) Cell Tyre G , lb /h-ft x 10- DN3R (W-3) (11L% Pcver)
) Unit 2.31 1.71 Corner 2.18 1.81 Wall 2.2h 1.78 Control Rod 2.20 1.83 The DNE ratios in all channels are high enough to insure a ecnfidence-pcpulation relationship equal to or better than that outlined in 3.2.3.1.1 for the hot unit cell channel. All of the vall, cerner, and control rod cells have DN3 ratics equal to or lever than that of the unit cell hot channel. This results frc a more favcratie flow to power ratio in these cells associated with relatively larger flow areas.
The DNE ratios were obtainem by ec paring the fuel rod 1ccal heat fluxes and channel coolant conditions with the limitations predicted by the correlation. Typical results are shown in Figures 3 L5 and 3 h6 for the nominal and vorst case conditiens in the unit cell. Fuel Ecd Fever Peaks and Cell Coclant Ccnditicns The nerinal case local-to-averaSe red pcVers and the 1ccal-t:- 5 average exit enthalpy rise ratics are shewn in Figure 3 L3 f er the
'/- het ccrner, vall, control red, and unit cells in the .ct fuel c:sc1-bly. '.' slues shewn are for netinal vater ga;r tetveer the 5:: fuel
- 60
- - 00'm ~r
. r_ _
2 l 1 I i i , i g asse bly and adjacent fuel asse blies with nc=inal flow to W ho ,(uel asse bly, and with a minimum intensity of turbulence, O, 1 equal to 0.02. q Additional tests are being run to determine the maximum values of
' intensity of turbulence associated with the fuel assembly. The expected value is greater than 0.02 since a value of 0.03 is ob-l tained in sdooth tubes. The fuel assembly spacer grids will in-
- i duce turbulence and improve coolant mixing.
4 i ! f The postulated worst case local-to-average rod powers (nuclear > i peaking factor) and exit enthalpy rise factors in the hot fuel asse bly are shown in Figure 3-4h. The facters were determined ; j for this case with the mini =u= vater gap between the hot fuel J asse=bly and adjacent fuel asse=blies, with minirum flow to the
- hot. fuel assembly, and with a minimum assumed intensity of tur-
.]. bulence, a, equal to 0.02. An evaluation of rinimum, nerinal, 7 and taximum spacing between asse=blies showed the mini =un to have t the lowest DNB ratios.
I' i
- A =ixing coefficient of 0.02 was used for both nominal and design
[, verst case analyses. Final design values of about 0.06 are likely. ]j The influence .of mixing coefficients is shown in Figure 3-29, which J- shows values ranging frcm 0.01 to 0.06. The value of 0.02 is sur-
- j ficiently conservative fcr design evaluation. The conditions an-C alyced to obtain the DNS ratios for various values of the mixing
"() coefficients shown in Fisure 3-29 vere outlined previously in 3.2.3.2.3.1.
4 I Fuel Asse:bly Power and Eated Flow Conditions 1 ! f The ncsinal and postulated vorst cases were run at llh percent reactor power' with the nominal and worst Fih factors shown in i )j - 3.2.3.1.1.c. The 1 50 modified cosine axial pcVer shape of 1 -Figure 3-10 was used to deceribe the worst axial condition. A -. [' p,a The hot'asse=bly flow under nczinal conditicns without a flov l' =a1 distribution effect is 96 percent of the average-assembly ; flow, and the reduction in f1cv is due entirely to heat input effects. The hot asse bly flow under the verst postulated b i (*)The intensity of' turbulence, a, is defined as V' /V r w( vhereV{_isthe.transverseecmponentofthefluctuatingturbulentvelocity, ik/ - and V-is:the.ccolant1 velocity.in the axial directien. This rethod cf cct- ,
.puting mixing is described by Sandberg, E.'0 , and.3ishey, A. A., CVTE'
! Thernal-Hydraulic Lesign for 65 lG ' Grcss Fissicn ?crer, CV::A-227.
'002?s
~3-61,
O g= conditions is 89 percent of the average asse=bly f1cv and cen-
- F siders the worst ec=bined effects of heat input and flev =al-distribution.
- k. DNE Results for Postulated Loss of an Internals Vent Valve The reactor arran6e=ent includes vent valves above the ccre to equalize the pressure between inlet and cutlet regions during a icss-of-coolant accident. The effective ccre flow will be re-l duced in the unlikely event that a valve disc breaks off. A 5
DN3 analysis was made to show the design cargin for a postulated accidental failure of one valve disc. i An arrange =ent censisting of valves with a 1L-inch diameter throat was investigated. In the event the disc frca one cf these valves-is cc pletely re=oved, a small reducticn in ef-fective core flcw for heat rencval vill te experienced. Ap-
- prcximately 5.7 percent of the inecting flev vill bypass the core through the valve opening. EcVever, the reduction of re-sistance results in an increase in total systen flew of about 1.1 percent. The net reduction of flow for core heat r,e cyal I is h.6 percent. .
The minimu: DN3 ratics for the reduced effective core flev cc - pare with the full flev ratics as follcvs:
) Percent Rated Power DNER (Full Flow)
DNER _( Reduced Flev) i '
!, . 100 2.21 2.05 i
l. t 107.5 (Trip 1.91 1 76 Set Point) i llh 1.71 1 55 1 DN3 ratics were deter =ined for the verst corner, centrol red, vall, er unit cell. The DN3 ratios in the het unit cell vere the levest.
"-a ~*-4 u: DN3 ratio3-"-at the trip set point of 107.5 percent power is well above the
- ecen= ended value of 1.30. The DN3 ratio of 1.30 is naintained up to 123 percent pcwer fer the postulated i verst case design conditiens.
A cceplete sensitivity analysis has been rade to determine the ef-fects of design ficv and unexpected -core bypass flow frc= the inlet F to the outlet cha:bers in the rcacter vessel. The results are
' shevn in Figure 3 h7 Bypass flow vas ; varied frc: 0 to 10 percent vhile holding a ecnstant core averaEe te perature of 580 F. A design allcvance' of 2 percent (2.63 x 105 -lb/hr) bypass flev fer vent valve seat and fitup leakage is included in all calculatiens for nc inal or maxi =u: design DN3 ratics. This design cenditien 2 ~
f is' indicated by Line 2 and is identical with rated conditicns in
. Figure 3-15 as previcusly discussed. Line 5 shcus the LNE ratics verets'pcVer fcr a ccndition of the;1 css cf ene vent valve disc s
002?9
\; 3-62
.\
- s. :
- ., , .-. -- , . -- , _ _ - , _ ,e
s plus a 2 percent typass. Additional lines are shcun fcr 0, 5, h and 10 percent variations in typass ficv.
'. 2. 2.7 2.L
. 7 . ' " a '. .' c . c # . . e ..a' e '.' e *.*. ". -a. ". a_
1 q A vapor Icek proble: could arise if water is trapped in the steam generator black.i43
- . . -
- b. e '.' cv c.',
. e -' a. a: ' .. ..- ,.- + k.. a. +c , c' . +.5..a. .ea-.-. *
. v uesa'-- *ca
. c^'d
-. .' ee, leak. Under this conditien, the stean pressure at the top of the reacter would rise and force the stea: tubbles through the water leg in the bettc cf the stea generater. This same differential pressre that develcps a water leg in the stea generatcr vill develcp a water leg in the reacter vessel which could lead to uncovering of the core.
The rest direct solutien to this rrchlen is to equalice the precsure acrcss u the ccre support shield, thus eliminating the depressien of the water level l in the core. This can be accc:;11shed by vent valves in the ccre st.pport
! shield which provide direct ec unicaticn between the reactcr ple--
the tcp cf the annulus. These vent valves 'cpen on a verv . lev rressure dif-ferential to allev stear generated in the core to ficv directly to the leak frc: the reactor vessel. Althety;h the flow path in the stear generatcr is blceked, this is of no ecnsequence since there is an adequate flev *,ath to i rencve the steam beinc-- cenerated in the ccre.
"2'h. e , ~.m. e l ' 3 . = ,, desien o'.
. _ . . 4.4s m v=.'ve d. e_c'.~.~.. . . . . ..
4 7.4-"-a e-- 51. s ~S.
. . a. v _= .' ".- a disc hangs c1csed in its natural position. A flat, stainless steel seat cal <..s., es .,ec:
-s .. . . . - - A 5 ce.
<.C,;, e.ees .-nr. . . ve+.< . . . - - . ~e..-l....--s...,_
. .. . c-- . . ...---a , . ,_
g" i t o '. .'.e u,, a.. ,. - "'er s..e-ke. ae s e..'l.. . I . *.k. e ev a.. .' a. r_- -.da.*., 4 . . - . ' .. '...a. 3 reverse pressu e differential vill cpen the valve. At all times during
. ... , , .ea-. .e c,e.-a*mc.., 4
- k.. a. y . a. .e. s ure i n
- k.e a ..."' " e . . . s
+..b.a. c"'. _e .i d .. e' the core support shield is greater than the pressee in the plenur cht. ter L en the inside of the core support shield. Acccrdingly, the vent valve vill l be .k. eld closed d" -i ~.e ..^-. a.' c, e. a+. 4 c . W.i *mk. #. c "-
- a. a . '. c c ^ ^ ' 5 .*. ,"--e cperating, the pressure differential is L2 psi resulting in a several thou-I sand pcund closin6- force en the vent valve.
Under accident conditiens, the valve vill tegin to cpen with a pressure dif-ferential in a directicn opposite to the ner:11 press = e differential cf aw- ..... o . 2.,. y~; . ca_ vale,s. ,- . 4. .3. - '.v. A . ". .. ' s ~f .i .+. , - hm a. c-a a. c. "..a. ". C. . a_ --- "r".-
~
..., .c. -. . e. n. .. a. .a*"-a'.
c 1ce..'..e
. '. c. e c .' *.h. a
- .a've.
. W . ".".. =_ ~,aee"-a. d' . .". a..-a . .-
..c, ,_-ei,
.;,,.._ e.,
e.e3+.e. .wn,. ~,a_..
. ...- +..".e . a' . ~. e - "la- k e ". "' ' , y . "n .i ". . . +.$. 4. s
, e. c_ s. -,- a. 4 -. 4 ". e . e . .* i a' . , *m*.e "$ '. e . .' a_- . a l # .. ...*e cc.-a 5 *
- . ".-"'d. k. a. -.a.~-"-. e **a_
- c-f ^. .# "...a. cc-a. . I*. ce.de. . '. c . * *.- -. a ac.e +~^- ke P.al' . ". c.e.-a#, a _c _e " 4. ..s~ _e c ' .4 d
. .-c. o .- 4. ...m., a k _'* ~. w.~ _
. . ' c .' *. ~a c e^ *. e , = ~.3e _c c "
- a. "'-.".e_-=...'.'=-' v .'_,.4 y..
+
' c. , ,.e ..u. .. . , . c . .e -, a- n.n <c, e u-c,e.<..e-
~
. .-. -.. . >m e de ve., c-y n. a
. . -u. a...e -
y . . a .. . . .c,. c. a. 3 times that required to open the valve ec:pletely. This is a censervative limit since it assures equal density in the ccre and the an.ulus surround-ing the core. The hot, stean-vater rixtre in the ccre v- 1_ . ave a density f much less than that cf the cold water in the annulus, and scntwhat greater
. n essv e A<.r.e e. e .r.+w .; .,.- , e C Cn.,A --w- vn .c1e. L . . . a.a_=
v a-r. C.v a- .. u. . a. $c-a , .- .o te ~.~~.c
. - %.~-. V. ~e .1 '. . uncovered. . \, )
I 00 3o- _.n
- ~-
)
Assu=ptions At the present ti=e, an analog co=puter si=ulation is being developed to evaluate the perfor=ance of the vent valves in the plent: chamber. This analysis vill be used to demonstrate that adequate stea= relief exists so that cooling of the core vill be accomplished. The basic model is a simulation of the reactor coolant syste which includes r the effect of the e=ergency cooling by the ECCS, the effect of stea= gen-eration in the once-through stea= Eenerators, the effect of stea generation in the core, and the effect of operation of the vent valves. The model is cc= posed of four basic regions that simulate the water volume in the annulus between the reactor vessel and the core, the water volu=e in the core, the , stea= volume above the core water level, and the stea volume in the region between the vent valves and the break location. Fluid flow between each of these regions, flow frc= the emergency injection syste=, steam flow through the ' break, and possible water spillaEe frc= the break are all censidered. The ccre volumetric heat Eeneration and heat transfer to a changing water level in a five-secticn core is censidered. Using the computer progra=, the required nu=ber and capacity of valves will be obtained. Conservative assu=ptions on, core decay heat, flov losses, heat
, transfer coefficients, and the available capacity frc= the e=ergency injec-tien syste=s.will be used.
The exact number of valves and their size vill be specified upon ec=pletion ~j of this analytical study. The perfor=ance of the valves =ust meet the cri-terion for core cooling that has been defined in lb.2.2.3.2 of the PSAR and is quoted below: ,
"The perfor=ance criterion for the_ emergency core cooling equip-
=ent is to limit the clad temperature transient below the clad
=elting point-so that fuel gec=etry.is maintained to provide core cooling capability. This equipment has been conservatively f
sized to limit the clad temperature. transient to 2,300 F or less as temperatures in excess.of this value promote a faster zir-ceniu=-vater reaction rate, and the termination of the transient near the =elting point would be_ difficult to de=cnstrate." d
, ,, 0023.7 .
3-6L'
v-I t S t i I . 3 2.h MS.NICAL IESIGN IAY0lff 3 2.h.1 Internal Layout Reactor internal components include the plenum assembly and the core support assembly (consisting of the core support shield, vent valves, core barrel, j lower grid, flow distr,1butor, in-core instrument guide tubes, thermal shield, and surveillance holder tubes). Figure 3-48 shows the reactor vessel, reactor
; vessel internals arrangement, and the reactor coolant flow path. Figure 3-h9 shows a cross section through the reactor vessel, and Figure 3-50 shows the core flooding arrangement.
Reactor internal components do not include fuel assemblies, orifice rod assemblies, control rod assemblies (CRA), surveillance specimen assemblies, or in-core instru=entation. Fuel assemblies are described in 3 2.L.2, control rod assemblies and drives in 3 2.h.3, surveillance specimen assemblies in L.L.3, and in-core instruzentation in 7 3 3
; Tne reactor internals are designed to support the core, r.aintain fuel assembly J align =ent, limit fuel assembly movement, and maintain CRA guide tube alignment between fuel assemblies and control rod drives. Tney also direct the flow of reactor coolant / provide gan=a and neutron shielding, provide guides for in-core i
instrumentation between the reactor vessel lover head and the fuel assemblies,
; support.the surveillance specimen assemblies in the annulus between the ther-l
- cal shield and the reactor vessel vall, and support the internals vent valves.
Tnese vent, valves are provided to relieve pressure generated by steaming in the core folleving a reactor coolant inlet pipe rupture so that the core vill re-main sufficiently covered with coolant. All reactor internal components can be removed fro: the reactor vessel to allow inspection of the reactor internals j _and the reactor vessel internal surface. *
. i. .
!i In anticipation of lateral deflection of the lower end of the core support as-
' sembly as a result of horizontal seismic loadings, integral veld-attached, de-flection-limiting _ spacer blocks have been placed on the reactor vessel inside vall. In addition, these blocks limit the rotation of the lover end of the core support assembly which.could conceivably result from flow-induced tor-sional loadings. The blocks _ allow free vertical movement of the lower end of the internals for thermal expansion throughout all ranges of reactor operating ccnditions, but in the unlikely event of a flange, circut'erential veld, or bolted joint failure the blocks will limit the possible core drop to 1/4 in, or less. Tne 'inal elevation plane of these blocks will be established near
.the same elevation as the vessel support skirt attachment to minimize dynamic leading' effects on the vessel shell or botto head. Freliminary calculations
~
findicate the 1: pact loading on the stop blocks for a,l/h in, core drop would
~
be approximately,5 g total. Block location and geometry will be evaluated and. determined to transfer this loading through the vessel support skirt to
,the' reactor building concrete. A significant reduction in impact loading can be achieved through- proper stop block design and detailed analysis. A l/h in.
ccre drop vill not allov -the 1cuer end of the. CRA neutron absorter rods to dis-engage frc their_ respective fuel assembly guide tubes if the CRA are in the full-out - position, since approximately 6-1/2'in. of rod . length would remain in g the fuel asse:bly guide tubes. 'A core drop of 1/h in. vill not result in a sig-D- nificant reactivity change. ' Tne core cannot rotate and: bind .the drive lines beesuse io:stiin cf the core support asse bly is prevented by the stor biceks. 3-65 0 0M'"'
r Qy The failure Of the core support shield and core barrel upper flanges, er re-lated flanges and other circunferential joints, is nct censidered credible on the basis of the conservative design criteria and large safety facters empicyed in the internals design. The final internals design vill te capable of with-standing varicus codbinaticns of forces and loadings resulting frc the static veight of internals (225,000 lb total, not including the plenue assembly which weighs 100,000 lb), core with control rod drive line (303,000 lt total), dy-naric Icad frc trip (l'O g gives 207,000 lb), seistic (0.10 g vertical cives 53,000 lb), coolant flcv hydraulic leading (230,000 lb), and cther related lead-ings. The algebraic su of this si plified loading case is 559,000 lb. This results in a tensile stress of about 585 psi in the ecre suppc:t shield shell, which is approximately 3 percent of the caterial yield strength. Final inter-nals cceponent veights, seistic analysis, dynanic leadings frc: ficv-induced vibration, detailed stress analysis with consideration for therral stress dur-ing all transients, and resolution of fabrication details such as shell rolling tolerances and* veld joint preparaticn details vill increase the stress levels
'is+oed
.. above. as a .'.4 .~c '. d --
e e d. e- .. s^ . .' *. e - 1 .* . +.k. . a c ^ -. =- c"--~~
--tr--- ^^--..*..*e
---r . "---
seet the stress requirements of the AS"I Code, Section III, during ncr a1 cper-ation and transients. The structur 1 integrity of all ecre supper; circunfer-ential veld Jcints in the internals shells vill be insured by cctpliance with the radiographic inspection require ents in the code abcVe. The seistic analysis vill include detailed calculations to determine the taxirus structural respcnse of the reactor vessel and internals. This analysis will be perferred as det-cribed in 3 1.2.k. s In the event of a major loss-of-coolant accident, such as a 36-in. diareter re-actor coolar:t pipe break near the reactor vessel outlet, the fuel assembly and vessel internals vould be subjected to dynamic leadings resulting frc: an cs-cillating differential pressure across the core. Scre deflection of the inter-nals structures would occur, but internals cc penent failure vill not occ6r. The occurrence of a less-of-coolant accident and resulting 1cadings will be evaluated during the detailed design period for the fuel assemblies and re-lated internals structural co ponents. The deflecticns and cove:ents described above would not prevent CRA insertien l because the control rods are guided throughout their travel. and the guide-to-fuel-assembly alignment cannot change regardless c: related cc ponent deflet-tions. CRA trip could conceivably be delayed comentarily as a result cf the oscillating pressure differential. However, the CRA travel tire tc full inser-tion suuld retain relatively unaffected as transient pressure oscillations are da:pened cut in approximately 0 5 sec. On this basis. the CRA travel time :: 2/3 insertion on a trip ccur.and vill be apprcximately 1.6 see instead cf the specified 1.LO sec. Also, this possible initial rincr delay in trip initiation would not contribute to the severity of the 1 css-Of-ccciant accident because at the initiation of CRA trip, the core would be suberitical frc voids. Material for the reactor internals boltire vill be subieeted tc ricid c.uality c'entrol require:ents to insure structural integrity. The bolts vill be in- . srected for surface flaw indications after all fabrica ict cc.eraticns have teen ccepleted. Torque values vill be specified for the final assemb13 :: de-velcp full-bc1 ing capability. All fasteners vill be lock-velded to insure as-sedbly integrity.
=
00m
-6c-
3 2.L.1.1 Flenu: Assembly The plenus asse:bly is located directly above the reacter core and is removed as a sirgle cc penent before refueling. It consists of a plenun cover, up;er grid, CFA guide tube asse:blies, and a flanged plenux cylinder with openings for reactor coolant cutlet ficv. Tne plenu: cover is a series cf parallel flat plates intersectirg; to,fc = square lattices with a perforated top plate and flange, and is attached to the plenu cylicder top flange. Tnree lifting lugs are provided fcr the plenum assembly hudlin;. Tne C;A guide tutes are velded to the plenue cover top plate and bolted to the upper grid. CRA guide assen-blies provide CM guidance and protect the CFA frc the effects of coolant cross-flow, and provide structural attachment of the grid assembly tc the plenut cover. Each C M guide assembly consists of an cuter tube hcusing, a counting flange, 12 perfcrated slotted tubes and fcur sets cf tube segments which are properly criented and attached to a series of castings to provide continucus guidance fcr the CFA full stroke travel. Design clearances in the guide tube vill ac-cc = cdate sc e degree of 21salig =ent between the CFA guide tubes and the fuel assemblies. Final design clearances will be established by tolerance studies and by the results of the Centrol Fed T ive Line Facility (CPIL) prototype tests. Preliminary test results are described in 3 2.h.3 5
~
Tne upper grid assembly consists of parallel flat bars intersecting to form square lattices. Tne bars are attached to a flange which is bolted to the plenut cylinder lover flange. Tne upper grid asse bly 1ccates the lover end _) of the individual CFA guide tube assembly relative to the upper end of the cor-responding fuel assembly. Locating key ays in the plenu asse bly cover flange engage the reacter vessel top flarge locating keys to align the plenu assembly with the reactor vessel, reactor closure head control rod drive penetrations, and the core support as-sembly. Tne bottes of the plenu assembly is guided by the inside surface cf the lover flange of the core support shield. 3 2.L.1.2 Core support Assembly Tne core support asse bly consists of the core suppcrt shield. core barrel, lcver grid assembly, flow distributor, thermal shield, in-ccre inst unent guide tubes, surveillance specimen holder tubes, and intern 11s vent valves. Static icads frc the assembled con;cnents and fuel asse:blies, and dynamic leads frc CFA trip, hydraulic ficv, thernal expansion, seistic disturbances, and Icss-cf-ecclant accident considerations, are all carried by t're core sup-port asseroly. Tne core support asse bly ecmponents are described as follows :
-. u% .,. -o e.
.v .-.w eu.. . e -, a Tne core support shield is a large flarged cylinder which cates with the reactor vessel opening. Tne top flange reste en a circu ."erential 1
- s. _ n. ;,_ 4,.,. u. w . ,. . . , . ..v~ . .m.ece.
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has two nozzle openings for reactor coolant outlet flow. The inside f] surface of the lower flange guides and aligns the plenum assembly relative to the core support shield. Three lifting lugs are pro-vided to handle the core support shield. These lugs are also used
+o handle the core support assembly.
L The core support shield outlet no::les are sealed to the reactor ves-sel outlet nozzles by the differential thermal expansion between the stainless steel core support shield and the carbon steel reactor ves-sel. The nozzle seal surfaces are fir.ished and fitted to a predeter-mined cold gap providing clearance during core support assembly in-i stallation and removal. At reactor operating temperature the mating metal surfaces are in contact to make a seal without exceeding allow-i' able stresses in either the reactor vessel or internals. Internals vent valves are installed in the core support shield cylinder vall to relieve the pressure generated by steaming in the core following a postulated cold leg (reactor coolant inlet) pipe rupture (see 3 2.h.1). i- b. Core Barrel i j The core barrel supports the fuel asser;blies, lower grid, flow dis-l tributor, and in-core instrument guide tubes. The core barrel con- {: sists of a flanged cylinder, a series of internal former plates { bolted to the cylinder, and a series of baffle plates bolted to the l- inner surfaces of the former plates to form an inner vall enclosing o . the fuel assemblies. Construction of the core barrel vill be sini-i ;s/ lar:to that of the reactor internals component developed by E&*n' for i the Indian Point Station Unit No. 1.
. Coolant flow is downward along the outside of the core barrel cylin-der and' upward through the fuel assemblies contained in the core bar-rel. A small' portion of the coolant flows upward through the space
, ;between the core barrel cylinder and the baffle plate vall. t-The upper flange of the core barrel cylinder is bolted to the mating lover flange of the core support shield assembly, and the lover flange
'is bolted.to the mating flange of the lover grid assembly. All bolts
~ vill be inspected.and installed as described in 3 2.h.1, and vill be
~
4 lock-velded after final assembly. Lifting lugs attached to the core barrel are provided for core support assembly handling.
- c. . Lever Grid Assembly i The lover grid assembly provides alignment and support for the fuel asse=blies, supports- the. thermal shield and flow distributor, and .
aligns the in-core instrument guide tubes with the fuel assembly in-strument tubes. The. lover grid consists of two lattice structures +- separated by short tubular columns surrounded by a - flanged cylinder. The top flange is bolted to the lower flange cf the core bsrrel. A perforated flat plate located =idway.between the two lattice struc-tures aids in distributing coolant flow. Ja
.\A)
[ .00E35 ' 3-68
, d. Flow Distributor i
The flow distributor is a perforated, dished head with an external flan 6e which is bolted to the bottom flange of the lower grid. The flow distributor supports the in-core instrument guide tubes and dis-tributes the reactor coolant entering the bottom of the core.
- e. Thermal' Shield A cylindrical, stainless steel, ther al shield is installed in the annulus between the core barrel cylinder and the reactor vessel inner vall. The thermal shield reduces the neutron and ga=ma internal heat generation in the reactor vessel vall and thereby reduces the result-ing thermal stresses.
The ther=al shield is supported on, positioned by, and attached to the lower grid top flange. The thercal shield upper end is positiened by spacers between the thermal shield and the core barrel outer cylin-der to minirize the possibility of thermal shield vibration. The thermal shield attachment is designed to avoid shear loads on fasten-ers. All fasteners are lock-welded after final assedbly.
- f. Surveillance Specimen Holder Tubes Surveillance specimen holder tubes are installed ~ on the core support assembly outer vall to contain the surveillance specimen asse:blies.
The tubes extend from the top flange of the core support shield to
} the lower end of the thermal shield. The tubes vill be rigidly at-tached to prevent flow-induced vibration. Slip joints at the in-termediate supports and top end of the assedblies accommodate axial motion caused by differential ther=al expansion,
- g. In-Core Instrument Guide TSbe Assembly The in-core instrument guide tube asseiblies guide the in-core instru-ment assamblies between the instruce penetrations in the reactor vessel tuttom head and the instrument tubes in the fuel assedblies.
' Minor horizontal' misalignment is accor=odated between the reactor vessel instrument penetrations and the instrument guide tubes assembled with the flow distributor. A perforated shroud tube, concentric with the instrument guide tube,' adds rigidity to the as-sembly and reduces the effect of coolant flow forces. Forty-six in-core instrument guide tubes are provided. The in-core instrument guide tubes are designed so they will not be affected by the core drop described in 3 2.h.1.
f
- h. Internals Vent Valves Internals vent valves are installed in the core support shield to prevent a ' pressure unbalance which cight interfere with ccre cooling
.following a loss-of-coolant accident. Under all normal operating
. conditions,.the vent valves will be closed. In the event of a loss-() of-coolant accident in the cold leg of the reactor loop, the valves 3-69 002E
~
F l t
- Q. Vill open to permit steam generated in the core to flow directly ci F to the leak and vill prevent the core from becoming = ore than 1/2 uncovered after emergency core coolant has been supplied to the reactor vessel. The design of the internals vent valve is shown in Figure 3-51.
t ?. Each valve assembly consists of a hinged disc, valve body with seal-ing surfaces, split-retaining ring, and fasteners. Each valve asse - bly is installed into a cachined counting ring, integrally velded in j the core support shield vall. The mounting ring contains the neces-i sary features to retain and seal the perimeter of the valve assembly. 1 Also, the mounting ring includes an alignment device to maintain the correct orientation of the valve asse=bly for hinged-disc operation. }h Each valve asse=bly will be remotely handled as a unit for re=cval i or installation. Valve component parts, including the disc, are of captured-design to minimite the possibility of loss of parts to the coolant syster, and all fasteners include a positive locking device. The hinged-disc includes a device for renote inspection of dise func-tion. The arrangement consists of lk-in dia: vent valve assemblies in-stalled in the cylindrical vall of the internals core support shield l (refer to Figure 3 h8). The valve centers are coplanar and are L2 in. above the plane of the reactor vessel coolant no :le centers. In cross section, the valves are spaced around the circumference of i, , the core support shield vall. a - i
,- , The hinge asse=bly consists of a shaft, two valve body journal re-ceptacles, two valve disc journal receptacles, and four flanged shaft
- l. journals (bushings). Loose clearances are used between the sh' art
)1. and journal inside diameters, and between the journal outside diam-l eters and their receptacles. ;- This feature provides eight loose rotational clearances to minitize any Sossibility of i=pairment of disc-free motion in service. In the ; event ? hat one rotational clearance should bind in service, seven loose rot. tional clearances vould remain to allow unhampered dise-free motion. In the worst case, at least four clearances cust bind or seize solidly to affect adversely valve disc-free motion.
In addition, the valve disc contains a self-align =ent feature so that the external differential . pressure adjusts the disc seal face to the valve body seal face. This feature minimites the possibility of in-creased leakage and pressure-induced deflection loadings on the hinge
. parts in service.
The external side of the disc is contoured to absorb the impact load of the disc on the reactor vessel inside vall without transmitting
~'
.. excessive impact loads to the hinge. parts as a result of a loss-of . _
coolant accident. 1 00?37 3-70
, ~
l-32.4.2 Fuel Assemblies 3 2.h.2.1 Description
- a. - General Description The fuel for the reactor is sintered pellets of low-enriched uraniu
dioxide clad in Zircaloy-h tubing. Tne clad, fuel pellets, end caps, s and the fuel support cc ponents for= a " Fuel Rod." T.c hundred and eight fuel rods, 16 control rod guide tubes, one instrumentation tube, I eight spacer grids, and two end fittings take up the basic " Fuel Assembly" (Figure 3-52). The guide tubes, spacer grids, and end fit-tings form a structural cage which contains the 203 fuel rods in a 15 x 15 array. The center position in the assembly is reserved for instrumentation . The remaining'16 locations in the array are pro-vided for the guide tubes which guide the control rods and provide the vertical support of the assembly. The complete core has 177 fuel assemblies which are arranged on a square lattice to approximate the shape of a cylinder. All asse -
- blies ~are identical in mechanical construction and interchangeable
- in the core and are designed to accept the centrol rod assemblies (CRA). The'. reactivity of the core under operating conditions is con-trolled by 57 CRA, of which 3 are xenon control rod asse blies.
These xenon control rod assemblies are identical in physical config-
~~
, uration to the CRA but have poison in the lover portion of the rod only. In the fuel assemblies containing no CRA, an orifice rod
_. / assembly. (Figure 3-53) or a burnable poison red asse:bly (Figure 3-54)
. is inserted into the upper ends of the guide tubes. Tnese assemblies I
- minimize guide tube bypass coolant flov. Tne lu= ped turnable poison I rod assemblies allow a lover boric acid concentration in the reactor i coolant,'thereby lowering the coderator temperature coefficient. 3e-
, cause of mechanical and geometric identity, the CRA, xenon control j_ rods, burnable poison rod asse=blies, and orifice rod assemblies are
- - designed to be interchangeable among fuel assemblies.
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Table 3-15 Puel Asse bly Components, Materials, and Dimensions Iter Material Eitensions, in. Puel UO2 Sintered 0 370 dia: Pellets Puel Clad Zircaloy-4 0.430 CD x 0 377 ID x 152.875 long Puel Ecd Pitch 0 568 Puel Asse bly Pitch 8.5S7 Active Pael Length 144 Overall Length =165
- I
- Centrol Rod Guide Zircaloy-4 0 530 CD x 0.015 vall Tube ,
Instrucentation Tute Zircaloy-4 0 530 CD x 0.402 ID Spacer Grid Inconel-718 Strips 0.016 thick i End Fittings Stainless Steel, Tp-304
- b. Puel The fuel is sintered and ground pellets of uraniu dioxide which are 1 fabricated frc: previously unirradiated caterial. These slightly enriched pellets are right circular cylinders with dished ends and a ground diaceter. The pellet ends are dished to minimize the dif-ference in axial expansion between the fuel and the cladding. The nominal density of the fuel is 93 5 percent cf theeretical.
Average design burnup of the fuel is 27,490 mwd /MTU. Peak desien turnup is 55,000 mwd /MTU. At the peak turnup, the fuel growth is cal-culated to te 9-1/2 volume percent by the nethod given in Reference
. 49 Radial growth of the fuel during turnup is accormodated by pellet i porosity, by radial clearance between the pellets and the cladding, 3
and by a small arount of perranent strain in the cladding. Eelev each fuel colunn is a support that axially locates the 'tcttet of the fuel eclunn and separates the fuel frc: the lcwer end esp. Means are provided at the top to maintain the fuel column in place during shipping and handling. 00239 2-72
M s. q-Fission gas release frc: the fuel is acco :cdated by voids within the fuel, by the radial gap between the pellets and the cladding, l l and by a void space at the top and bottc= ends of the fuel rods. 1
- c. Fuel Asse=bly Structu e
^ (1)- General , The fuel asse=bly shcvn in Figure 3-52 is the canless type in i which the spacer grids, end fittings, and the guide tubes for: the basic structure. Fuel rods are supported at each spacer grid by contact points integral with the valls of the fuel cell a boundary. The guide tubes are per=anently attached to the upper and lever end fittin6s tying the asse bly tcgether. The use of si=ilar =aterial in the guide tubes and the fuel reds results
, in _ tinicum differential thertal expansion. The fuel rods tottc=
cn the grid of the lever end fitting.
-(2) Spacer Grids Spacer grids are ecnstructed fro: strips which are slotted and fitted together in "eggerate" fashion. Each grid has 32 strips, 16 perpendicular to 16, which fer= the 15 x 15 lattice for the fuel rods. The square valls forced by the interlaced strips provide suppert fcr the fuel rods in two perpendicular direc-j tions. - Centact peints en the valls of each square opening are -
integrally punched di:ples in the strips. Y~ (3) Lover'End' Fitting J. The lever end fitting positions the _ assembly when inserted in [ theLlover ccre grid plate and supports the fuel asse=bly veight. The. lever ends of the fuel rods rest en the grid of the lever i_ end. fitting. Penetrations in the fitting are provided for at- [ teching the. control rod guide tubes and for access of the instru-4 4 tentation tube. a
-(4) Ucper End Fitting The upper end fitting positiens the upper end of the fuel as-se=bly.in.the upper ecre grid plate structure and p-ovides means
. for -coupling the handling equip:ent. An identifying nu:ber. cn each ' upper end fitting provides positive identification when
_ handling. r-j- An -internal hollev post, velded in the center of the end fitting
- provides ceans fcr retention of the~ orifice rod asse=bly and turnable poison red asse bly.
- Attached to;the upper end fittings are fcur holddern springs.
These springs provide a. pesitive' holddevr. targin to cppcse _ . .l~.. hydraulic ferees. r -- . i . 00240
't
- j. .
-13 ;) .-
m (5) control 1:cd Guide Tubes J The Zircaloy guide tubes prcvide continuous guidance to the con-trol rods within the fuel asse bly during operaticn and provide structural continuity for the fuel assently. Welded to each end j of a guide tube are flanged and threaded sleeves, which attach the tubes to the end fittings by lock-velded nuts. Eadial re-straint of the guide tubes is provided 13 the spacer grids. (6) Instrutentation Tbbe This Zircaloy tube serves as a channel to guide, position, and contain the in-core instrumentation in the center of the fuel as-
; sembly. The instrucentation string is guided up through the I lover end fitting and through the tube to the desired core ele-i vation. The instrurentaticn tube provides no structural support of the asce bly and is retained axially 13 the end fittings and radially by the spacer grids.
k 4 h i i
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3 2.1.2.2 4 Evaluation
- a. Fuel Rod Assembly (1) General The basi,s for the design of the fuel rod is discussed in 31.2.4 Materials testing and actual operation in reactor service with Zircaloy cladding have demonstrated that Zircaloy-k =aterial has sufficient corrosion resistance and techanical properties to i maintain the integrity and serviceability required for design burnup.
4 (2) Clad Stress and Strain The cladding of fuel rods is subjected to hydrostatic pressure, gradually increasing internal pressure, thernal stresser, vitra-tion, and to the effects of differential expansion of tha fuel i and cladding caused by thertal expansions and by fuel gro:eth due .
; to irradiation effects. In addition, the properties of the clad-i ding are influenced by ther al and irradiation effects, which are analyzed below.
- Stress analysis for cladding is based on several cense vative assu=ptions that take the actual nargins of safety greater than
, . 2. those calculated. For. exa:ple, it is assured that the clad with
-J the thinnest vall, the smallest fuel-clad gap, and the greatest ovality- permitted by the specification is operating in the re-gion of the core where performance requirements are cost severe.
Fission; gas release rates, fuel growth, arj changes in techani-cal properties with irradiation are based on a conservative evaluation of currently available data. Thus, it it 'inlikely I- that failure of the cladding vill result during operm lon. Pressure Effects Clad stresses due to external and internal pressure are consid-erab3v below the yield strength. Circunferential stress due to external pressure, calculated using those combinations of clad dimensions, ovality, and eccentricity that produce the highest
-stress, is shcvn in Table 3-16. The maximum stress of 33,000-psi compressicn, at the system design pressure of 2,500 psi, is the sus of 22,000-psi co pressive re=trane stress plus 11,000-psi co:pressive' bending stress due to ovality at the clad OD in
.L the expansion void and at the beginning of life. The taximum 1 stress in the heat-producing zone xis 32,000 psi at design pres- 'i sure and'27,000 psi at operating pressure. At this stress, the i caterial ay creep -enough .to allow an- increase in ovality until
- further creep is restrained by support fro: the fuel. Centact loads between fuel and cladding for this case are abcut 20 1t/
in. of length. {) 00M2 3-75
At the end of life, fission gas pressure ray exceed operating
) pressure when the fuel red is at operating te perature. The calculation of fission gas release is discussed in 3 2 3 2 3.h.
The value of 3,300 psi is used as a design internal pressure. The taxitut design pressure differential of 1,115 psi gives a resultant circunferential stress of 9,000 psi. This is about 1/h of the yield strength and, therefore, is n:t a potential scurce cf shcrt-time burst. The possibility of stress-rupture burst has teen investigated using finite-difference tethods to esticate the long-tice effects of the increasing pressure on the clad. The predicted pressure-time relaticnship produces stresses that are less than 1/3 of the stress levels that vculd produce stress rupture at the end cf life. Outpile stress-rupture data vere used, but the greater than 3: 1 cargin on stress is core than enough to account for decreased stress-rupture strength due to irradiation. s h t a V 00':; 2-76
O Table 3-16 L-} Clad Circumferential Stresses Ultimate Calc. Yield Tensile Stress, Strest., Stress, Coerating Candition psi nsi psi
' 1.
30L " - Operating at Design Pressure Conditions Total Stress (Me=brane + 3ending) Due ' to 2,500 psig Syste: Design Pressure Minus 100 psig Fuel Rod Internal Pressure Stress i Average Clad Te perature - Approxi-i =ately 625 F (Expansion Void) -33,000 h6,000 !~ i 2.- EOL - Maxiru Overnover Conditiens A .Syste Pressure - 2,185 psig Fuel Rod Internal Pressure - ' 3,300 psig. l Average Tc perature Through Clad j Thickness at Hot Spot - Approxi-cately 725 F Stress
- Pressure Stress Only(b) 9,000-Including 4,000 psi' Tnertal Stress 13,000 36,000 38,000 3 EOL - Shutdown
'Inmediately-After Shutdovn Ccnditions Syste- Pressure - 2,200 psig-Fuel Rod Internal Pressure -
1,750_psig: " (* Cladding vill-be specified with h5,000 psi minitu: yield strength and 10 per-cent rinitu: elongation,'toth at 650 F. 'Minicus roc: terperature strengths vill be approximately 75,000 psi yield strength (0.2 percent offset) and
~
S5,000 psi ultimate tensile strength.
- d"-- :(b) Cladding stresses due to P;el'svelling are discussed fcrther en r.nher
-pre _ c' 3 2. h.2. 2.
,_77 . 00?4. a..
a
, Table 3-16 (Ccntd)
Ultimate Calc. Yield Tensile Stress, Stress, Stress, Operating Condition psi usi usi Stress Average Clad Te:perature - Approxi-cately 575 F -4,003 L5,000 kg,co3 i 3 Hours Later F Conditions (50 F/h Pressuricer Cooldown Eate) Puel Rod Internal Pressure - 1,050 psig Syster Pressure - 6EO psig J Stress
. Average Clad Terperature - Approxi-cately 425 F 3,300 52,000 55,000
.g , The total prcduction of fission gas in the hottest fuel red as-f settly is tased on the hot rod average turnup of 35,000 :Gd/2.TJ..
The correspondin6 taximum design burnup at the hot fuel red mid-point is 55,000 igd /FIU. The fission gas release is based on temperature versus release fraction as shown in Figure 3-36. Puel temperatures are calcu-lated for small radial and axial increments. The total fissicn gas release is calculated by integrating the incre: ental re-leases. The maximun release and gas pressure buildups are deternined by evaluating the following facters for the rest conservative condi-tions: (a) Gas cenductivity at the end of life with fission gas present. (b) Influence of the pellet-to-clad radial gap and contact heat transfer coefficient en fuel tenperature and release rate. (c) Unrestrained radial and exial thermal growth of the fuel pellets relative to the clad. (d) Ect red local pea?.ing fac:crs. (e) Radial distributien of fission gas producticn in the fuel s_s
$ *tellets.
00W5 2-78
t (f) Fuel terperatures at reactor design cverpcVer. h The fuel temperatures used to deterzine fiesien gas release and internal gas pressure have been calculs:ed at the reactor cver-power condition (llh percent). Fuel terperatures, total free gas volute, fission gas release, and internni gas pressure have been i evaluated for a range of initial distetral clearances. This evaluation shcvs that the highest internal pressure results when the taxicum diacetral gap is assured tecause of the resulting i high average fuel temperature. The release rate increases rapidly with an increase in fuel terperature, and unrestrained axial growth reduces the relatively cold gas end plenu vcluxes. A con-
; servative ideal therral expansien todel is used to calculate fuel terperatures as a function of initial cold diacetral clearance.
Censiderably lover resistance te heat transfer tetween the fuel l and clad is anticipated at the end of life due te fuel fracture, swelling, and densification. 2ne resulting taxinur fission gas release rate is h3 percent.
; Colla _rse Marc _ ins Short-tite collapse tests have dencnstrated a clad collapsing pressure in excess of h,000 psi at expansic void taxinun ter-
, perature. Collapse pressure cargin is approxicately 1.7 Ex-trapolatien to het spct average clad te perature (zi25 F) indi-j cates a ecllapse pressure of 3,500 psi and a nargin cf 1.4,
) -
which also greatly exceeds require:ent. Ou:;ile creep ec11 apse
, # tests have deconstrated that the clad teets the lcng-tite (creep-I collapse) requirement.
l i< a Fuel Irradiation Grcuth and Fuel-Clad Differential Therral Expansion 1 i I The results of tests and the cperation of Zircale~v-clad UO-i fuel rods indicate that the rods can be safely cperated tocthe j point where total per anent strain is 1-1/2 percent, or higher, i in the terperature range applicable to FK?, cladding.(50) The allevable design strain is abcut 1 percent (3 1.2.4.2.c). Fuel rod operating ecnditions pertinent Ic fuel swelling consid-erations are listed telev for end-of-life ecnditicns. Burnup (Design Value), mwd /KrU 55,000 Minicus Fuel-to-Clad Gap (Eeginning cf Life), in. 0.00h5
. Pellet Kctinal Diameter, in. 0 370 l Pellet Density (Percent of Theoretical), %
- 93 5 Cladding (Zircalcy-4T, in. 0.0265 Wall
.~
' n 110 :er%..
-(,
m
The capability of Zircaloy-clad UO2 fuel in solid rod form to perfort satisfactorily in service has been de onstrated through operation of the SA-1 assembly in the Dresden and Shippingport cores, and thrcugh results of their supplementary development progra=s, up to approximately 40,000 igd /MTU. As outlined 'celov, existing ex'perimental inferratien supports the various individual design parameters and operating condi-tions up to and perhaps beyond the taximum design turnup of 55,000 igd /PTU, but not in a single expericent. E:vever, the S&W High Burnup Irradiation Progra: currently in progress does combine the primary items of concern in a single expericent, and the results will be available to contribute to the final design. Anplicatien of Exnerimental Data to Desien Adecuacy of the Clad-Fuel Initial Gap To Accortodate Clad-Pael Differential Themal Expansion Exneritental Work Six rabbit capsules, each containing three Zr-2 clad rods of 5-in. fue)51)len Test Reactor \ gth, were irradiated at power levels up in to the Westinghouse 24 kW/ft. The 94 percent theoretical density (TD) UO2 pellets (0.430 OD) had initial clad-fuel diacetral gaps of 6,12, and 25 mils.
. No dimensional changes were observed. Central melting oc-curred at 24 kW/ft only in the rods that had the 25 til initial gap.
Two additicnal capsules were tested.(52) The specimens , were similar to those described above except for' length and initiel gap. . Initial gaps of 2, 6, and 12 tils were used in each capsule. In the A-2 capsule, three 38-in.- long rods were irradiated to 3,450 IGdh2U at 19 kW/ft
~
maximu=. In the A-4 capsule four 6-in.-long rods were irradiated to 6,250 IGdhcu a,t 22.2 kW/ft maximu=. no -
. central melting occurred in any rod, but dia:eter in-j creases up to 3 mils in the A-2 capsule and up to 1.5 mils
- in the A-4 capsule vere found in the reds with the 2 til-initial gap.
. Application In addition to demonstrating the adequacy of Zircalcy-clad - UO2 pellet rods to operate successfuuy at the power levels of interest (and without central melting), these experi-cents demonstrate that the design initial clad-fuel gap of
~'=- 4 to 8 mils is adequate to prevent unacceptable clad dia -
- eter increase 'due to differential ther al expansion between
! the clad and the fuel'at beginning of life. A taxi un ! - local diametral increase of less than 0.001 in. is inci-I' cated for fuel rods 'having the minicus initial gap, cperat-l- _h; ing a- - " " xitur'cVerp:ver ecndition. !~ d 00?qy 3-a0 t L.
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in the 93 5 percent pellets is filled. Frc: that time cn, svelling is assumed to take place at 0.7 py-cent LV/1020 f/cc until the taximum turnup of 13.6 x 10 #b f/cc (55,000 9 d/MTJ) is reached. Studies of clad strain at various gaps indicate that the rod with the minicum cap experiences the greatest clad strain in spite of its improved gap conductivity. Clad permanent strain reaches a maximum at the end of life, and is 0.7 percent for nominal density fuel. Clad strain for fuel rods with taximum density allcved by the specificatien vill also meet the design's maximum allevable permanent strain. Fuel Svelling Studies at E&W Erperixental fuel swelling studies under inpile ecnditicns siculating large reactor environnents are under way. Para - eters centributing to swelling are turnup, heating rate, fuel density and grain size, and clad restraint. These are being studied systematically by irradiating a series of capsules containing fuel rods. Test variables are shown in Table 3-17, and the prcgrat's schedule is given in Table 3-18. See also 3 3 3 3 3 Test variables include heat rate, turnup, clad thickness, and fuel-to-clad gar. Festirradiation exacinaticn vill
=d include investigaticn of dimensional changes, retallo-graphic examination of fuel and cladding, ficsion gas release correlations with test conditiens, and other re- . lated observaticns.
i a 9 U 00'"19 3-se
f . W . ] Table 3-17 BfW Ifigh Burnup Irradiation Program - Capsule Fuel Test Burnup ilant Rate , Identification Diametral Clad Irradiation IND/MIU Tisalons/cc Irrndiation Time Initini, Finni, Gap, Thickness, Capsule Fuel Bod Facility x 10-3 x 10-20(1) enlendar months (2 kv/ft kv/ft mils - mile B-1 B-1 RS-3 10 25 4 18 17 5 h-5 25 B-2 2.5 17 5 7-8 25 B-3 25 17 5 7-8 15 B-2 B-20 RS-b 17 3.8 10 18 16.9 Powder 25 B-19 26 6.5 16.9 h-5 25 B-3 B-7 ns-6 30 .5 11 18 16.1 h-5 25 B-8 7.5 16.1 7-8 25 D-9 75 16.1 T-8 15 Bb B-10 RS-1 b5 10.05 17 18 1h.9 Powder 25 B-11 11.25 1h.9 h-5 25 B-5 IL13 DS-5 55 13.75 21 18 1h.1 h-5 25 (., 1414 13.75 14.1 7-8 25
- Ibl5 13.75 14.1 7-8 15 W IL6 B-31 RS-2 70 15.63 26 18 13.3 Powder 25 B-17 17.5 13.3 7-8 25 lb 7 B-5 PG-5 80 17.87 30 18 12 5 Powder 25 ilk 2G. 0 12 5 7-8 25 B-8 B-22 It!-1 30 7.5 11 21-1/2 20.5 h-5 25 B-23 7.5 20 5 7-8 25 B-33 7.5 20 5 T-8 15 IL9 B-25 Rir? 55 13.75 21 21-1/2 19.3 4-5 25 lbr6 13.75 19.3 78 25 B-27 13.75 19.3 7-8 15 b-10 B-28 n r.- 3 80 20.0 30 21-1/2 17.8 T-8 25 B-?9 20.0 17.8 T-8 25 B-30 20.0 17.8 7-8 15 B-11 B-2h RL b 70 17.5 26 21-1/2 16.5 7-8 15 C B-3h 17.5 16.5 7-8 25 C B-35 17 5 16.5 7-8 25 11-12 Ibl6 DG-3 65 1h.50 24 18 13.3 Powder 25 N Ib32 16.2h
- 13.3 7-8 15 C
Ilaned on 200 Mcv per finston. Pened on 80 per cent react ar e rrictency.
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4, Effect of Zircaloy Creep The effect of Zircaloy creep on the arcunt of fuel red growth due to ibel swelling has been investigated. Clad creep has the effect of producing a nearly constant total pressure on the clad ID by permitting the clad dia eter to increase as the fuel diareter increases. based on cut-of-pile data,(>E) 1 percent creep will result in 10,000 h (corresponding apprcxi-rately to the end-of-life diaretral swelling rate) f~ct a stress of about 22,000 psi at the 2720 F average terperature through the clad at the hot spot. At the start of this high swelling period (roughly the last 1/3 of the core life), the reacter coolant syste: pressure vould =cre er less be balanced by the rod internal pressure, so the total pressure to produce the clas stress of 22,000 psi vould have to cc=e free the fuel. Contact pressure vould be 2400 psi. At the end of life, the rod internal pressure exceeds the syster pressure by about 1100 psi, so the clad fuel contact pressure vould drop to 1300 psi. Assuring that irradiation produces a 3:1 increase in creep rates, the clad stress for 1 percent strain in 10,000 h would drop to about 15,000 psi. Contact pressures veuld be 1800 psi at the beginning of the high swelling period, 7JO psi at the end of life. Since the contact pressure was assired to be 825 psi in calculating the contact coefficient used to de-ter=ine the fuel pellet therral expansics, there is cnly a short period at the very end of life (assuring the 3: 1 increase in j creep rates due to irradiation) when the pellet is slightly hotter than calculated. The effect of this would be a slight increase in pellet therral expansion and therefore in clad strain. Consicering the improbability that irradiation will actually increase creep rates by 3: 1, no change is anticipated.
- b. Overall Assembly (1) Assurance of Centrol Rod Asse bly Free Motion The 0.060 in. diametral clearance between the control rod guide tube and the control red is provided to eccl the centrcl rod and to insure adecuate freedc to insert the control rod. As indicated below, studies have shown that fuel rods will not tov sufficiently to touch the guide tute. Thus, the guide tube vill not undergo deferratica caused by fuel rod bcring effects.
Initial lack of straightness of fuel red and guide tube, plus other adverse tolerance conditions, conceivably could reduce the 0.093 in. nominal gap between fuel rod and guide tube to a minimum of about 0.055 in., including a:plification of bowing due to axial frictits leads frce the spacer grid. The maxi =u expected flux gradient of 1.176 across a fuel rod vill produce a terperature difference of 12 F, which will result in a thernal bov cf less than 0.002 in. Under these conditicns, for the fuel rod to touch the guide tube, the thernal gradient f across the fuel zod diareter vculd have to be ce the crier cf v 100 F. s
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The effect of a DIi3 occurring on the side of a fuel rod adjacent to a Euide tube would result in a large terperature difference. In this case, however, investigation has shown that the clad terperatu"e vould be so high that insufficient strength would be available to Eenerate a force of sufficient ragnitude to cause a significant deflection of the guide tube. In addition, the guide tube vould experience an opposing gradient that would resist fuel rod bowing, and its internal cooling would maintain temperatures =uch lower than those in the fuel rod cladding, thus retainin5 the Euide tube strength. (2) Vibration The serie pirical erpression developed by Burgreen( ) vcs used to calculate the flow-induced vibratory amplitudes for
. the fuel assembly and fuel rod. The calculated a plitude is 0.010 in for the fuel asse=bly and less than 0.005 in. for the fuel rod. The fucl rod vibratory a plitude ccrrelates with the measur'ed a:plitude obtained fro a test on a 3 x 3 fuel rod assembly. In order to substantiate what is believed to be a conservatively calculated a plitude for the fuel asse=-
bly, a direct ceasurement vill be obtained for a full sine prototype fuel asse:bly during testing of the asse=bly in the Control Rod Drive Line Facility (CRDL) at the NN Research Center, Alliance, Ch5 o. (3) Demonstration In addition to the specific iter.e discussed above, the overal'1 rechanical perfor:ance of the fuel asse bly and its individual
.; components is being demonstrated in au extensive experimental
. pro 6ra in the CRDL. .
'3 2.4.3 Control Rod Drive syste i
l'3 2.h.'3 1 Description i The control rod drive syster includes drive rechanists which actuate control rod assemblies and xeno.n control red asse blies, drive controls, power sup-
. plies, position indication, operating panels and indicators, safety devices, i enclosures,' housings, and countings. Criteria appl 1 cable to drive rechanisms
.for both. control rod assemblies and xenon control rod asse:blies are 5 1ven
- in -3 2.4.3 1.1. Additional recuirements for the rechanisms which actuate ,
only control rod assemblies are givin in 3 2.4.3 1.2. , 3 2.4.3 1.'1 General Design Criteria -
- a. Single Failure No-single failure shall inhibit'the protective action of the con-
~
trol rod drive syster. The effect of a single failu-e shall be 11=ited to one control rod. drive. k~ [ oo?ss 3-66.
- b. Uncontrolled k'ithirawal_
- c cingle failure er sequence of dependent failures chall cauce un-centro 11ed vii wal of an centrcl rod acrcnbly (CRA).
- c. r,uiptent Renova 2 The disconnection of plug-in connectors, =cdules, and cutasse:blies frc the protective circuite chall be annunciated er chall cause a reactor trip.
- d. Fosition Indication Continuous position indication, as well as an upper and lower posi-tion limit indication, shall be provided for eaci. control rod drive.
The accuracy of the position indicators shall be consistent with the tolerance set by reactor safety analytic.
- e. Sycte Monitcring The control rod drive control syste: shall include provisions for monitoring conditions that are important to safety and reliability.
These include rod position deviation and power supply voltage.
- f. Drive Speed The control rod drive control cystem shall provide for single uni-
-s/ for: speed of the cetbacic=. The drive controls, or techanis:
, and motor corbination, cnall have an inherent cpeed limiting fea-ture. The speed of the techanic: shall be 30 in./=in 6 percent of the predeter=ined value for both incertion and withdrawal. The withdrawal speed sh<111 te limited so as not to exceed 25 percent overspeed in the event of speed control fault,
- g. Mechanical Stops Each control rod drive shall have positive techanical stops at
' both ends of the strcke or travel. The steps chall be capable of receiving the full cperating force of the rechanists without fail-ure.
3 2.4.3 1.2 ' Additional resign Criteria The following criteria are applicable cnly to the techanics: which actuate control rod assemblies.
. a. CRA Positioning The control rod drives shall provide for contrclied withdrawal or insertion of the control rod acsecblies (CRA) out of, cr into, the reacter core to establish and hold the power level required.
The drives are also capable of rapid insertien er trip for trercency t s-3 reacter ccnditicns. 00'M 3-87
- b. CRA Trip The trip command shall have priority over all other cc=nands. Trip action shall be positive and nonreversible. Trip circuitry shall provide the final protective actics and shall be direct-acting, in-car mini =um delay, and shall not require external pcwer. Circuit-interrupti 2 devices shall not prevent reacter trip. Puses, where used, shall be provided with blevn indicatcrs. Circuit breaker position informatica chall also be indicated.
- c. Grcup Withdrawal 1
The control rod drive system allows only two cut of three regulat-ins CRA Sroups to withdraw at any ti=e subject to the conditicns described in 7 2.2.1.2. 32.4.3.2 Contrcl Red Drive Mechanists I The control rod drives provide for controlled withdrawal or insertien cf the control rod assemblies out of or into the core and are capable of rapid in-sertion or trip. The drives are hernetically sealed, reluctance motcr-driven i screw units. The centrol rod drive data are listed in Table 3-19 i Table 3-19 Centrol Rod Drive Design Data 16 I Item Data t i Nu:ber of Drives 57 i Type Her etically Scaled, Reluctance
' Motor-Driven Screw I
Location Top-Mounted i Direction of Trip Down Velocity of Ucrmal Withi-aval and Insertion, in./rin 30 1',aximum Travel Time for 2/3 Trip Insertion (93 in.), s 1.40 j Length of Stroke, in. 139 4 f Design Pressure, psi 6 2500 Design Tarperature, F 650
-5 s The drive techanis censists of a notor tube which houses a lead cerev and s-its rcter asse bly, and a tuffer. The end cf the retor tute is cirsed by a
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cap and vent assembly. A totor stator is placed dcun cver the cotor tube precsure vessel, and position indication switches are arranged cutside the motor tube extension. The centrol rod drive output elenent is a translating screw chaft which is coupled to the centrol rod. The screv is driven by an anti-friction nut ele =ent which is rotated magnetically by a notor stater located cutside the pressure boundary. Current impresced on the stator cauce: the ceparable nut halves to engage; a =eebanical cpring causes them to direngage the screw in the absence of a current. For rapid insertien, the nut ceparater to release the screw shaft which then falls into the ccre by gravity. A buffer within the upper housing decelerates the falling asse bly to a 1cv speed a chcrt distance above its full-in position. The final deceleratien is acect=cdated by the down-step buffer spring. This recha*4"- d- orpcrates proven principles and caterial cc binaticas and is based en extensive analytical, develeprennal, design, test, and nanufac-turin6 experience cbtained ever the years for Shippingpcrt and the Nucicar Navy. - The control red drive is ch vn in Figures 3-55 and 3-56. eubasce blies cf the control rod drive are described as follows:
- a. Mater Iube The noter tube is a three-piece velded asce bly cesigned ani ranu-factured in accordance with the requirements of the ASME Code, J
Section III, for Class A nuclear pressure vesseln. The noter tube vall between the rotor asseibly and the ctater is ccnctructed cf ragnetic raterial in crder to present a stall air Eap to the Otcr. This regicn of the cotor tube is of lov alloy cteel clad cn the inside diameter with stainless steel er with Incenel. The upper end of the rotor tube acts only as an enclosure for the withdrawn lead screw; this end is made of a stainless steel and is transition-velded to the upper end of the lov alloy =cter section. Tne icver end cf the lov alloy tube cection is velded to a stainlesc centrif-ugal casting which is flanged at the face in centact with the
. vessel's ccatrol rod no:cle. Double Easkets with a test port be-tween are used at the connectica between the noter tube and the reactor vessel.
- b. Meter The notor is a synchroncus, canned reluctance unit with a slip-on stator. The rotor element is described in Paragraph (f) belev.
The stator is a 48-slot, four-pole arrangement with water cooling coils wound on the outside of its casing. The stator is encaptu-lated after vinding to establish a hermetically sealed unit. It is six phase, star-connected for cperaticn in a pulec-stepping rede, advancing 15 rechanical degrees and 30 electrical degrees per tiep. The stater assembly is acunted over the ctor tute houcing ar chcun in Figure 3-56. ts- 8 O O '9 -)n n% 3-59
- c. Cap and Vent Valve The upper end of the retor tube is closed by a cap containing a vapor bleed pcrt and vent valve. The valve and bleed per and the cap-to-noter tube closures are arranged to have double seals.
The cap is retained by a bolting ring threaded to the outside of the ctor tube. The retaining bolts are accenbled to the belting ring so that'they cannot be dropped. The bolts are made long to as to be clastic enough to provide pccitive ceal prelcad at any assenbly terperature fro: 20 to 650 F. The minicus prelcad is equal to the 3750 psig proof pressure force.
- d. Actuator i The actuator consists of the translating screw chaft, its rotating nut asse bly, and a torque restraint for the screw. The actuator travel ir about 12 feet,
- e. Screv Shaft 1
The screw shaft has a lead of 0 750 in. The thread is a double
! entry with a spacing of 0 375 in. Thread lead error is held to , 0.0005 in. raxinus in any 6 in., so that good lead sharing is obtained with the roller nuts. The thread for: is a dified AELS with a flank angle that allcvs the roller to disengage without lifting the screw.
s h
- f. Rotor Assembly The rotor assembly is a pair of scissors arms containing ball bearing-scunted threaded rollers skewed at the lead screw helix angle so as to engage the screw thread. The scissors arms are pivoted on a hollow support re ber (rotor tube), so that the rollers can alternately engage or disengage with the screw thread. The rotor tube re=ber is rounted on ball bearings cupported by the rotor
, tube.
- g. Roller Hut The roller nut asse:bly is a cylinder with circunferential rings matching the for anc spacing of the screw shaft thread. It cen-tains integral, angular contac*,, ball thrust bearing racer at each end thrcuch which it is supported by a chaft fixed at the c:rev helix angle to the scissors ar:s. Two such asce bliec are counted to each scissors art, so that the four together form a complete nut to catch the screw shaft.
- h. Scissors Arms The scirscrs arts are ragnetically acted upon by notcr ::ator.
The end of the arts belcv the pivot support the : ier nute and a a, separating spring. Carrent in the rotor cauces the arts to :ve
"<) radially elecer to the notor be's inner vall, thereb," c;;1cing 00'7a 3 co
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tions cate and crient the tcrque restraint tube and the otor tube just belev the cap. The male serraticas are nachined on a shculder
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O l 1 upper section which produce the damping pressure drep. The number y, e of holes presented to the buffer chamber is reduced as the red
=cves into the cere, se that the da ping ccefficient increases as the velocity reduces, thereby providing an apprcximately unifer:
deceleration. A large helical spring buffer is e=plcyed te take the kinetic energy cf the drive line at the end of the water buffer streke. The spring tuffer accepts a five-fce; per second inpact velocity cf the drive line and control rod with an instantanecus cvertravel of one inch past the normal dcwn sicp. The inclusica of this spring buffer permits practical clearances in the water buffer.
- 1. Lead Screw Guide The lead screw guide bushing acts as a primary thermal tarrier and as a guide for the screw shaft. The tushing acts as a prirary therral barrier by alicving Only a small path for free convection cf water between the rectanis ~and the closure head noc le. Fluid tenperature in the rechanis: is larEely governed by the ficw of water up and devn through this bushing. The diametral clearance between screw shaft and bushing is large encugh to preclude je=:ing the screw shaft and small enough to hold the free convection to an acceptable value. In order te obtain trip travel tires of acceptably exall values, it is necessary to provide an auxiliary flow path arcund the guide bushing. The larger area path is necessary to reduce the pressure differential required te drive water into the rechanis: to equal the rev displace ent. The
,;) auxiliary flow paths are closed fcr a:all preccure differentials (several inches of water) by gravity relief valves which prevent the convective flows, but open fully during trip.
(~ . -. w g,q~: /U- 4 s s
l i i l 7 c. Position Indication and Controls i.~- The position transducer consists of a series of ragnetical.ly operated
' reed switches =cunted in a tube parallel to the motor tube extension.
Each switch is her=etically sealed. Switch contacts close when in
.; proximity to a pe =anent magnet mounted on the upper end of the lead
{ screv extension. As the lead screw (and the control rod assembly) moves, switches operate sequentially. The closures are detected, and an analog voltage of position is produced at the output of the position con-verter. The output is utilized for control and reter indication. The accuracy of position indication is about f1-1/2 in. Similar switches are used as travel limit transducers, which are energined i '~ by the same internal cagnet. Switch closure can be used to operate
'i ala m and stop rod motien devices directly.
A position indication is provided by driving a pulse-stepping motor, which in turn drives a position indication potentio=eter. This out-put drives a conventional panel meter. The accuracy of position indication is about 12-1/2 in.
}- *
- n. Motor Tub'e Design Criteria
'f. The motor tube design co= plies with Section III of the ASG Boiler
] and Pressure vessel Code under classification as Class A vessels.
l The operating transient cycles, which are considered for the stress j
' analysis of the reactor pressure vessel, are also considered in the i motor tube design. "
l Quality standards relative to =aterial selection, fabrication, and inspection are specified to insure safety function .of the housings essential to accident prevention. Materials conform to ASTM or ASG, Section II, Material Specifications. All velding shall be perfor=ed by personnel qualified under ASG Code, Section IX, Weldins Qualifications. These design and fabrication procedures establish quality assurance of the asse=blies to contain the reactor coolant safely at operating te=perature and pressure, , In the highly unlikely event that a pressure barrier component or the control rod drive asse=bly does fail catastrophically,
- ie, ruptured completely,' the following results would ensue
l4 ._(1) Control Rod Drive Nozzle The assembly would be ejected upward as a missile until , it was stopped by the missile shield over the reactor.
- This upward motion would have no adverse effect en adja-cent asse=blies.
m 0060 ' 2-93
, (2) Motor Tube The failure of this corpenent anywhere above the 1cwer flange would result in a missile like ejection into the missile shielding over the reacter. This upward notien would have no adverse effect on adjacent rechanists.
3 2.4.3 3 Centrol itad Drive Control Syster (Centrol Package) The control syster for the control rod drive is designed to energine and position the control rod drive, provide a reactor trip, indicate the control rod assembly (CRA) position in the core, and indicate ralfunctions in the
., syster. The control syster consists of:
A. Systec Centrol
- 1. Individual and Group CRA Ccntrci (Cperator's Fanel)
- 2. Position Indication 3 Autcratic Sequencing
- 4. Position Deviation Monitors B. Power Supply (Motor Controller)'
- 1. SCR Prograrrer (CRA Speed Standard)
- 2. SCR Banks 3 CRA Grcuping panel J
- 4. Transfer Control C. Triu Figure 3-57 depicts in block diagra: fer: Items 3 and C with cct:and inputs fro: Ite: A.
The reactor operator is provided with en cperator panel and controls which perrit manual or autcratic group operation, manual single rod operatien, group sequencing and position indicarica. All =anual cerrands, including operator-initiated trip, are made frc the opera:cr's panel. Position indication is provided on both individual r:ds and rod groups. Iniividual position =cters indicating percent withdrawn are visible frc the operator's panel. Four Ercup pcsitien reters are prcvided at the cperator's panel. These group reters indicate the position cf either the four safety rod groups or the regulating grcups, whichever is selected by the reter selector switch. Automatic sequencing of the regulating grcups is provided. The sequencer pro- , vides overlapped withdrawal and insertion of Crcups 5, 6, and 7 within the limits of 7 2.2.1.2. This sequencing is provided in both autcratic and asnual
=0 des of centrol. Sequencer logic is derived frc pcsitien transducer and limit cwitch signals.
.o The syster control provides the 1ccic tc cc :and the prcper Crcup ;;wer suppl;..
't-# The grcup pcwer supply prcvidec d-c p:ver frc a thrwe-:'r ace ::urce in: L;;1ic:
,_ct 00261
it as directed by the program er controlled SCR to the CRA nechanism. The pcwer supply consists of an SCR progra==er, SCR gate drivers, SCR tanks, trans-fer relays, input power transforrers, and rod group patch panels. Input ccanands of in-hold-out are received by the progranner, which in turn generates the gating sequences for the SCR banks. The programmer consists of a synchronous =otor.cperating on 60-cycle, a-c power, driving a ecded disk thrcugh a light bear. The coded light bear driver photo detectcrs, and the photo detectors drive S R gate driver arplifiers, which in turn gate the SCR. The SCR banks apply a steady voltage to successive motor windings and are line-commutated by the input a-c power. The rotor is 6-phase, star-ccnnected and produces 15 degrees of techanical rotatica per switching cycle. The group power supplies contain redundant SCR banks, each fed frca a dif-ferent power source but driven frer a ec= cn but dual-channel progranter. Eight grcups of drives are in the rod drive syster, each having its evn power supply. A ninth pcwer supply ic prcvided fer single rci ccncrol and as an cperational spare. - Any rod cay be operated in the single rod mode by transferring to the ninth power supply. Reactor trip is initiated by de-energicing two circuit breakers supplying control roi drive power or two contactors supplying SCR gate power. Ecth circuits have two devices in series.
, j The reacter protection syste trips the power circuit breakers and ccntrol circuit contactors through two out of four logic as thcwn in the block dia-gras (Figure 3-57). .
3.2.4.3.k Control Rod Drive System Evaluation E. Design Criteria The system vill be designed, tested, and analyced for cc pliance with the design criteria. A preliminary safety analysic of the ccntrol rod drive rotor control subsyster was ccnducted to de-terzine failures of logic functions. It was concluded that no cingle failure in any CRA control would prevent CRA inserticn, n:r cause inadvertent CRA vithdrawal of another CRA cr CRA grcup.
- b. Materials Selection Psterials are selected to be ccrpatible with, and operate in, the reactor coolant. Certified mill test reports containing chemical analysis and test data of all asterials exposed to the reactor system fluid vill be provided and naintained for the control rod drives. Certificates of cc=pliance for other naterials and cca-penents shall also be provided.
)
()C'i$t325 1 l I <hk L
- c. Relation to Design Temperature All parts of the control rod drive exposed to reactor coolant are designed to cperate at 650 F, althcugh it is expected that all parts vill operate considerably cocier. Some tests have been ect-pleted, and addinional tests are planned to determine the cperating terperature gradients throughout the drive techanis: during all phases of operaticn. These tests will also provide an indica:1cn cf the arount of convection that takes place within the water space of the mechanist. The more significant terperature changes vill be caused by displacement of reactcr coolant in and out of the rechanis: vater space as the drive line is raised and levered,
- d. Design Life The expected life of the centrol red drive control syster is as e ,__ _. . s .
.n .
(1) Structural pcrtions, such as flanges and pressure housings, have an expected life of h0 years. (2) Moving parts, such as lead screw and roller nuts have an expected life of 20 years. (3) Electronic ccatrol circuitry has an expected life of 20 years. 3 2.4.3 5 Centrol Ecd Asse:bly (CRA) Each control rod assembly is cade up of 16 centrol rods which are coupled to a single Type 304 stainless steel spider (Figure 3-58). Each centrol rod consists of an absorber section of silver-indium-cadrium poisen clad with cold-worked, Type 304 stainless steel tubing and Type 304 stainless steel upper and lover end pieces. The end pieces are velded to the clad to for a water and pressure-tight container for the poison. The centrol reds are loosely coupled to the spider to perrit maximum conformity with the channels provided by the guide tubes. The CEA is inserted through the upper end fitting of the fuel assembly, each contrcl rod being guided by an in-cere guide' tube. Guide tubes are also previied in the upper plenu= assembly abcVe the core so that full length guidance cf the control rods is provided throughout the stroke. With the rea'eter assembled, the CEA cannot be with-irawn far enough to cause disengagement of the control rods frc the in-core guide tubes, pertinent design data are shcvn in Table 3-20. () Oe .R)mt% 3-9b
7 1 O i
, Tuble 3-20 Control Rod Asce bly Design Data Ite: Data i
Uu-ber of R d Asse blies 49
; Nu=ber of C trol Rods per Asse bly 16
'l l Outside Dia:eter of Centrol Rod, in. O.440 i, i Cladding Thickness, in. 0.019 I Cladding Material Type 30k ES, Cold-Werked poisc Material 60% Ag, 159 In, 5% cd length of poisen Section, in. 134
, Stroke of Control Rod, in. 139 This type cf CRA has been developed under the USAEC large Reactor Development progra: and offers the following significant advantages:
~} a. More uniforn distribution of absorber thrcuchout the ccre volune.
- b. Ehorter reactor vessel and shorter internals owing to elimination of control rod followers.
- c. IcVer reactor building requirements cving to reduction of reactor coolant inventory.
- d. 3etter core power distribution for a given CRA vorth.
A CRA prototype sinilar to the B&W design has been extensively tested (EI) at reactor ter;erature, pressure, and flow conditions under the IRD program. The silver-indiu=-cadriu poison caterial is enclosed in stainless steel tubes to pr: vide structural strength to the control rod asse blies. These rois are designed to withstand all cperating loads including these rcruiting frc: hydra" ^ "--^es, thernal gradients, and reactor trip deceleraticn. The cladding of the poison section also prevents corrosion and eliminates possible silver contamination of the reactor coolant. The ability cf the poison clad to resist collapse due to the syste pressure has been dc---e- = ad by an extensive collap;e test procra cn cold-v:rked stainless steel rods. The actual collapse cargin: are hicher than the re-quirenents. J (){}:'di4: 2 c,a (-
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i I 1 , , Table 3-21 Xenon Control Rod Assenbly Desi6n Data + Ite= Data i Nur.ber of Rod Asse=blies 8 t Number of Xenon Control Rods per Assenbly 16 1 Outside Dic=eter of Xenon Control Rod, in. O.hhO j Cladding Tnickness, in. 0.019 Cladding IM:erial Type 30h SS, Cold-Worked Poison laterial 80% As,15% In, 5% Cd Length of Poison Section, in. 36
~
Stroke of Control Rod, in. 139 3 2.4.3 7 Eurnable Poison Rod Asse=bly (EPPA) , Each burnable poison rod asse=bly consists of 16 burnable poison rods which , j are coupled to a single Type 3ch stainless steel spider (Fi6ure 3-5h).
- burnable poison rod consists of clad Bhc in Al 023 . The end pieces are Each i velded to the clad to for= a vater and pressure-tight container for the burn-
. able poison. Tne EPPA are Suiced by the fuel asse=bly guide tubes, and coupled with the fuel asse=bly by =eans of a positive coupli 5 =echanis= provided on the burnable poison rod spider and the fuel assenbly hold-down latch. In addition to their nuclear function, these EPRA also mini =1:e guide tube bypass coolant flov. Pertinent design data are shown in Table 3-22.
Table 3-22 Burnable Poison Rod Asse=bly Design Data 4 i Ite Data Nu=ber of Rod Asse=blies 72 Ita=ber of Earnable Poison Rods per Asse=bly 16 - l 1 ,'Outside Dia=eter of Burnable Poison' Rod, in. O.h30 i Cladding Tnick=ess, in. 0.035-
, l' Poison tuterial EkC in Al 023 i Length of P icon Section, in. 126 00266 3-9: . ,- . , , , , _ _ -
--..Ane
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e
33 tests AI:D II:SPECTIO!!S 331 I.JCLEAR TESTE A';D II SFECTICI: 3 3 1.1 Critical Experiments An experimental program (58-60) to verify the relative reactivity worth of the CRA has recently been ccepleted. Detailed testing establishei the worth of the CRA under various conditions similar to those for the reference core.
; These parameters include control rod arrangement in a CRA, fuel enrichtents, j fuel element geometry, CRA materials, and soluble boron concentration in the 1
toderator. I Gross and local power peaking vere also studied, and three-ditensicnal power-peaking data were taken as a function of CRA insertion. Eetailed peaking data vere also taken between fuel asseiblies and around the water holes left by withdrawn CRA. The experimental data are being analyzed and vill teccre part of the experitental bench tark for the analytical codels used in the design. l L 3 3 1.2 Zero Power. Approach to Power. and Power Testing .l
- Boron worth and CRA vorth (including stuck-CRA vorth) vill te determined by e physics tests at the beginning of each core cycle. Recalibration of boron worth and CRA vorth is expected to be performed at least once during each core cycle.' Calculated values of boron worth and CRA vorth will be adjusted . ,gh to the test values as necessary. The boron vorth and CRA vorth at a given time in core life vill be based on CRA position indication and calculated data as adjusted by experimental data. ,
The reactor coolant vill be analyzed in the laboratory periodically to deter-tine the baron concentration, and the reactivity held in boron vill then be l calculated from the concentration and the reactivity worth of boren. 1
. The nethod of raintaining the hot shutdown targin (hence stuck-CRA targin) is I, related to operational characteristics (load patterns) and to the power-peaking I restrictions on CRA patterns at power. The CRA pattern restricticns vill in-sure that sufficient reactivity is always fully withdrawn to provide adequate shutdown with the stuck-CRA margin. Fover peaking as related to CRA patterns and shutdown targin vill be monitored by reactivity calculations.
Operation under power conditions vill normally be conitored by in-ccre instrumenta-tion, and the resulting data vill be analyzed and ccrpared with cultidiren-sional calculations to provide support for further power etcalations. I j 332 THERMAL AND HYDRAULIC TESTS AND II;SPECTIOI: I 3 3 2.1 Reactor Vessel Flow Distribution and Pressure Eron Test A 1/6-scale codel of the reactor vessel and internals will te tested to evaluate :
,g^ a. The flow distribution to each fuel assembl:. of the react:r ccre o .- and to develop, if necessary, devices rer ired to pr;'use M.c desired flow distribution.
00?.67 3-100
4 l
; 'T - b. Fluid mixing between the vessel inlet no::le and the core inlet, and between the inlet and outlet of the core.
- c. The overall pressure drop between the vessel inlet and cutlet 1 no::les, and the pressure drop between various points in the
, reactor vessel flow circuit.
l d. The internals vent valves for closing behavior and for the effect on core flov vith valves in the open position. j Tne reactor vessel, thermal shield, flow baffle, core barrel, and plenum asse=- bly are made of clear plastic to allow use of visual flow study techniques. 4 All parts of the model except the core are sec=etrically similar to those in the prototype reactor. However, the simulated core was designed oc =aintain dynamic similarity between the model and prototype. t Each of the 177 si=ulated fuel assetblies contains a calibrated flow no::le. The test loop is capable of supplying cold water (80 F) to three inlet no::les i and hot water (180 F) to the fourth. Temperature vill be =easured in the inlet and outlet no::les of the reactor model and at the inlet and outlet of each of i the fuel asse=blies., Static pressure taps will be located at suitable points i along the flow path through the vessel. This instru=entation vill provide the data necessary to secomplish the objectives set forth for the tests. 3 3 2.2. Fuel Assebbly Heat Transfer and Fluid Flow Tests 1 . ,/ B&W is conducting.a continuous research and develop =ent progra for fuel asse - i'
.bly heat transfer and fluid flow applicable to the design of the reference re-t iactor. Single-channel tubular and annular test sections and =ultiple roi as-se=.blies have been tested at the B&W Research Center.
The reactor thermal design is based .upon burnout heat transfer experi=ents with (a) =ultiple ' rod,' heated assedblies' vith uniform heat flux, and (b) single rod,
- annular heaters with nonunifor= axial heat flux, at design conditions of pres-
- sure.and = ass velocity. These experiments are being extended to test non-unifor =ultiple rod heater assedblies as described in 1 5 2. The results of these . tests vill be applied to the final thermal design of the reactor and the
! specification of operating limits. , .3 3 2.2.1; Single-channel Heat Transfer Tests A large quantity of unifor= flux, single-channel, critiesl heat flux data has been obtained. . . References to uniform flux data are given in EAW-168(64)and 3.2 3 2 3 of this report. The effect on the critical heat flux caused by non- , uniform axial power generation in a tybular test section at 2,000 psi pressure .. was investigated as early as 1961.(61) This progra: vas extended to includ pressures (of)1,000,1,500, and 2,000 psi and mass velocities up to 2 5 x 10g lb/h-ft ,2 62 ."The effect en the critical heat flux caused by differences in
.. the radial andataxial investigated reactor'power distribution-in(an) design conditions. b3 annular test section was recently Data were obtained at press
~
of 1,000,1,500, 2,000, and 2,200 psi and at mass velocities up to.2 5 x logres "A lh[g.ftd, W 00768 3-101
,~ .. . - - . ...z --,.z.- . . .- . _ - . .
1 l The tubular tests included the following axial heat flux shapec where P/P is local to average power:
- a. Uniform Heat Flux (P/E) = 1.000 ccnstant
- b. Sine Heat Flux (P/E) x = 1 396 @ 50% L
- c. Inlet Peak Heat Flux (P/P) = 1930 2 25% L
- d. Outlet Peak Heat Flux (P/E) = 1 930 @ 75% L Tests of two additional, nonuniform, 72-in. heated length, tubular tests were undertaken to obtain data for peaking conditions more closely related to the reference de ign. The additional flux shapes being tested are:
- a. Inlet Peak Heat Flux (P/P) , = 1.65 @ 28I L
- b. Outlet Peak Heat Flux (P/E) = 1.65 @ 72% L These tests will cover approximately the same range of pressure, mass velocity and LT as the multiple rod fuel assembly tests.
. 3 3 2.2.2 Multiple P.od Puel Assembly Heat Transfer Tests Critical heat flux data have been obtained fro 6-ft long, 9-rod fuel asse -
blies in a 3 x 3 square array. A total of 513 data points was obtained cover-ing the following conditiens: 10 =< ATS
= 300 .
1,000 I P 5 2,h00 0.2 x 10 6 <
= G = 3.5 x 106 where ATg = inlet subcooling, F P = pressure, psia G = mass velocity, lb/h-ft The geccetry of this section censisted of nine rods of 0.L20-in. diameter en a 0 558-in. square pitch. Analysis of the last data of this set is in process.
3 3 2.2 3 Fuel Assembly Flow Distributien, Mixing, and Pressure Drop Tests Flow visualication and pressure drop data have been cbtained fro: a 10-times-full-scale (10X) model of a sir.gle rod in a square ficv channel. T;.ece data have been used to refine the spacer ferrule designs with respect to tixing turbulence and pressure drop. Additional pressure drop testing has 'reen cen-l_) ducted usir.g L-rod (5X), L-rod (lX),1-rod (1X) and 9-r:d (1X) niC_r .
,)OO 'O 3-102
\ Testing to determine the extent of interchannel mixing and ficv distributicn also has been conducted. Flev distribution in a square L-rod test asseibly has been reasured. A salt solution injection technique was used to determine the average ficv rates in the simulated reactor assembly ccrner cells, wall cells, and unit cells. Interchannel mixing data were obcained for the same asseibly. These data have been used to confir: the flev distributien and mixin6 relationships employed in the ecre thermal and hydraulic design. Flcv tests on a acekup of two adjacent fuel assemblies have teen conducted. Addi-ticnal tixing, flev distributien, and pressure drop data vill be Ottained to improve the core pcVer capability. The follcuing fuel assembly gecretries will be tested to provide additicnal data:
- a. A 9-rod (3 x 3 array) mixing test assembly, Of the same bundle gecretry as the EN3 bundle described previcusly, has been con-structed to determine flow pressure drop, flow distribution, and degree of mixing present during the EXE investigations.
Testing with the asseibly is in prcgress.
- b. Several 6L-red assemblies simulsting larger regions and various techanical arrangements within a 15 x 15 fuel assembly and between adjacent fuel asseiblies vill be flow tested. The hydraulic facility has been constructed.
3.2.; FUEL ASSEGLY, CC: TROL ROD ASSEGLY, A :D CC:CRCL RCD ERIVE !GCHAI;ICAL TFETS A :D II;SFECTIO!! To demonstrate the techanical adequacy and safety cf the fuel assembly, cen-trol red asseibly (CRA), and control red drive, a nubber of functional tests have tion. been performed, are in progress, or are in the final stages of prepara-3331 Prototype TeStin8 A full scale prototype fuel assembly, CRA, and control rod drive are presently being tested in the Centrol Ecd Erive Line (CREL) Facility located at the B&'a' Research Center, Alliance, Chio. This full-size loop is capable of simulating reactor environ ental conditiens of pressure, temperature, and coolant flow. To verify the techanics1 design, operating ec:patibility, and characteristics of the entire centrol rod drive fuel asseibly system, the drive vill be stroked and tripped approxicately 200 percent of the expected c;erating life require-rents. A portion of the testing vill be perferred with maximun nisalignment condi-tiens. Equipment is available to record and verify data such as fuel assembly pressure drop, vibration characteristics and hydraulic ferees, and to demon-strate control rod drive operation and verify scra: times. All prototype cct-ponents vill be examined periodically for signs of material fretting, vear, and vibration / fatigue to insure that the techanical design of the equipment teets reactor operating requirements. Preliminary test results are given in 3 2.L.3 5
, m-,
3-103
o 3332 Model Testing Many functional imprevements have been incorporated in the design of the prototype fuel assembly as a result of redel tests ru.- to date. Fcr exas;1e, the spacer grid to fuel rod contact area was fabricated to 10 times reacter sice and tested in a loop simulating coolant flow Feynolds nutters of interest. Thus, visually, the shape of the fuel rod suppcrt areas was c;tiniced with re-spect to minimicing the severity of flow vertices. Also, a 9-red (3 x 3) actual sice model was fabricated (using producticn fuel assent:3 sterials) and tested at 6h0 F, 2,200 psi, and 13 fps coolant flev. Frincipal objectives of this test were to evaluate fuel rod cladding to spacer grid centact wear, and/or fretting corrosion resulting free flow-induced vibration. A vide range of centact leads (including small clearances) was present in this speci-men. No significant wear or other flev-induced damage was cbserved after 210 days cf 1 cop operation. 3333 Cerr:nert and /cr Material Testing 33331 Fuel Ecd Cladding Extensive short time collapse testing was performed en Zircalcy L tube speci-tens as part of the B&'4 cverall creep-collap'se testing prograr. Initial test specimens were 0.h36-in. OD with vall thicknesses cf 0.020 in. , 0.02h in. , and 0.023 in. Ten E-in. long specitens of each thicknees were individually tested at 630 F at sicvly increasing pressure until ec11 apse cecurred. Cellapse pressures.for the 0.020-in. vall thickness specirens ranged frc 1,800 to 2,200
,/ psig, the 0.02L-in. specimens ranged fro 2,800 to 3 200 psig, and the 0.026-in. specimens ranged frc: h,500 to h,900 psis. The sterial yield strength of these specimens ranged frc= 65,000 to 72,000 psi at rc : temperature, and was 35,800 psi at 680 F. -
Additional Zircalcy h short time collapse specimens were prepared with a ta-terial yield stress of 78,000 psi at roc temperature and LS,500 psi at 61[ F. Fifteen specimens havin6 an OD of 0.L10 in, and an ID cf 0 365 in. (0.0225-in. nc=inal vall thickness ) were tested at 615 F at increasing pressure until cel-lapse occurred. Collapse pressures ranged fre h,h70 to L,960 peig. Creep-ecllapse testing was performed en the 0.h36-in. OD specirens. Twelve specimens of 0.02h-in. vall thickness and 30 specimens of 0.C25-in. vall thich-ness were tested in a single autoclave at 620 F and 2,050 psig. Turing this test, two 0.02h-in. vall thickness specimens collapsed during the first 30 days and two collapsed between 30 and 60 days. Nene of the 0.025-in. vall thickness specimens had collapsed after 60 days. Creep-ecllapse testing was then perferred en thirty 0.h10-in. OD by 0 365-in. II (0.C225-in. nominal vall) specirens for 60 days at 615 F and 2,1h0 psig. Kone of these specimens collapsed, and there were no significant increases in ovality after 60 days. Results of the 60-day, creep-collapse testing en the 0.h10-in. OD specimens shoved no indicatien of incipient collapse. The 60-day period fcr creep-collapse testing is used since it exceeds the point cf primary creep of the material, yet is sufficiently long to enter the stage when fuel rci pressure q tegins to build up during reactor operation, ie. past the ;; int cf taximun
differential pressure that the clad v:uld te sutlerted te in the ree:ter. -
a U,p <rn ed d.f.
,0~,
c-1
. - - = - _ . - _ _ - - _ - . _ _ _ - _ _ _ - . . -_. -
}'$ In order to help optimize the final clad thickness, additional clad-collapse testing is scheduled for 1969 using specimens fabricated to the reference design fuel clad dimensions, =aterial specificaticns, and operating conditions.
33332 Fuel Asse=bly Structural Components The structural characteristics of the fuel asse=bly which are pertinent to loadings resulting from normal operation, handling, earthquake, and accident conditions vill be investigated experi=entally in test facilities such as, i the CRDL Facility. Structural characteristics such as natural frequet;y { and da: pin 6 vill be determined at the relatively high amplitude of interest l I in our seismic and LOCA analysis. Natural frequencies and amplitudes re-sultin6 from flow induced vibration vill be teasured at various temperatures and flow velocities, up to reactor operating conditions. In the mechanical design of the spacer grids, particular attentien is given to the ferrule-to-fuel-rod contact points. Sufficient lead rust be applied te pcsition the fuel rods and to minimize fuel rod vibration, yet allow axial
, thernal differential expansion, and not produce fretting wear in the fuel rod ! cladding. Static lead and functional testing of the prototype grids will de=-
l onstrate their adequacy to perfor= within the design requirements. 33333 High Barnup Fuel Irradiations The primary purpose of the B&W High Burnup Irradiation Prcgra is to determine the swelling rate of UO2 as a function of burnup using fuel rods of the sa e
/ design as the core. In addition to determining the swelling rate, the effect . of several other variables includin6 the density, heat rate, and cladding re-i: straint vill be investigated.
t . The program consists of capsules so=e of which will operate at a heat rate cf i 18 kW/ft and otheis at a heat rate of 21 5 kW/ft. The pellets, other than U-235. content, vill conform to the reactor fuel specifications. The burnup vill range from 10,000 to 80,000 mwd /NTU with eight capsules exceeding h5,000 mwd /NTU. The capsules will not operate vith an external pressure. However, two different cladding thicknesses, 0.015 and 0.025 in., vill be used to vary the restraint offered by the cladding. The fuel rods vill operate with a clad-niing surface te=perature of 650 F. The diametral gaps between the pellets and cladding vill vary from h-5 to 7-8 =ils, to give s: eared densities of abcut 92 3 and 90.8 percent, respectively. These gaps and creared densities are consistent with the fuel rod specifications. The insertion date for the first capsule was September 5,1967 See Tables 3-17 and 3-18 herein. The tests are oriented toward the deter =ination of the behavier of materials Lin an irradiation environ =ent and'to deter =ine the optimum geometric and ca-2 terial properties for the specific application. The information is essential
+ for advance =ent of the art, but.is not considered critical in the sense that all of the programs must be cc=pleted to insure safe operation.
3 3 3.h Control Rod Drive Tests ar- 1hipection 3 3 3.h.1 Control Rod Drive Developmental Tests y b . The' prototype'. roller nut drive is.under test at the LEW Research Center, Alliance, Ohio. 00272 3-105
Wear characteristics of critical cc=ponents have indicated that esterial cct-
\ patibility and structural design of these ec=penents vill be adequate for the life cf the techanism.
The development program has been eczpleted, and the complete protctype cen-trol red drive is being subjected to environmental testing under simulated reacter conditions (e,xcept radiation) in the Contre: Ecd Drive Line (CRLL) Facility at Alliance. Environmental tests include: Operational Tests Operating speeds. Temperature profiles. Trip times for full and partially withdrawn control rod assemblies (CRA) fcr varicus flev-induced pressure drcps across the CRA. Life Tests (With internals assembled to maximum :isalignment permitted by drawing dimensions and tolerances.) Uc. of Partial Stroke Span of Centrol Rod Strcke Strcke Cycles Length, in. Fre: "?ull-In" Fcsiticn. in.
) 1,550 83 Frc: 56 Tc 139 5,400 50 T' 121 8,500 25 11h 139 8,500 13 126 139 .
No. of Trip Cycles 500 139 Frc: 0 To 139 Misalignment Tests 100 full strckes and 100 full stroke trips with internals tclerances altered to 1 5 tires maximum allcvable misalignment. Ccurling Tests Check of ecupling cperations after testing. The cycles above meet the total test requirements of 5,000 full strokes and 500 trips. The assembly vill be completely disassembled and i;.spected at varicus E&W facilities after completion of envircnnental tests. 3 3 3 1.2 Control Rod Drive Contr:1 Systen revelc; ental Tests
,g A ccrtr:1 red drive pcuer supply unit has been built in grcup pr:tecype fort.
(_/ Fell:ving the ec bined test of the pcuer supply and meinanism. therral. life, and si ulated failure tests vill te ccniucted. The ri lite? :a__;re :est well te designed to verify tha safety analysis. 3_1c c- Q()'M'.{
-%g The control rod drive control system will be tested in conjunction with the centrol rod drive power supply to insure proper creration. Simulated failure testing vill also be perferred on the ec:bined syster to insure that pro-tective requirements are being ret.
The position indicator and limit switch subsystet has been built in prototype fort and life-tested rechanically under expected envirenr. ental conditions. Further testing, both'rechanical and electrical, vill be done under expected environmental conditions at the B&'4 Research Center. Characteristics to be determined vill include accuracy, repeatability, li earity, short ter s t abil-ity, and long ter stability. 3 3 3.h.3 Production Tests Production tests discussed in this section vill be perferred either en the drives installed, or on drives manufactured to the same specificaticns. The finished control rod drive vill be proof-tested as a ec plete syster, ie, reenanists, notor control, and system control working as a systet. This prcef-testing vill be above and beyond any develeprental testing perferred in the product development stages. Mechanist production tests will include:
- a. Anbient Tests Coupling tests.
Operating speeds. Position indication. . Trip tests.
- b. Operational Tests Operating speeds.
Position indication. Partial and full stroke cycles. Partial and full stroke trip cycles. Control syste: productic;. tests vill be perforted as described in the fellev-ing paragraphs. The finished hardware vill be systematically cperated throuSh all of its cper-ating modes, checked over the full range of all set points, and checked fer proper operation of all patch plugs. This will check completeness and pre;er functioning of viring and cc penents. The crerating codes to be checked vill include such things as autcratic opera-Ne tien, manual 6roup operaticn, tric cr single CRA cperaticn, pcsition indica ice Of all CRA travel lini cn all CRA, trip circuit cperati:ns, I:, cu rand. CUT i ect snd, etc. 3-107 0027a '
l
- l l
1 i The trip circuit or circuits vill be tested by repeated operation. The over-all trip time vill be teasured. The accuracy and repeatability of the position indicaticn and litit switch sys-tems vill be tested. Power supply tests will be perforced to deter:f.e the upper and IcVer operating voltage and to prove inzunity to switching transients. Fault conditions vill be s1=ulated to prove that no unsafe acticn results frc: defective co=ponents, circuits, or viring. Ability to detect unsafe fault conditions at the operating console vill be determined. Typical of faults to be simulated are:
- a. refective limit switch or circuit.
- b. I preper CRA grcup patch.
- c. :efective patch plugs.
- d. Eefective group sequencer.
- e. refective clock.
- f. Defective autcrstic control signal.
- g. Defective cortand line.
- h. Iefective fuses.
. 1. Eefective single CRA centrol circuit or switch.
- j. Eefective power supply,
- k. refective noter translator.
- 1. Lefective : ter cable.
- c. Defective positien trans=itter.
The finished hardware vill be visually inspected for quality of workranship. This inspection vill include an examination cf the enclosure, cable entrances, dust-tightness, maintenance features, drawers and cable retracters, fasteners, stiffeners, codule counts, wire harnesses, and other similar details. 3 3.L II;TEFl'ALS TESTS AND INSPECTION The internals upper and lever plenu hydraulic design vill be evaluated and guided by the results frc: the 1/6-scale codel flow test which is described in detail in 3 3 2.1. These test results vill indicate areas of gross flov tal-distributien and allow verification of vessel flow-pressure drop cc putaticns. In addition, the test results will provide reasured pressure pulses at specific locations to aid in assessing the vibration response characteristics of the in-ternals cc penents. The effects cf internals tisalignment vill be evaluated on the basis of the test results from the CRDL tests described in 3 3 3.L. These test results, when correlated with the internals guide tube final design, will insure that the CRA vill have the capability for a reactor trip or fast insertien under all codes Of reactor operatien in the reacter ecclant env'----= - 7 ese tests vill not include the effects of neutron flux exposure.
'a'ith regard to the internals surveillance specimen helder tubes, the :sterial irradiatien surveillance prcgra: is described in L.L.3
( , m .n 190 y/ti 3-105
i i l l
'g All internals cc ponents can be removed frc the reactor vessel tc a11cv inspection of all vessel interier surfaces (see L.L.1). Inte rn11s cct-ponents surfaces can be inspected when the internals are rencved tc the canal storage location.
The internals vent valves are designed to relieve the pressure generated by stes ing in the core.folleving the LCCA so that the core vill rerain suffi-ciently cceled. The valves are designed to withstand the fcrees resulting frc rupture of either a reactor coclant inlet er cutlet pipe.
- a. A full-cice prototype valve asse:bly (valve disc retaining tech-anist and valve bcdy) has been hydrostatically tested to the taximum pressure expected to result during the blevdevn.
- b. Sufficient tests have been conducted at cere pressure to deter:ine
+
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valve ecver, and the ccver retcund resul*ing frcr irract cf the cover on the seat so that the valve rest.ense to e.velic b1cvdcvn ferees say be determined analytically. A prototype valve assembly has been pressuriced to determine the c. pressure differential required tb cause the valve te begin to cpen. A determinatien of the pressure differential required to cpen the valve to its maximum cren pcsition vill be simulated by techanical means. I d. A prototype valve assembly has been installed and rencved renntely in a test stand to confir: the adequacy of the vent valve handling tool. An analysis indicates that the vent valves vill not cren during Operation as a result of vibration caused by transtission of ecre suppcrt shield vibratiens. Tc verify this analysis 3&W has perferred vibration tests en a' full scale p;ctetype vent valve. The prototype valve was ecunted in a test fixture which duplicated the method of valve counting in the core support shield.
.e test fixture with valve installed was attached to a vibration test c._ chine and excited sinusoidally through a range of frequencies which enect-a..'.**' . r . . . s.v.
+w "h.ich ='.e_d #^-.. ... ** e ^.^-a. .e ".
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there was no relative ection of the valve to its seat for ceniitiens sicu-
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During refueling cutages after the reactor vessel head and the internals ple-nur assembly have been recoved, the vent valves vill be accessible for visual and techanical inspection. A hook teol vill be previded to engage with the previcusly centioned valve dise hcok or eye. With the aid of this tool, the valve dise can be manually exercised to evaluate the disc freecer. 2ne hinge design vill incorperate special features, as des critei in 3 2..1.2.h, to sinitize the possibility of valve dise notion impair:ent curing its service
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e u . 14. k. .e 'Q e ". e ..'. '.k. a*. a .. v ..' , e y e. *. e ". . ~ uld . ~ "a 4 '. d"a.*..s- '
. ~v. , +
'a.*.'.".e v~- '.' ka .= .-"...e~. '. .u. . .. '.ee 2d..- ..e e i g *. 1'. i c a ..*. 4. ..'.. 4 c a+ 1. c . o' .eu-b.. a -. - ..
s retien as a result of altered rctational clearances.
- 2. .L 2. .r--...:_: m-.a M
(1) 5. ".
- n ._ , G . r . , '."CPT C - n* T. o .-+. . a ". r w- g . c. .'o-
. . . . C='~ 'a**.e "..*..sm..*
-. y c '.
T.T.'C 4
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F = . +.. ' e 1 e s i . Cy'.'. ..#. e . . .e , - ~ .
%. . .- .. -Wa.+. e . .". . '. e .. 'de, n' ". e .* .v , n . "r . , P. . e ". .- a d.. i c t '. c. . c '. ".. eu*. .-~. . n*. . e " c . *. .- . . '. ..
(2) Nar7_^' m
-- 25, 2
n* ~r.41 1^M .
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. , W A ' ;*.'..P../. 0"t 2 .
.w. .
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**.r.., 5 a e..4 V. n (L\J *4..*,
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.... , v. . . J . a..2 C a. '. '. aS -'. . .a .. , J. C., .A J
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n (r.. w.., r. r - 1.'. - c
- t .
s mn.-5, n. C.~..__...c.<,.., ..e..... ~_ _ .<. r. . (6) . ,e w. .e , O . o. a.2 .e,.-
. .. . . . . . . .. .ss s, M,. C.,.~.w -- . w.
fusion Equatien Progra: for the Philec-2000 Corputer, WAE- 3-2hl.
- r. , F. . C . , 'nm. . .e.-u.c , A m..e. alim. i .. -' c..e. .y . e.u.....,", .- e .r.. =..a.ec. -
(7) u..c . .a.. 2 2 _ ..
.-. ,..,cee.
an . . . :
. .a. . :r. 0. 6.
^
( 4.,) *
~~
~ a . e e .~ , =. . =. . , a ". .'. =. ." .e ' " e . , C . , . . ~. .~. . .= ^ s =. < . . ' =_ 's - l. _. _
C .=. ' ". * ' '_. , W .w . e-, "..a*...., T. 3 . _.., r.e.
. . .e - v .u*. . c .. c '. *...
- e ."aw mL . a. . .. e 4. . . .o.'. .."."e4~..-
v. h N
- , ' , . 4. e. . . 7. . v%.1 e .. W.*." N.".*/. . L U . -
(w' av. . AS r ("% *f9
.O_ ..' 1 o.
'9g (26) Owen, D. E., Factors for One-Sided Tolerance Limits and for Variable Sampling Plans, SCR.607, March 1963 (27) Ecwring, R. W. , Fnysical Mcdel, Eased on Eubble Mtachnent, and Calcula-tion of Stea: Voidae. in the Subcooled Regicn of a Heated Channel, HFR-10, OECD Halden Reactor Project, December 1962.
(28) Zuber, N. and Findlay, J. A. , Average Volumetric Ccncentrations in Twc Phase Flow Systems, Presented at the ASME Winter Meeting, 19t~. To be published in the ASME Transactions. (29) Maurer, G. W. , A Method of Predicting Steady-State Boiling Vapor Frac-tions in Reactor Coolant Channels, Eettis Technical Review, WAPL-HT-lo. (30) Baker, O., Simultaneous Flow of 011 and Gas, Oil and Gas Jcurnal. Vel 5?. pp 185-195, 195L. (31) Rose, S. C., Jr., and Griffith. P., Flow Prc;erties of Eubbly Mixtures, ASME Paper No. 65-HT-38, 1965 (32) Haberstroh, R. D. and Griffith, P. , The Transiticn Frc: the Angular to the Slug Flow Regime in Two-Phase Flow, MIT TR 5003-28, repartment cf Mechanical Engineering, MIT, June 196h. (33) Bergies, A. E. and Suo, M., Investigation of Sciling Water Ficv Fegites at High Pressure, NYO-33Oh-8, February 1, 1966. (3L) Notley, N. J . F. , The Ther:11 Conductivity of Columnar Grains in Irradi-ated UO2 Fuel Elements, AECL-1822, July 1963 ,
. (35) Lyons, M. F., et al., UO2 Fuel Rod Operation With Gross Central Melting, GEAP-L26h, October 1963 (36) Notle'y, M. J. F., et al., Zirealcy-Sheathed UO Fuel Elements Irradiated at Values of Integral kd9 Ectueen 30 and 83 v c=, AEC1-1676, Eecenter 1962.
(37) Bain, A. S., Melting of UO3 During Irradiations of Short Duration, AECL-2289, August 1965 (38) Notley, M. J. F. , et al. , The Longitudinal and Diametral Expansicns cf UO,c Fuel Eletents, AECL-21h3, Noyetter 1961 (39) Lyons, M. F., et al., UC2 Fellet Thermal Conductivity Fre: Irradiations With Central Melting, GEAP h62L, July 1966 (LO) McGrath, R. G. , Carolinas-Virginia Kuclear Power Associat. s , Inc. , Re-search and Develeptent Pr gram, Quarterly Frogress Fepcrt fer the Perici April - M.ay - June 1965, CVNA-2L6. (L1) Ross, A. M. and Stoute, R. L., Heat Transfer Coefficients Eetween UO, and Zirealoy-2, AECL-1552, June 1962. '
?-111 l
'N (9) Mar 1 cue, O. J. , nuclear Reactor Depletion Programs for the Philco-2OOO Cc.puter, WAFD-TM-221.
(10) Lathrop, K. P., DIF-IV, A FORTRAN-IV Program for Sc1ving the Multigroup Transport Equation With Anisotropic Scattering, LA-337?. (11) Joanou, G. D. and D;dek, J. S. , GAM-1: A Ccnsistent P1 Multigrcup Code for the Calculatien cf Fast Keutron Spectra and '!ultigrcup Ccnstants, GA-1850. (12) Baldwin, M. N., Physics Verification Experinents, CORE I, c28 and Initial Conversion Eatio Measurements, EAW-S' L5h. (13) Clark, R. H. and Pitts, T. G. , Physics Verification Experiments, Core I, EAW-TM L55 (1L) Clark, R. E. and Picts, T. G., Physics Verifiestien Experiments, Ceres II and III, EAW-D' L56. (15) Spinks, H. , "The Extrapolation Distance at the Surface cf a Grey Cylin-drical Control Ecd," Nuclear Science and Engineering 22, pp 87-93,1965 (16) Neuhold, R..J., Xenen Oscillation, EAW-305, 1966. (17) Randall, D. and St. Jchn, D. S. , Xenon Spatial Oscillations , Nucleonics , March 1956. (18) Randall, D. and St. John, D. S., Xenen Spatial Oscillations, Euclear Science and Engineering. Ih, No. 2, October 1962. . (19) Control of Xenon Instabilities in Large FWR's, WCAP-?630 L, July 1967 (20) -Poncelet, C. G. and Christie, A. M. , The Effect of a Finite Time Step Length en Calculated Xenon Stability Characteristics in Larce FWR's, ANS Winter Meeting, Nove=ber 1967 (21) Clark, R. H., Eatch, M. L., and Pitts, T. G., Lumped Burnable ?cison Progra: - Final Repert, EAW-3L4?-1. , (22) Tong, L. S. , ENE Prediction for an Axially Sc nunifor: Heat Flux Distri-bution, WCAF-556L, Septenber 1965 (23) Tcng, L. S. , An Ezaluation of the Departure Frc: Uueleate Boiling in Bundles of Reactor ?. el Rods, Nuclear Science and Engineering. 33, pp 7-15, 1968. (2h) U.S.-Eurato: Joint R&D Program, Burncut Flev Inside Ecund Tubes With Honunifer: Heat Fluxes, The Eabcock & Wilcox Cc pany, EAW-3233-9, May 1966. (25) Jens, W. H. and Icttes, P. A., Analysis of Heat Transfer Burncut, Pres-
. sure Drcp, and rensity Leta fer High-Fressure Wnter, ANL L627. ':ay 1951.
s_/ es w . 8 6 3-112
(L2) Hoffman, J. F. and Ccplin, D. E. , The Release of Fis s icn Gaser Fre:
-h U .-a . . iL.. "d.ex43. m e e 1.1 e *. .ru _7
. . a 0-a_, . m'+.e' a+ "*t"- ". e . , ^ . ". " - a .e , Gr!" L"C..
.s -- ,#
Septerter 196L. (b-) .e,a. ,. a ..d .e , C . "o . '..' .e_3 -a.*...,* .r . u. . . , : e _c .. 4. "". = '..4' . . T..'_c_'^. _ . Gec : n.- '_ a =_ _c a. From Uranium Dicxide, GEAF L31L, July 1963 (LL) Rcbertsen, J. A.' L. , et al. , Iehavior cf Uraniu Cicxide as a Eeae:cr
- n. a el , e.m_:-t. 4.0: , . c. ,c e .
(L5) Parker, G. W., et al., Fission Freduct Felease Fre: UC2 by High Ten; era-ture Diffusion and Melting in Helium and Air, CF-60-12-1L, OENE, February 1961. (L6) runcc be, E., Effects cf Fuel Cracking, Veid Migratien, and Clad Cellapse in Oxide Fuel Ecds, Trans. ANS 11(1), p 132, June 196c. (1 -) t ca ..1 , n.. e., .n.e
. . . ._cegr.,,
4
- e. . . , . n. . . a- - .4 r. e-
. s a' c>ac-'
- w. .. e. a a " .: .------..r.an.
- -- ". . e e ' e .-
Studies en UCp Fuel Elements, AECL-2539, June 1966. i (LS) Ealfeur, M. G. , Fest-Irradiation
- ~
Examinaticn cf CVTE Fuel Assemblies, r sun: : 25v,-2,., . arc.n 19tc. (L9) raniel, R. C. , et al. , Effects of Eish Burnup en tirealcy-Clai, Bulk Uc,. Plate Fuel Ele:ent Samples, WAFD-26;, Septenler 1962. ( c,0) 7.-a c'+.". . e o .' C.v ' 4. .'..
- c a l. Puel . ...: .d Cla4'*._- 3 d".e *.^. 0. 1c_ _e .* -. ..
T . .c
..+.e ., '. 4 ' ' *. . .
--) . . . . -
...,.,.-,1, a n. u i 5 n.y. 4 2, . ,o..
(51) Duncan, R. N. , Eabbit Capsule Irradiatien of UC , 2CVNA-1L2, June 1962.
~
(52) Duncan, R. N. , CVTR Fuel Capsule Irradiaticns , CV';A-153, August 1962. (53) Frost, Eradbury, and Griffiths (AEEE Ear ell), Irradiatien Effects ir Fissile Cxides and Carbides at Lev and High Burnut Levels, Freceedings of IAEA S'rpesium en Radiatien ramage in Solids and Eeacter Materials, Venice, I: 31y, May 1962. (Sh) Gerhart, J . M. , The Fest-Irradiation Exacination of a Puo,-UO, Fast Eeacter Fuel. GhAF-3b33. (55) Physical and Mechanical Frc;crties of Eirealcy-2 and ~
-c, WCAF-3260 L1.
v..<3. . e . . (56) 32rgreen, D. , Byrnes , J. J . , and Senforade , D. M. , "Vibrat'.cn of Ecds Induced by Water in Parallel Flev," Trans. ASME EO, p 991, 1953. (57) Large C1csed-Cycle Water Eeactor R&D Pregram, Fregress Eeport for the Ferice January 1 to March 31, 1965, WCAF-3269-12. (55) Cla,rk, R. E. , Physics Yerificatien Experixents , Ce res IV and '.~. EAW-2'-
)7Up
. M (g .wy-4. . C_. p- 6 w . . e, 1hb.h.w 4
%/
00280 c,,' .1
(59) Clark, R. H., Physics Verification Experiment, Core VI, EAW-TM-179, December 1966. (60) Clark, R. H., Physics Verification Experiment, Axial Fcver Mapping on Core IV, EAW 'D!-255, Decenter 1966. (61) Svenson, H. W. , Carver, J. R. , and Kakarala, C. R. , The Influence of Axial Heat Flux Distribution on the Departure Frc Uaeleate Boiling in a Water Cooled Tube, ASME Paper 62-WA-2f [j (62) Barnout for Flow Inside Round Tubes With Uonunifor Heat Fluxes, EAW-3238-9, May 1966. (63) Nonunifor Heat Generation Experimental Program, BAW-3238-13, July 1966. (6h) Wilson, R. H. and Ferrell, J. K. , Correlation of Critical Heat Flux for Boiling Water in Forced Circulation at Elevated Pressures, The Labcock
& Wilcox Company, BAW-168, Novenber 1961.
g. 002S1 3-11L
4 6 4 , Arial Power Profile for 5 5 ?, i nsertion is shown_on g Figu re 3 -3 =I
.- 1
-t 1
I 1.7 - I
/
/ s r ;> y
/
/
x s ic_ 1.6 / g
. m a.
- . /
/
m =. >
.- o J- g , n. ' I5 e ao
- 2 m.
e - ..x. ,
,= ,
> 1.4 '
E $ .}
. / 1.3 /..
s 10 n. 20 30 40'. 50 6 0. ' 70 60 Rod Insertion. ",
.) I. } >i
'r i
(1-AXIAL PEAK T0 AVE RAGE P0nER
.?-
- i,e VERSUS XENON OVERRIDE R00 INSERTION
- b. p ' ' ,
k v. r .
,'t s 4 e j l
.. l t
l .; 1.8 16 / \ _l 1.4 / 1.2
- - 1.0 5
0'8 ' l" '\ \
-j j N
'i -{ f 0.6 \ ,
! O4 \
< g 0.2 T 0 ' I44~
10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 Distance from Bottom of Active Fuel, in. i AXIAL POKER PROFILE, XENON OVERRIDE ROOS 55 PER CENT INSERTE0 figure 3-3 00'2-S't
X X X X X X X X X X X X X X X X X X X X X X X X X X XXX X X X x X X X X X X
~
X X X XX X X X X X X X X X X . X X X X X X X X X X X X X X X X X X LOC A TI O N O F F UEL A SSE WB L IES CONT A IN ING B UR N AS L E POISDN RODS ,3 Figure 3-4 00785
4 4-
~
. N.
t
}
'l.
.110 4
t 100 T i N
\
\
\
\
- i 80 ;
1 i- ! t
- i
- 2 i t e t'" t
'a o- 50 t
, . = i c c \
~*
--~
.2 1 2. Si:.kik
- E, ' a Sz s
- a. t
\
- 40
- t .
\
+
~
5.l*:skik \' s
.\
4 ~~ _, 20 ( 1 0-0 I 2 4 5 E 7 8 Time. sec 3 PER CEt;I NEUTRON pawer
~
I VERius fire FOLL0nitlC TRIP. Fi:ure 3 5'
. v' '
4 - 000mN.$ 't* h 4 Y< k-+ ', ,e ,+ 4 e-v* e m~ r -wer e - -- - i,----~~----
. . _. .~ . . _ _ _ , . - - +
Q. 2.8
- 2.6 2.4 -
2.2 ~ 2.0 - 1.8 - - l.6 -- Loser Core
' ,,, , s /
f' ' N \ f
,. j /
/ / / \
P/P I4 -'s f
' ~~
Y #
\
\
\ # /
/
s j f f \ g f l / g,y , / ~' /
/ ~ ~ ~ ht -
N \
,## g % . ' fg '-
1
- m 1.0 - N
%/
f \\ --
/
f
/
\
\ /
f \ \
/ g 3 C m '
* .8 - N w/ N / \ / \
\
,g Upper Core y/ s / g i iO o
- .6- N !Al "n , @ - ; .4 -
r- e O" .2 - 5% 0=
"4 e m i i i i i
= 0 1 2 3 4
%5
.oc 5
Time (I), days 3O NOTES:
,, (( I. Power Ratio taken 36 in, from top and bottom of active fuel.
Case i n o m m No temperature steration, I fuel
- l.400 F.
m
;, ; Case 2 - Temperature iteration eith I fuel. 1.400 F.
C" Case 3 - Temperature iteration eith fuel 900 F.
'." m' o 2. Oscillation initiated at i . 2 days.
m =
2.4 2.2 -
~
tower Core
- 1. 8 - --
, . - . ~
1.6 - \
\
1.4 - pfp 1.2 - 1.0 - u
/
.8-h
.6 - ,/
.4 - N .,_
\-
.2 - Upper Core
.0 , , ,
0 1 2 3 4 Time (T). days Notes:
- 1. Power Ratio taken 36 in. from top and bottom of active fue1.
Case 1 - Temperature iteration with T fuel - 1.400 F. Case 2 . Temperature iteration with 900 F. fuel -
- 2. Oscillation initia ted at T - 300 days.
EFFECT OF FUEL TD/2ERATURE (DOPPLER) OM XENON OSCILIATICNS - LTAR DiD OF LIFE Figure 3-7 J 0 0 '537
2.6 - 2.4 - s 2 . 2 ._ Upper Core / \ 1 / g 2.0 - \
\
I 1.8 - ,~s I \ N I 1.6 -'g \ 2 s P/P 3,4 _
\ \
I
\ i \
sj ~
- 1. 2. -
\ I g
\ , I \
s i
\
/ '
\
1.0 - \ v \ / j
\
\
Y s ,
,g _ s -
\ ' }
\ ' \
.6 - I 4-
\
/ \
\ f
\
Lower Cere JN / \N
.2 - N e
%/ s,
.0 , 3 0 1 2 Time (T). days NOTES:
- 1. Case 1 - Divergent escsflation (without temperature iteration).
Case 2 . Power ratic variation with control (without t empe ra tu re iteration).
- 2. _ Oscillation initiated at T . 200 days, 4
CONTROL OF AX t AL C SC t LL ATIOh WITH PARTIAL 8.005 Figure 3-8
-; e 00288
J 9-3 erugif ^ "
)1
_8 1 0llAR 8hC SU$REV P.l QNI ,P- .OEICETORP sciTALUP08 O
)3-s ( citaR SND 42 22 02 8I 61 41 21 0l {'
3933S 0 10000 0
) snoi ti dnoC ngiseD t a RB ND muminiM ( _
- 17 1. RBND g -
9539 0 1000 0
)
f - - 3 03 I. RS ND g .., 999 0 1 00 O O I _ de t c et o r P noitalupoP %28 99
~
99 0 10 0
~
detcetorP noiti' -3 49 30.. 10
- ' 231,0 - no i ta iv eD d ra dnatS -
5 39 0 . e ulaV naeM _
~ _
908's stnioP a ta D f o r ebmuN ~
)p( )P-1(
00u96 . r
, 1.8 ji W i P/i' = 1.70 (Partial 1.6 V Rod insertion) 1.5
\ i i l l r e -T '
j P/P 1.50 1.4
/ ,
i s X (Modif led Cosine)
~
f i
/
I \ 1.3 / i ,
/ (
1.2 i \
! \
j
\ \
/ \
g 0. 9
, , s 0.8
/ ,
i
\ \
I \ Q
/
0.7 f i
\
/ , \
0.5 ,' , 3 's _, 0.4 i \ 0.3 ,. Fuel Midplane 0.2 Core T I Core Top \ 01 - Bottom i 3 m 144" g
=
)
0 ' 20 40 :l 60 80 100 120 140 10 30 50 70 90 110 130 Distance f r on Bo ttom of Active fuel, in. g, FORER SHAPE REFLECT'lHG INCREASED
'd AXIAL PCWER PEih FOR 144 ihCH CORE 00290 Figure 3-10
1.80 1.60 x s 1.40 1.20
.5 1. 00
% ~a E
,- 0.80 ,
O C
*I
~
0'60 2 O.40 \ 0.20 0.00 0 10 20 30 40 50 60 70 80 90 100 Percentage of Fuel Rods with Higher Peaking Factors Than Point Values, ',
~
I
- 3. _,\
OlSTRIEUIl010F FUEL R03 FE AKING 00?SJ. Figure 3-ii
110 Line 1 S5', F l os 100 m 5 m C O. 50 l 5 1.3 ONSR Limit o.
- " 60 E ) %
e 40 5
- 114% Overpower t
S , 5 20
/
B A I' 0 100 110 120 130 f ! Rated Power (2,452 Ot), s i i POSSIELE FUEL E00 DNC'S FOR ruit'JC DESIGN CD:,0IT IC::S - 35 E16 - E00 CGE l 00293 riEure 3-i2
i l:)
'l 50 E
o M
- 40 0
3 - E 30
=
i M ~
. 4.aximum 3- Overponer (1145) 2 20- '
6
. r- E l F
(
. ;10 '
D /
/
0 - I-100 .105 110 115 120 125 130 Rated Power 12.452 Pnt). .ir 4 1 FOSSIELE FUEL'R00 UNB'S F0F, MOST
.J
. FR00AEL E C;'0I T IONS - 3E E15-ROD CC;:E
''"" 3-'3 D O. ':)'). -
(1-P) (P)
*' n- O.1 ! . 0. 9
- f ,
i
! a f
l t i I I i l l I l l
- 0. 01 ,
0.99 i . i i
\ !
\ ! i
\ ! !
\ u I i i
\ l 0.001 . , 0.559 3 i
\ f
\ ! ! I ! '
\
- _i !
\ i b
\ \ r r ,
\ 1305 _)
\
0.0001 N s 0.9999 N \N
'A !
_ t 7 --%- _ i t E y ! mwi i 114'J / j NN I l k sl w 0.00001 . 0.55559 0 10 20 30 40 50 60 70 80 50 1 00 Percentage of Rods with a Loner Value of P,5 ,f- s_ DISTRIEUTION OF FDPUL ATION FROTECTED s P. AND 1-P VE . SUS WEER OF F0DS F0? KOSI PR05: ELE CONDII10NS Figure 3-14 0 0 " . 4- '
~
2.2 20 1 994 Confidence Basis Es b 1.8 - \
}.7l L, .
u 3 1.6 S a. g. e 1.4
~
b 3 [ 1.- 2 \
. .S
[ Oesign Overponer (1149) E o 1. 0 X
- 0. 6 l
100 110 120 130 140 150 Rated Paner (2.452 Mat).f,
% >. e ani S (r-33-is n01 u:a1 Cttt
'007.45 VEESUS REACTOR PCCER Figure 3-15
O l' 4 00 t
?O t
U02 Melts + g 3.00 7
'i
.e h ,
] Data Based On MOD GE A P-4624 2~ ( )
i U l k de - 93 */ cm ! s )
=
~
2 3 e
- 2. 00 5
S. p .} .l
- l
/
1 00 ( 0 1000 2000 3000 4000 5000 Temperature. F ooass THER"Al CON'OllCTIVITY OF 'J0 2 Figure 3-17
I
,20
+ SS f
+18 g7 5 )
l
+16 i f j "
90 l / ; a
+14 /- g3 g dg Design
/
/ g O ve r po.e r (1143,)
.12 /
c0
/ /
l / IS
+10 og 2"
[ E
= :'
.8 / /i '
103 5; lt /
, .\
d, f / Ic E p + E / /
/ 109 3e a / ,/ :
+4
/ < 116 2i
~
fl b
+ '2 f l g
127 $ j ' ' ' E O > l Oua li ty , 344
-2
/
/ Subcoolec 5 E
/. - 5
. 4 l ---
2.120 psig g 2.185 psig ; I 100 -110 120 130 14 0 150 Rated Power (2.452 Fat).- -
.- MXIMUF HDT CH %EL Elli Oc LITY VERS!JS E CTCR F0tER-Figure 3-16 009.am-se
- ~g 6000 002 Welting Temperature}
5000 / s l 4000 , I I f ' J I l E
\ l i 1 1 3 3000 I I .==
I I I i 100% Fonet 1 l E m i 2000 1 I I I i 114', Poner l 0' C0 i l I I I I I i 1 1 0 1 1 5 10 5 20 5 Linear Heat Rate,kw/ft s d FUEL CENTER TE'.',FEEATURE ;T ige EDT SFOI VERSUS LINEAR FCLER gg Figure 3-12 Y
N 70 - Gaussian Distribution 60 - y" 50 - _ f -
/0 -
30 . m . 20 _ 10 . ~ k 0 r f-q
, . . , , , , , , , N , -, , ,-,
0.6 0. 8 1.0 1.2 14 1.6 1.8
<'/6 EC J
NU!'5ER OF DATA POINTS VERSUS eg/6C' 0 0 ?/ 6 Figure 3-19
/
7 l 1 1.025 - 1.020 - 1.015 - F g" 1.010 - 1.005 -- g Fa .
, . n 1 000 60 70 80 90 100
- .= 0.S95 . .
0 990 - E . O.985 -- Fg (Interior Bundle Cell) 0.980 - 0.975 - I C.970 - Fattall Buncle Cell. 0.565 - 4 I O.SSO -
. Feculatica Pr ct ected .'
i liDT CHANNEL FACTORS VERS'JS. FER CENT FCPULATION FROTECTED Figure 3 20-i - v. a
-00'300 t . . . - .. . . _ . _
.-_2 _
4 d j 4 100 I i p 4 . Infin a te Sample - d 10% Conf idence a U f Finite Sample-9% Confidence E
-_ l l
% ] f in ite Sanple-i j' 70 #J II' Confidence E
, 60
-50 1-O 1.1 1,2 13 1.4 1.5 1.6 17
- Burnout Factor; DNB Ratio (9-3)
-BURNOUT FACTOR (n-3) VER$US POPULA110N FOR VARID'JS CONFIDENCE LEVELS
, Fogure 3-21 Ou,.01 1
4
/
W
.20 t
l ' l 18 l l 1
+16
,14 '
i l i l Het Channel nith 5'. ; I
+12 l
Fica Distribution Fatict
\ '
l k .10 ' I. l l l
! i i
.a t I
i ! ! -
! I
.E ,
., i ,
, l ht.ir.a! Channel O +4 t it t! cut F i ca -
3 o ' l ! D stritution Fatter i
.... +2 .
I I
! I
'- l Ouality
. l l l
-2 /[/ '
Suttocled
., !/ I Design Overpener l ! -6 ,
l I !
! 100 110 120 130 140 150 IED R a t e:: Poner (2.452 Wati l
DESIGN HDT CH AWEL AND N0ulNAL CHbNEL E XIT QUALITIES VERSUS FE ACTOR FDPER (alTHOUT EN;t NEERING EDT C-: wet f CTORS)
~
F iEure 3-22 ( l L' 00.'1C17 1
e
'\
- Bundle Burnout Test Conditions there Stable Operations Were Otse rved.
A worst Conditions,1140 Power
- m. mor st Cond itions,1304 Pon er e Nominal Conditions.114% Power
+ Ncminal Conditions.1305 Poner 3.0 25 .
A .. . + . a
'?= .
_[ 2.0
* 'N 7 . ..
} . , *. .
J .
.m 3 o 1.5 o
O g E ,
, g, ,_ , .. ..
Bubble To Annular s
't.0 (Baker)
,5 -
Bake 0 5 10 15 20 25~ 30 Quality (10 v apor total I b ) . '-
.4 % I FLC REG;L'E r F FGR THE NOT usti CELL Fiture '~-23 00;':03
l
+ Bundle Burncut Test Conditiens there Stable 0;eraticas Iere Otserved.
A torst Ccnditions,114; Pener E torst Conditions.130s Poner O Nominal Condi' ions.114% Poser
+ Nedinal,Ccnditions, 130% Power 30
+
+ ++ +
+
+
2.5 O (r +
+
+
+ 4 + +, * ++ ,
+
A I E
'"b +
; 2.0
+ t#+ + ++* ,
.+ ,,
- f. ' +
m-
) O 1.5 ., .t-
++ 4+++ t , #+ F + . , Bubble To
, {
" . Annular
+, (Saker) h t
=
i ' 1.0 I 4- +. +ew + ++ ++- ++ + + ,#.
+
+ . ,.- .
* +
+ r Bubble To *
'5 Slug (Saker) 5 10 ~ 15 20 25 - 30
-5 0
- Quality (10 vaper ' total 13 ), 4
%' FLDs REGIF'E tiiP FOR TEE- HOT CChiR3L RCO CEl i'
rigure 3-g ,,,c,
- Sundle Burnout Test Ccnditiens Where Stable Operations
-- '+
J sete Onserved. A nor st Condit ions.1140 Pcaer E ter st Conditions.130% Pceer O Ncminal Cond itions.1145 Poner
+ Nominal Conditions 130', Pceer 3.0 I '} .. . ..
l p 1 2.5 0 -
* =
ij A , . .t i g n4 - a . 1 '2.0 . Ms * .. o 2 g ,
, y_ . .
@V. o Sutule To
=
15 Annular s (Eaker)
~
i' O
= .
, g, ,
1.0
? .
I , i, .
,~~
Butule Tc -
$ .5 Slu;; < Eaker ;
10 15 20 25 3 2
-5 0 5-Quality (lo vapcr total 10 ) ,
, : P. O rtc< n c ws n e rc; w a r u t :stt
- l. bf[JrE 3-26 1
{ ? g } .3 i
- Bundle Burncut Test Conditions there Stable Operations
-) Rete Observed A tcrst Ccnditiens.1140 Poner E Rorst Conditions,130', Power
$ 9 Ncminal Condi tions,114', Poser
+ Nor,inal Condi tions,1300 Power
<1 9
+
2.5 . m . . !! 1::. A .t a . 'l - g 2.0
. 7 2 = =
i .e 1 * . . . . .
}x . , ,
e I E 15 Subble To
; Annular .
l i- 0 (Dake r) i E
. . g g , , *. * .
j 10 m
. . .~j c Bubble To _
/
4 -Stug (Saker) i M E-4 10 15 20 25 30
-5 .0 5 Quality (Ib vapor' total Ib). L
, 7
'i;; .
FL0f RE G il'E l'.i? F CF T HE H0i C0F.NEF, CELL FiEare 3-26 4 003M
~- y , k--- g, ,y. , -- , * -
1 l
. \
l 1 l
} 22 i
I ! ; !
. t i i
/ Design Overponer (1145) l I i l
\ \ I 1.8 \ \ -\ 1 I i i
^
n l b 1.50 Cosine l
.2 ' i 16 '
b \ N
= \ s s :
l i b 1.4 \ I
.e J 1.65 Cosine N
1.2 \ ls i \ \
\ '\
1~0 ' l l 1.80 Casini Kg ;
\ \
0.6 - ! ! 100 110 120 130 140 150 Rated Poner (2,452 MWt),4 HOT CHANNEL DNS P.ATIO (1-3) VERSUS P0rER FOR j VARIOUS AXIAL FLUX SHiPES r i tu e -27 00307
3 m't 150 _
., e
+. 140 ,
Design _Flcerate (13132 x 106 I t 'h r )
/
130
-/ 4, e
o. q ' 34 1 120 / h.
.c. . - s -) ~
o m
. 110 o
,e r j x5 - g DESIGN OVERFORER
= ; (114 x 2452 Met)
.e N. ,
.= .100 ..
7 s
.;g s
~
llli xw3 s l
\ , +
DNER (8-3) - 1.30
~ I-
__t
.t
. {' - .
.l.
l ..
.l ,
-2400 {E00 2600 .3000. 3200 3400- 3500
,. 1
,y Rea: tor Core Power. Mut' cf 'i, , s t '4 a *,
-_ ) % -
_a > FEICICP. CChiNT SISTEV FLCt YERSUS iC;iE M' 08; .riter: 3 2s
l 2.4 - t
. l I
22 , l WillNG LINE FLOR COEFFICIENT 1 110'e . 0 2 ., l 2 1005 02 , - 2.0 t i i
- I 3 9 0~- .02
! : 4 100', C E ..
l5 100'
. 01 .
2 16 i \1 - l 5 \ ; i i
= l \\
\\ ,
, m 15
= 3
\, >
N e e\ \ 4 5
- 1< - .~ i .
\~ \ l
= 1 N \j 1 30(*-3)
__ ._ s. _ _ _ _. _ _ _ _ . _ _4 i . 12 -
.I M \
2 ; 3 , l l - in I \ \\ \ ' 0.B i-1 00 110 120- 130 140 150 Rate:: Peaer (2,452 Pet) ,
- HOT CHANNEL ONS R ATIO (7-3) VERSt'S *:t ER t tTH
. ('} ' RE10iGR SYSTEu FtC* 'N3 ENEFGY u!!!'d.15 FAR W ET ERS. g ),c q F :i.::e 3-25
4.00 3 ~ g. g B&W Design Value _ l --- CVNA-246 1 -- DEAP-4524
\
~
~ \
i 5 w \\ 3.00 - L \
=
\ -
_ U02 Melts B '
- \
\
E _
\ CVNA 246 5 \
1 -
' /
$ \
/
% 2.00 \
5 \ /
~
E \ /* x / g B&W 0esign Value
\ j
", /
% -y GE A P-4524 1 -00 I I I l I i !
0 1000 2000 3000 4000 5000 Temperature, F~
.ji , THERFAL C0h0VOTivlTY OF S3.5-PER CENT DENSE SINTERED UO 2 PELLETS 00310 FiE" 3-30
l l 6000 5500
/
f '
/
U02 Me I t in g T er:p
' ] 7
' ' -L m 5000 --
__/ _.- - E ,
* /
! 5 / L/ 2 1004 Power % 7 /
' ^'
E^ 4500
?
e ff/ af/ b / A 0e s i gn Ove r pon e r E [j (1145) u 4000 j / . (( ' 3500 2
/
3000 6&W Oesign Value
-- GE A P-4 624
- - - - - CVNA-246
.2500-l 6 8 10 12 14 16 16 20 22 24 26 '28 30'
- Linear Heat Rate,kW/f t s-
-\_.]-
FUEL' CENi!F. IEuFE MTL'E FC' EE 31).N thG-0F - L IF E ' CON 3 t T IONS
*
- 3 -"
00R!.1
1 l 60C0 5500
/
/
U02 Melting Temperature
' I'
^
[ 5000 ! m Design Overpc et z' j, g ( 114 ", ) / E / d 5 4500 - ll! h /
=
- h W 3
=
0 4000 l 1001 Poner 3500 - 3000 GE AP-4 624 B&W Oesign Value
- - - - - CVN A-2 4 0 2500 I I I i 1 !
6 8 10 12 14 16 18 20- 22 24 26 28 30 Linear Heat Rate,kW/f t 1
- ;' FUEL CENTER TEPPER!iURE FOR END-CF-tlFE 00N0 lit 0NS
gur e 3-32 00312
)
.o il 3,300 1 1:, 2.900 r 2,500 J
.. 3
- 1. ;) 3 l $ 2.100 1 :
i . I 1,700 i
- 1. 3 M 0 20 40 60 80 100
, Vol. Fraction of Total Fuel.L l-,
(at or above Fuel Temperature)
,L s- FUEL TEMPERATURE VERSUS TOTAL
'G FUEL VOLUt'E FRACT t CN FOR
' [' - EQUILIERIUM CYCLE AT END OF LIFE OO;t (3 Figure 3-33
s 1
\ _
\
5-a 3 __.---2 - _- 3
--2--- --3 1 3----
1
--i 07E0\. 0 930 0 E39 0 !?E 1. 043 0.EEE 0.995 1 005
\
' 1, 3 3 2 l3 1 1 i
05\ s 0 738 0 794 1 054 1 C05 1 311 0 970
\_
3 2 2 3 2 1 C.E$\ 1. 097 1.162 1 C76 1 062 0 781 N - . 2 2 2 1 s 1.202\ 1.148 1.091 1.034 s Ne 3 1 1 1 043's 1 216 0.798 N
- N um::e t Cycles Burne:
14 0 911 .
\ Assc.01) PP TVPICAL RE ACICR FUEL ASSEuRLY PO*EP GiSTRIEUTICN AT ENO CF LIFE E0'JILIS-Ilp CYCLE CON 0lil0NS f 0R 1 E CORE Figure 3-34 A
003i4
N i .I t Fuel I ' Clad L._._
.! 1400 3000
\ 10 As/ft 1
2600 w \ I - 2200 s . I
- 3dip 6 ku/ft E
c 1600 E
\
o t -
, ~
1
! 1400 l!
1000 580 F Tavg Coolant s 600 I ! ! I . 0 .04' .06 12 .tB .20 fuel Rcc Radius in.
\-
FUEL E00 IEtFEEaiURE FEGF!LES 11 e is: 1: an Figutt 3-35
!* A
. -g m
h ' 8 3 g '% 4- , -
*E85 0"*4 ~$ "
m
. . .7 h
' ~
a
+ and?~ R
\ --
. < 3040 R o*+< -
8 .
- w. l. N 3 1 <
. < 8 %
i
=
n t
+ -
a @
~
r t
<l J Qvi 1 4 8 3 g
) -
\@ < r g i E
a -
+
s n a N $ nN 3 .
. I l
3 o O 00 N - .- 4 sn n
, l o 5 ?
,p -
.. o 8 8 8 8 8 S e 8 g d d ; d o d g paseagay seg uoissig A
J PES CEmi flS$10m Gas EELE ASEO 15 i funtilta 05 iME ArEt1GE TEarEti:UEE OF THE UD T '.' E L 2 003;E ri pre 3-3t
J l.8 l 3
.l l
' N P/P - 1.70 (Partial Rod
, / \ i insertien)
/ '
I l5
/
r
\ f -T x g/p - i,5o l4 / l N J (Maci fiec Cosine)
/ / \
g 100 Days EU EU i , g l.2 j f
/ '
300 Days U/E} i . . 1.1
#f ,
I g 930 Days-
% l s xw / " EWRJ -
bl / I \ l
. o. m 0.9 V# / l 1 \ \T i ll 7 ,
n
" : . 8 s\
\ \\
" N/ // i x s\ \\
" /E l N x\ \\
0.5 'M < / Ifg l N% \13
- 0. ti
// I i
\ y\{\
\
r/ i N \j 0.3 i Y"- F Fuel Midplane 0.2 Core i \ Eottom I Core S h-Top \
. i I 6 l j 0.0 I ' ' ' ' ' I ' I I I 20 % 60 80 100 120 1%
10 30 50 70 90 110 130 Distance from Bottom of Active Fuel. in. ArIAL LCCAL-TC-AVE: 15E ELauf 1RD INST'ktAkECUS F0=ER C04PartsCns Oca(7 figu tt 3-37
50 Design limit 40 t 30 E E 1 50 BU And ' .50 g Axial Shape c.s
=
2 / b 1.50 BU And 1.70 E [ Axial Shape 10 N _ 930 0ay BU And 1.~70 Axial Shape 1 0 ' ' ' ' ' O 1 2 3 4 5 6 7 8 9 10 i nitial Cold-Diametral Clearance, in. x 10-3 FISSION GAS RELEASE FOR 1 50 AND 1 70 MAX AVG AXIAL-POWER SHtPES 0031.s- Figure 3-3B h
3M0 ' i
' i I
DesiEU l "i\ 3000 Closed Pares 2500 l c. I 2000 a '
/
5 0 1 50 Axial foner j 1500 - And 1.50 Burnup Shape
. /
J &, u 1000 1 A / w Open Pores 500 1.70 Arial Power And 930 Day Burnup Shape 0 ' ' ' ! ' ' 4 6 8 10 0 2 Initial Cold Diametral Clearance, in. x 10-3 (s~S.> GAS FRESSURE INSIDE THE FUEL CLAD FOR ViR: US AXI AL EURNUP AND FO*ER $ dices FOR IDE AL THER!>Al EXPANS10N F.00EL 00ft J.9 Figure. _-a
2000 [ 1800 -
/
/
"r f
- 1600 ,A 2 /
$ 1400 / !
f 5 1200 M
~
"O Contact" Cracking Effe /
E 1000 l' o 7
/
k 800 /
- / /
'i 600 /
7 2.5-Mil Cracking Effect N/
/
/
5 400
/ l deal Thermal 200 > / -/ Expansion
-/
0 2 4 6 8 10 12 14 16 18 20 22 linear Heat Rate, kW/ft 1 SENSITIVITY ANALYSIS OF Tile EFF ECTS OF FUEL t[!; CRACKING ON FUEL-TO-CLAD GA? CCHDUCT ANCE Figure 3-40 00R20
3500 , , , g i 3 Design Limit 3000 2500 k ideal Thermal p Expansion
= 2000 /
7 E
, 2. 5-M i l C r ac king E f f ec t 5
1500 # 7 7 1000
- /
"O Contact" Crackir,g Effect 500 0 1 i i i i 0 2 4 6 B 10 Initial Cold Diametral Clearance, in. I 10-3 a
\./ SEl:SITIVITY 1.!;ALYSIS OF THE EFFECTS OF FUEL CRACKING ON INTER;;AL PRESSURE OOP,'d'1' F iture 3-41
~
3500 i i Desiin Limit i 3000 Ideal Thermal-
\
l Expansion 1 I
.; 100% :l 114k Power Power
; [ l l E; 2.5-Mii Cracking m l
= 2000 Effect 5 I e
.e_ .
/lJ l 5 1500 l
i 3 l
'"O Contact" 5 I Effect j 2 \
C racki ng/ 1000
' /
j l I 500 l I I , 1 0 {- 2000- 2200 2400 2600 2600 3000 Reactor Power, NW
$ '. SENSITIVITY ANALYSIS OF THE EFFECTS 0F REACTOR F0 ER CN ll;TER'?AL PRESSURE
, FiEure 3-42
~
d 9 I 1 043 1.050 1.058 1.060 1. 046 1.025 1.0 57
- 1. 0 0 0 0 0 0 S 950 1 053 1. 44 1.046 1.052 1.034 1 015 1 00 0S7]
025 (10 1.05E 1 045 .042 0 1 045 0 1 1.010 0 1 00 S 971 l('h'j/ l ((h} l t('h'/ ' l ((" i (h l' 1.058 1. 051 1. 045 i 1 034 (h"/ 1 C(h/ I \' '3 i t"' 1.015 1.017 0 999 0 557
" "\/ i f f a f\
2 1 047 C f ') i t 3 i C "/ i (0aaj i C(/ i t['"/ i t'71 1.034 1.025 1 02 1 014 1 013 0.552 0 559 0 0 0 - S 09> i v N y ,011v i v i o i v 577( 9 >77 S 1.030 1. 015 1. 1. 01 6 1 014 1 002 0 0 543
- 0. 9>1.008 vivid 0 S
- 1. 004 0
1.002 i 1 001 v i v i v i t /77 I S 0.SS3 S 0.576 S 0 951 9 ~ 0 53 I
- 3. 9 0S .5 y vi 9 SS iev i v S S SE7 o
.0.978 T 0 975 0 S73 1 0 57'_ 0 SE1 I 0
'.S Id'v 0. ,S 2 i 0 ESs b Nucteaf Pe an in; f act c :
@ HOT UNIT DELL
@ HOT t1LL CEtt NOWINAL FUE L R00 P0n ER PE AKS AND HOT C0hER Mu CELL EXIT ENTHALFY RISE R ATIOS
@ HOT CONTROL R00 CELL F i gu re 3-43 J
00 R3
~
( l 1 I d--- lNST .035 .1.01r 1,007 1.00 - D 9 9. -0 SSS -0.577 . - - 1 02 1.03 1 . 04 1 04 1.03 1 020 1 014 I5 03 0 v i v i v i (50> C 9
- 1. 0> 1. 031.027(40>1.025 1.035 1 01 i
v i ( 5 761 1 005 1.005 1.00 1 01 03 0 0 01 SSE 1.03 1.02 1.02E 1.03 1.0 1.002 1.007 1.000 f f f 49ie<\ e i e[\ l.040 i 1.034 1.031 i 1.025 e i e[ m 1.007 i 1 005 fei& 1 0C5 0 551 f f f f i"\'/ DH
' t '"/ i O i O i M, o/ i O< 1005O i t'k 1.031 1.018 1 011 1.00 1. 00 1 i
1 000 0 SS"
~
t>2i Ry1005v i o0 v >02v t>51 0 {0.5s 04 S7 5 S 1.021 1.006 . 1.011 1.007 1.0 0.9 53 0 SS4 h th (h (h (h (h Ch ( 1.017 1.014 1. 13 1.010 1 003 0 54 0.565 0 . 9 77 m <m rm rm <m rm rm rt 1 011 1.010 011 1 004 0 555 0.ji5 C.57E b hutieat Feaking Fatter h HOT UNIT CELL - Enthalpy R i se Factor C HOT eALL CELL WallWUW FUE L E03 POKER PEAF.S
@ HOT CORNER CELL A hD C E LL EIli ENT HALPY RISE R4TIOS , h H OT CONTROL R00 CELL F gur e 3-44 y fl 0034 i
1.6
} ,
,.5 \
G - 2. 55 x 106 lb/nt-t12 14 \ ! ' I I I l3 , 1-3 DNS Heat Flux - (Design Limit) 1.2 i ' ' i N 1.1 \ m 10 , 2 0.5 7 2 0.E
, Minimum DNER - 1. 97
. 0.7 b
O. 6 E l
% 0.5 o r N 0.4 _
O.3 , alculated Sur f ace 0.2 j Heat Flux 0.1 I O 540 560 580 600 620 640 E60 680 700 720 Local Enthalpy, Stu/lb
') CALCULATED AND DESIGN LIVIT LOC AL HE AT FLUX VEF. SUS ENTHALFY IN THE HDT UNIT CELL AT THE NC"!NAL C:'i0li!CN Figu re 3-45 003:25
y
.j' :}
1.6
+
l 4 1.5 y I G - 2.31 x 106 l b/h r- f t2 1.4
\
1.3 \ W-3 DNB Heat Flux I2 , (Design Limit)
! =
g 1.0 0.9 \ { j Minimum DNBR 1.71 b 0.8
,_ ) 2' } 0.7-t- -
f 4-g 0. 6 U 0.5 N
~
[
' 0. 4 0.3 Calculated Surface 0.2
/ Heat Flux v ', \-
0.1 0 540 560 580 600 620 640 660 660 700 720
' Local Enthalpy. Stu/lb
-4
'~ #
CALCULATED ANO DESIGN LiiilT LO !L EEAT Flux YS ENTHALPY'IN THE HDI thlT CELL Ai THE DESIGN C0X0li 1(N 0 0:I M ig.,,, 3_,3
24 O 2.2 i Curve Leakage (>, af 131.3 x 106 It/hr) I O 2 2 (Design) 2.0 \\ 3 5 4 10 5 6. 6 (Design Plus One Vent Yalve Disc Removed) n
, 1. 8 \ ; \
a m E 1.6 '
- \
. i !
5 1.'4
\
_ _ , 1.30 (W-3 ) 1.2 N NN ' \ \
\ \
1.0 cA XM
- 0. 8 /
100 110 120 130 140 150 Rated Power (2.452 Mtt). 5 v i D!!5 RAT 10 (T-3) YERS!!S FC;ER FC:. Yif.lCUS IRLET-iC-0UTLET CCRE 6ff ASS' LE AKiGE 00:U;7 " P ' ' 3 -4 7
*h. .3.
Co. C. # v ,C.3t O g
. 4 e_
7 _ _ t._9 s j.
' lf jp.mg=.
= =
~Mi
. .p r,d i ,
y A l , s . g A . . . . . . .. . . . . . . . . . . . s y Aa ~, 4 a, f- ,; a, ai g a s e ' s g% 3 %._._ o a i v g'p fp x i , i . me N IJst Q/ l 19h A
/r x -
( i i 4,n/c
.. u ..m .
Jk ;Ml ,g, _7 - a " i l p^ ;y y _ b c ~~
)s f c
s + w7u w i. a az9, s P
< . i
- r. s e s N E> \ =
/q.-' i s @ ; N !
/y' N -
%s%1 \. \
s c-s s
.s e g& ,
- s,,
b x s ......,,,, s q N: // . ,
.q s
.....a s h N- / ,g s, s r x u
,/K s j- 3)s 4,
s g _ , . . - . . . . . . . s - x s fj .' _
\; / iI s {<j }N, l
......... s 3.- / = \ ,. ;-
x b, / \! 3
'N sv r.u.s A 'b.5h1 N 4.A ..
NN k bd++eww&J S HH; yHHH gr 'fl0){l)d[j]0 % yh . . _ , . . . . _ . , u
\ M Nj x I N j
)i
.- \
t
"~
N[m I1i[ ,
, .I e i! i e
_s
- i. REACTOR VESSEL AND INTER!MLS - l 4 GENERAL ARRANGEMENT 003V8 Figure 3-48
FUEL ASSEMBLY
,/
t l/l / N. T
. [M ,M ,[
SURVEILLANCE SPECIMEN
.L ._ 9_
I I E_ r l') k 13 } = IGr 4l b,O\'f i HOLDER TUBE [
~~
, 13 = OlT(0) . lel___Lo \ , ? j/ l WI O lSl 'r@ *13 l 1 :
\\ CONTROL ROD O] Jei le .l4l l lef. !ej iM / ASSEMBLY (49) y c:_xl3HOx e! tel 2 4r[o
/ !
31_ l a . Il l ! l GI 7 3
-4 != C , !?! je @-t p e- - ] [ %'s )L-cs)- t I
- 1:.1 l 8i ./
INCORE INSTRUl/ENT h:
- is al3 ist el le @ .
13 _: LOCATION (46)
/ ol !Gl G =
ll @ G l* 3 1 8 e
- l @l @
iol l REACTOR VESSEL e l 1
. l fl. ' 0(. l[l.j O
( THERMAL SHIELD
//////l CORE BARREL XENON CONTROL ROD ASSEL!ELY (C) y O REACTOR VESSEL 'ND INTERNALS-OU"OhhSECil0S Figure 3-49
5 4 . I i 6' L__J , C3 ] g i /,.J y , 1 s I t IOt
,/ /
, .; /
/
l /;
} .
5
.. h
' /
, ).1 '{ .
/ d 9 1 LJ\ .
; ) ) )x M -
\
+ s "I CORE FLOODING
'i E
- NOZZLE
}. .,
'j b
h '
' 7' ;I
; NT Y.
l lg,I t W+ .j N
*w
.U C0FtE F L OOD IN G APF. WS EPE N T-
-003.30 ritne 3-se
~
r 9 \' a\ \ / N N l'/ / i
'p /,/I' (N s
\
\\s
,/ .
c 4'L.N A sN i......... hl, xM6' 3'~
$n l @<'
..M, x\
\\
l ,// //' 's\ V, ','h / d \ W.: i ';qj a d .g N
,, b Nx
,s h# b)'V,j;!i. '"
,~
,, s
\\\
p ,/ /
,/ ,
\\
.f s x'N S S L.-
SECTION ZZ Z l I i d '
', s
/
/ a_
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