ML20054G254

Forwards Response to Questions in NRC 820226,0311 & 0409 Ltrs Re Prototype Test Program.Responses Will Be Incorporated in Amend 69 to PSAR

ML20054G254
Person / Time
Site:	Clinch River
Issue date:	06/17/1982
From:	Longenecker J ENERGY, DEPT. OF, CLINCH RIVER BREEDER REACTOR PLANT
To:	Check P Office of Nuclear Reactor Regulation
References
HQ:S:82:047, HQ:S:82:47, NUDOCS 8206210319
Download: ML20054G254 (68)

Text

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u Department of Energy Washington, D.C. 20545 Docket No. 50-537 HQ:S:82:047 JUN 17 U32

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Mr. Paul S. Check, Director CRBR Program Office Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Conunission Washington, D.C. 20555

Dear Mr. Check:

RESPONSES TO REQUEST FOR ADDITIONAL INFORMATION

Reference:

Letter, P. S. Check to J. R. Longenecker, "CRBRP Request for Additional Information," dated February 26, March 11, and April 9,1982 This letter fonnally responds to your request for additional information contained in the referenced letter.

Enclosed are responses to Questions CS 210.12, CS 270.12, CS 220.34, and CS 421.04. These responses will also be incorporated into the PSAR Amendment 69, scheduled for submittal in July.

Sincerely, Jo3nR.Longen er Acting Director, Office of the Clinch River Breeder Reactor Plant Project Office of Nuclear Energy Enclosures cc: Service List Standard Distribution Licensing Distribution 8206210319 820617 PDR ADOCK 05000537 A PDR

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l Ouestion DCS 210.12 Describe the CRBR prototype scale model testing program and its correlation with the CRBR preoperational vibration testing program f or assessing CRBR Internals vibration motion. In addition, describe the acceptance criteria for Internals vibration tests.

Resnonse 4

Section 3.9.1.3 has been completely revised to incorporate the current CRBRP vibration test program f or reactor Internal s. The test program including i scale model testing is described in Section 3.9.1.3.1.4. The correlations developed f rom the scale models to predict CRBRP vibration are also given in Section 3.9.1.3.1.4 under the paragraph titled "Model to Prototype Scaling I Ratlos." Utilization of the scale model test data in the CRBRP vibration

.u= ing and considerations f or developing the acceptance criteria are discussed in Section 3.9.1.3.5.

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1 3.9.1.3 Dynamic System Analysis Methods for Reactor Internals 3.9.1.3.1 Summary of Overal1 Flow Induced Vibration Assessment Program for CRBRP A comprehensive vibration assessment progre to assure that excessive flow induced vibratory motion of reactor internals does not exist, is part of the CRBRP progre. The progre, schematically illustrated in Figure 3.9-4, assures the structural Integrity of reactor Internals. The progra accomplishes this objective by means of the following key actions:

1. Design to avoid flow induced vibration

2. Evaluate susceptibility of each component to flow induced vibration

3. Analyze to provide a model evaluation of each component

4. Test both scale model and f ull system simulations

5. Factor FFTF operational measurements and experience, as well as relation between FFTF scale model testing and in-reactor data, into CRBRP eval uations.

6. Monitor upper Internals structure vibrational response during pre-operational testing to supplement and confinn vlbration results from the model test program.

The progrm Integrates the input from these sources to assure that excessive flow Induced vibratory motion of reactor Internals does not occur.

The overall program has been formulated to meet the Intent of NRC Regulatory Guide 1.20. The program relles on highly Instrumented model tests to assure that flow induced vibration problems will not exist in the reactor. This approach is an extension of the vibration assessment progra developed for FFTF. It utilizes the exoerlence gained from FFTF model testing, where models i were used to assess the potential for flow Induced vlbration, with the l

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~~ -s n Pana . 2 {g,pg]pg3 objective of developing correlations which will predict CRBRP prototypic motions using scale model test data. Where scale model tests may not be adequate, such as f or long, slender members with flow thru clearances, full  ;

scale tests are perf ormed.

Details of the program's key features are provided in the following sections.

3.9.1.3.1.1 Design CRBRP components are designed to avoid excessive flow induced vibration.

Heavy structures are utilized, where possible, to withstand vibratory forces.

Generally, structures and components are designed to have natural frequencies as high as practical to avoid coincidence with known f orcing mechanisms.

Mechanical stresses that are caused by flow induced vibrations that are a'unidable must meet the f atigue requirements of the ASME Boller and Pressure Vessel, Section Ill, and Code Case N-47. When combined with the ef fects of other operating conditions, maximum displacements will be limited, as required, to assure proper f unctioning of the components.

3.9.1.3.1.2 Canponent Evaluation A general survey of all CRBRP reactor internals was made. In this survey the following considerations were given to each components o Type of possible component excitation, vortex shedding, turbulence, pump pulsation, jet Impingement, Jet reaction, gap modulation or other fluid elastic mechanisms.

o importance Index - based on the relative importance of the component and type of excitation.

o investigative priority - priority is based on degree of concern and whether the component can be evaluated by state-of-the-art methods.

Table 3.9-9 summarizes this survey and provides the methods selected to address the question of flow induced vibration.

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u wo Paga - 3 [8,22]#23 3.9.1.3.1.3 Analysis

, Existing analysis methods (ANSYS, etc) are used to compute the natural t

frequencies and mode shapes. Component response is calculated, assum,Ing a fluid forcing mechanism, and the resulting stress or component motion eval uated. Presently, one of the most of fective means of designing'against ,

the occurrence of severe flow induced vibration problems is to separate l structural natural frequencies f rom expected excitation f requencies; generally, natural frequencies of Interf acing components are separated.

By utilizing existing flow induced vibration analysis and Information, the f ree and forced response of the CRBRP structures are investigated. The prediction of flow Induced vibration response is limited to the state-of-the-art in describing the fluid to forcing function associated with fluid exe::ellcr. mechanisms and damping. idany excitation mechanisms could be responsible for CRBRP component vibration. Sane examples are cross-flow vortex shedding, parallel flow boundary layers, fluid borne noise, wakes f rom adjacent components and fluid elastic coupling between adjacent components.

Although all of these mechanisms are of concern, generally characterization is possible for only the first two mechanisms when each is assumed to act alone.

Even then the forcing f unctions are based on assumption. Therefore, experimental evaluation of components using scale models is relied upon. The results f rom the experimental portion of the program will be f actored into the analysis and final assessment of the component acceptability regarding flow induced vibration.

3.9.1.3.1.4 Testing For many regions of the reactor, the complex interaction of the flow field with the reactor structures precludes the utilization of analytical methods for evaluation of component flow induced vibration. For this reason a comprehensive series of system tests, both scale model and f ull scale, have been planned for CRBRP. These tests reproduce flow fields closely matching l

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Pega - 4 [8,223#23 those which will exist in the actual reactor. The planning of the experimental programs was carried out in consultation with recognized personnel in the fleid of flow induced vibration f rom the Argonne National Laboratory, Hanford Engineering Development Laboratory and the Westinghouse Research Laboratories. In addition, during performance of the progran, the test methods, similarity requirements, test results, etc. are continually reviewed by these personnel as well as by personnel from other organizations knowledgeable in flow induced vibration. Every ef fort is made to ensure that the results obtained from the model studies are correct and directly applicable to CRBRP.

The major models ur,ed in the experimental portion of the CRBRP comprehensive vibration assessment program consist of the following:

. '-!ct Plenum Feature Model

b. Bench Test Models

c. Selected Full Scale Models

d. Integral Reactor Feature Model (IRFM)

A description of these various models follows. Since the IRFM test program Is the most significant part of the total test program this test is discussed in detall in later sections. Results f rom the various model tests are provided i n Section 3.9.1.3.2.

a. Inlet Plenum Feature Model (IPFM)

The IPFM models all hydrodynamically wetted surf aces in a full 360 degree (0.248 scale) sector of the inlet plenum and lower Internal components.

l This test, initiated in 1974, was completed in 1976. Seven of the 61 i lower inlet modules (LIM) were dynamically simulated with accelerometers mounted on four of these seven modules to monitor their vibrational response. The LIMs which were dynamically simulated (i.e., equal Strouhal number) are those located nearest to the reactor vessel Inlet nozzles, because they are subjected to the highest cross-flow velocity and thus most susceptible to vibration. Results from the IPFM test series are summarized in Section 3.9.1.3.2.

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b. Bench Tests Based on component evaluations, an experimental progran to evaluate the flow induced vibration characteristics of selected regions of the Upper l I

Internals Structure sas established. Testing was performed by Argonne National Laboratory in a 1/3 scale f acility which structurally and hydraulically simulated the Instrumentation post and a chimney with a lower shroud tube.

Results of these bench tests are given In Section 3.9.1.3.2.

c. Full Scale Model Testing it was anticipated that during the perf ormance of the experimental program, the need for additional, selected f ull scale model tests might be identifled. For exampie, it became necessary to conduct seiected f ulI scale tests on portions of the UlS where similarity criterla or Reynolds number simil arity could not be satisfied by scale model testing. These f ull scale tests include the gap modulation flow induced vibration test of the control rod upper to lower shroud tube slip joint, the UlS chimney vibro impact test, and the IVTM port plug gap modulation flow induced vibration test.

In addition to these full scale tests, a special test is being conducted on the IVTN Port Plug in conjunction with the gap modulation flow induced vibration test. During the test it is planned to assess the suscep-tibility of the IVTM Port Plug to self-excited vibration. If the levels do not increase, it will be a positive indication that the potential for sel f-excited vibration does not exist.

Finally, full scale testing of a prototype Primary Control Rod System - a Primary Control Rod Drive Mechanism, Primary Control Rod Driveline, and Primary Control Assembly in sodium at 4000F - has been perf ormed. The Primary Control Drive (PCRD) was instrumented with three pairs of 3.9-1f Amend. 69 May 1982

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Pcg3 - 6 [8,22]i23 accelenncuneters to measure its flow induced vibration response. Emphasi s on the PCRD was at the dashpot / cup area. The PCA was Instrumented with four accelerometers which were positioned at two elevations on the outer duct of the assembly. The in sodium tests were conducted in an aligned and misaligned configuration, the test variables for a given configuration being rod withdrawal height and prototypic sodlum flow through the PCA and the shroud tube enveloping the driveline. Results of this testing are given i n Section 3.9.1.3.2.

d. Integral Reactor Flow Model (IRFM)

The outlet plenum region of the reactor contains components which could potentially have flow induced vibration problems. Therefore planning for a 1/4 scale flow and vibration system model of this region was initiated in 1975. The objectives of the IRFM test progra are as f ollows:

1. Measurement of the velocity pattern in the outlet plenum and in the vicinity of major outlet plenum structures to provide Input for the prediction of flow induced vibration.

2. Evaluation of flow Induced vibration characteristics of selected outlet plenum structures.

3. Evaluation of flow Induced vibration characteristics of the primary and secondary control rod drivelines, it was decided that to satisfy these objectives the model test progrm would be conducted in two phases since the design of all outlet plenum components was not finalized. For the initial phase of IRFM testing, Phase I, approximately 100 accelerometers were located on critical regions of the outiet plenum components as shown in Figure 3.9-5. The Iocation of this instrumentation is based on modal analyses of these components.

Table 3.9-10 summaries the Instrumentation planned f or the Phase ll vibration tests in IRFM.

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k2g N 7 [8,22]i23 in order to satisfy the objectives of the test progran the foilowing tests were performed during Phase I:

1. Yelocity Test (outlet plenum for both three-loop and two-loop operation, inter-chimney region, chimney, core assemblies - upper internals).

2. Flow induced Vibrations (preliminary evaluation of current designs).

A flow step summary of the flow induced vibration test is provided in Table 3.9-6. As the table shows, vibration data was obtained on the model for the f ull range of ant!cipated flow including the refueling mode. in addition to the flow induced vibration data, experimental modal analysis (shaker tests) of the UlS and outlet plenum internals was conducted for two conditions: (1) the model dry in air and (2) no flow but rnodat vacsel fulI of water. For the Phase il testing, the flow induced vibration test sequence will be repeated.

The shaker tests will also be repeated for the final designs.

Model DescrIntion The IRFM is an approximately 0.248 scale model of the wetted surf aces in the CRBRP reactor outlet plenum. It is a 3600 model, including three outlet nozzles as in the CRBRP reactor, and is capable of two or three loop operation. The vessel support is not prototypic, but is provides sufficient f requency separation between support, vessel and components.

Both hydraulic and vibrational testing are performed in IRFM. However, since flow induced vibrations are of concern for only certain outlet plenum structures, only those structures will reflect hydraulic and structural simulation. Table 3.9-8 is provided to define the components which require one or both types of simulation.

in the design et the IRFM, provision was made for simulating the height variation of the reactor assembly exit nozzles. Provisions were also made 3.9-lh Amend. 69 May 1982

.. oss, Pcon - 8 [8,22]i23 for simulating misalignment between the UlS and the core. Two height and two radial alignment variations have been tested. Also, the ref ueling position mode, in which the upper Internals structure is raised 9-1/2" and decoupled f rom the lower Internals, has been tested.

For the Phase I series of tests the following components were modeled in IFRM:

o Upper Internal . Structure o Instrumentation Post o Instrumentation Conduit o Liquid Level Monitor o IVTM Port Plug o Ex-vessel Transfer Machine Guide Tube o Upper Control Rod Shroud Tube a Control Rod Driveline The second phase of IRFM testing structurally simulates the final, released designs of outlet plenum hardware. Phase 11 model simulations provide data on prototypically modeled components prior to reactor operation and will be used to predict prototypic vibration of outlet plenum components. To date the IRFM has been modified to include the thermal liner, the outlet nozzle !!ner, a redesigned Liquid Level Monitor Port Plug (LLMPP), and a dynamic simulation of the vortex suppressor plate. For the final series of IRFM Phase ll tests, the simulation of the UlS is being modified to include dynamic simulation of the following: two chimneys, the thermal liner plates in the mixing chanber, and the UlS Jacking mechanism method of UlS column support.

In addition, it is planned to locate four blaxlal accelerometers on the UIS model to monitor its gross motion during a simulation of the reactor's preoperational test program. The accelerometers are located on the model in the identical positions of the plant accelerometers. By correlating the data f rom these two sets of accelerometers, predictions on the motion of the prototype unit can be made f rom data obtained from the highly instrumented IRFM model.

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Pcg3 - 9 [8,22]i23 Vibration Modelina Considerations The similitude requirements for valid flow induced vibration model testing are general ly wel I known (see Ref s. I and 2).

Similitude ratios which are pertinent to flow Induced vibration modeling are summarized below. Subscripts m and p refer to model and prototype, respectively.

(Re)m/(Re)p, where Re is the Reynolds number (DU/y),

Sm/Sp, where S is the Strouhal number (fD/U),

is the structure material density and p is the (fs/)m/(Ps/)p,where#fluidd8nsity,and P P (Sm/6p), where 6 is the log decrement.

For a given structure in incompressible flow, and ignoring (a) externally driven boundary motions, (b) externally generated forces (such as pump pul sati ons), (c) surf ace wave of f acts (Froude number), and (d) surf ace tension ef fects (Weber number) the dependent paraneters of interest are a function of four main Independent parameters.

y/D = p $

tt 3 P Y

) 3 ,

where y = vibration amplitude D= characteristic length (diameter) f n= natural frequency U= flow velocity I

l Ps= density of structure P= density of fluid V= kinematic viscosity 6= mechanical damping, log decrement l

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Pcgs - 10 [8,22]#23 An equivalent set of parameters is C 7 y/D . p K A_ Dit

_ (A2 D ) r3 Y) pud 2 ;

where K = spring rate, Ib/in C = mechanical damping constant, Ib-sec/in For geometrically similar structures, the scaling law between model and prototype using subscripts m and p respectively, is p m This relattonshIp wllI hold when the four similitude paraneters are the same

'- --dal and prototype.

Table 3.9-7 shows the flow vibration paraneters and associated ratios for the model and prototype, in the following paragraphs, the similitude ratios are discussed as they pertain to the IRFM simulation. The model fluid will be water at 1000F and the prototype fluid is sodium at 9500F. The structural material is stainless steel.

Reduced Frecuency and Density 0

Using stainless steel and a model temperature of 100 F wilI result in com-ponent stif f nesses which are approximately 24% greater than properly scaled val ues.

Because the stif f ness is about 24% higher, the vibration natural frequencies will be about 11.4% higher as a result. To obtain the same values of reduced frequencies f the model flow velocities must be 11.4%nD/U in the model as in the prototypehigher than the prototyps flow velocItles. Th will be properly scaled, but the fluid densities will not.

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uew n Pags - 11 [8,22]i23 Under these conditions, the fluid elastic parameter K/pu2D wil l be .83 times the prototype value, which is conservative.

Revnolds Number A modeling scale Dmodel/D prototype = 0.248 has been chosen. A flow velocity ratio U conditio*8dal/U prototype = 1.1 was chosen, corresponding to an operation D in model and giving the same redetced frequency of structural vibration, fUnder prototype. rs the.se conditions the ratio

- Du/v) wilI be:

Remodel/Reprototype =0.1 ,

Reynolds number will not be duplicated in the model. This is not considered to be significant except possibly in a limited flow / vibration regime. This is f urther discussed below.

Circular cylinders in steady cross flow exhibit regular vortex shedding with well defined shedding frequency in the Reynolds number redime 801Ref.5 x 105 (subcritical) and Re 1 3.5 x 106 (transcritical), where Re = DU/V is based on the cylinder diameter D and cross flow velocity U. For 5001Re13.5 x 105 the sheddingffrequency whereas or Re 13.5fsxis10given ,f =with

.27 fU/D.

airly good accuracy In the by fs subcritical = .2 U/D Hz*,

and transcritical regimes, coincl@ence between vortex shedding f requency sf and structural vibration natural frequency f can forces and large lateral vibration ampil9udes, give in 1Serise to larheical supercrl alternating regime, 3.5 x 105 1Re13.5 x 10 , any shedding is irregular and the resulting 6

vibration excitation forces are random. Therefore in the supercritical region there is no possibility of driving frequency coincidence with the structural vibration natural frequency.

On account of the foregoing considerations, it is important to identify the Reynolds number regimes of cross flow past cylindrical structural components of CRBRP prototype and flow vibration models.

• In the formulas for shedding frequency, the units of U and D are such that U/D has units of sec~l.

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Pcga - 12 [8,22]#23 Values of Reynolds numbers in the model and in the prototype, determined f or the af orementioned assumed conditions are given in Figure 3.9-3 as f unctions of the product (prototype cross flow velocity) x (prototype cylindrical-component dieneter). Also indicated in Figure 3.9-3 are the limits defining the supercritical regime of random excitation, it can be seen that the model will be conservative f or subcritical flows but non-conservative f or trans-critical flows. That is, there is a DU regime (11 in-ft/see SDU,1107 in-f t/sec) where vortex shedding frequency coincidence that might take place in the model would not take place in the prototype; and there is another DU regime (116 in-f t/see SDUI 1200 in-f t/sec) where vortex shedding frequency coincidence that might take place in the prototype would not occur in the model. CRBRP model vibration test results must be viewed with this point in mind. The mismatch in Reynolds numbers emphasizes the need f or analytically examining upper Internals components f or susceptibility to vortex shedding

---t+.+[on, Density Ratlos The value of density ratios, (ps/pJ/(Ps/p f rom Tabl e 3.9-7 is 0.851. Thus the ef f ect of the virtual mass of the flu d,is somewhat greater in the model,,

yielding lower ef fective model natural frequencies. The fluid excitation of the structure by the fluid is more ef fective in the model, yielding somewhat greater amplitudes. Both ef fects are small and considered conservative.

Vibration Disolacements With respect to the modeling based on the requirements that the ratio of model-to-prototype Strauhai number be unity, the previously cited regimes of DU will establish the scalability of model results. In those regimes where the model is conservative, the ef fect of density ratio damping should be f urther conservati sm. Then the ratio (y/D) ( y/D) is considered I conservative. The model results obtained l$ =he t nonEconservative regime and

1. 1 are not directly scalable to the prototype, and the the regime results wherein must be SdS@her Fur analyzed based upon the test circumstances to estabilsh applicability.

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w.~u.a.o o Pcg3 - 13 [8,22]i23 Model to Prototvoe Scaling Ratlos Based upon the values of Table 1 and the geometric scaling ratio of 0.248, the following are model-to-prototype ratios of measured parmeters:

f m/fp = 4.432 (frequency)

Am/Ap = 0.248 (displacement)

F m/Fp = 0.076 (force)

'Am/Ep= 4.871 (acceleration)

To aid concentration of the testing on areas of significance, guidelines have been established for vibration levels rquirir.g detailed measurements and assessment. If the vibration measurement for a component show eithcr acceleration levels greater than 0.3g's, occurrance of impacting or a flow rate dependent resonance, the component data is assessed for potential design impact with additional measurements perf ormed if necessary to obtain con-clusive data, if neither of these of fects occur, the component does not require f urther vibration assessment.

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Ac;e - 14 [8,22]#23 3.9.1.3.2 Test Results A. IPFM Data f rom the IPFM tests show the maximum prototypic vibrational amplitude of a LIM is 0.2 mils. This value is very low and is structural ly acceptable. The tests have also shown that dynamic coupling between the Core Support Structure and the LIMs is negilble.

B. Bench Model Tests Test results from the 1/3 ANL scale model of an instrumentation post Indicate vibrational response is extremely small, less than 0.4 mils in bending. Testing of an outer chimney (chimney located above a k8 anket reglon) showed small vibration amplitudes and only slight rattling between the chimney and its support. Testing of a central chimney with a shroud tube (chimney located above an active region of the core) Indicated simil ar results.

C. Full Scale Model Testing o The upper to lower shroud tube slip joint test have been success-fully completed with no excitation mechanism discovered for simulated prototypic leakage conditions.

o The Uls chimney vibro-impact test has been completed and the data is being evaluated. Preliminary evaluation of test data Indicates that no gross motion of the chimney occurs at flow rates up to 95%

of design flow rate. At 95% of design flow rate, signals from displacement transducers Indicated fluctuations. It does not appear however that impacting is occurring at the 95% flow condition. Additional data evaluation Is being perf ormed and the model will be visually inspected for signs of impacting at di sassembly.

Primary Control Rod System Tests o Low energy level flow !nduced vibration occurred across a broad i

f requency spectrum.

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P g$ I 15 [8,22]i23 o Acceleration levels increased according to the square of flow velocity without Indicating any resonant peaks or Instabilities, o Vibration behavior of the system was not significantly af fected by either rod withdrawal height or misalignment, o Low level flow induced impacting did occur in both the dashpot and control assembly areas. The acceleration magnitude and rate of impacting also increased approximately as the square of the flow ,

rate.

D. IRFM Phast i Test results of the Phase I tests can be summarized as follows:

o The measured responses of Instrumented components were generally smal and proportional to the flow rate over the flow regime tested, o There were no observed unstable dependencies upon flow rate nor abnormally high fluid excited f orced components.

o There were no observed occurrences of vortex shedding synchronous with component resonant frequency over the f ull flow range tested.

o The vibrational characteristics of the UlS and Instrumented outlet plenum components were essentially Independent of core config-urati on, loop mode operation, and UlS/ core alignment.

o impacting was observed on the following components, with the location of impacting coinciding with assembly gaps incorporated in the prototype design and scaled in IRFM; UIS Kevs/ Core Former Rino At 110% flow a maximum impacting of approximately 2.1 g's peak-to-peak was measured on an accelerometer mounted on the model core f ormer ring radial ly outward f rom the keyways. The cyclic rate of i

this impact was less than 1 cycle /sec. The corresponding g value in the plant is 0.43 g's.

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UlS Unoer Shroud Tubs / Lower Shroud Tube At 110% flow a maximum impacting of up to 3.5 g's p-p at a cyclic rate of 4 to 5 Impacts /sec was detected. These measurements were recorded from a shroud tube containing a control rod driveline.

Another shroud tube without a CRDL exhibited impacting level of 3 g's p-p with a cyclic rate of 2 Impacts /sec. These g levels when scaled to the reactor are 0.71 and 0.62 respectively. Based on these results, a f ull scale model test of the upper to lower shroud tube joint was perf ormed f or the final design configur-ation. As noted in C above, no vibrational problems were found.

Chimnev/ Solder / Lower Shroud Tube This region was not directly Instrumented f or Phase I but responses on neighboring accelerometers Indicate impacting was probabl e. The results of the f ull scale chimney test are described in C above.

UlS Column / Closure Head Indications of Impacting were detected with maximum levels of response less than 1 g p-p and a cyclic rate less than 1 Impact /sec. The corresponding g value for the plant is 0.21 g's.

Control Rod Driveline The control rod driveline response was acceptable at 100% and higher fluid velocities. Impacting at the dashpot /plsion Interf ace was infrequent.

IRFM Phase ll Results from the completed portions of the Phase ll testing are as f ol lows:

o The measures response of the Phase ll components was generally small.

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us esso Pcge - 17 [8,22]i23 o The termal liner, outlet nozzle liner, and the Liquid Level Monitor Port Plug measured responses exhibited no unstable, flow rate dependency nor excessively high fluid excited response over the entire range of flow rates tested. Suppressor plate measured response was generally consistent and small over the range of test flow rates. The measured displacement amplitudes were less than 0.5 mils rms (model scale) at the maximum observed response l evel s.

o The gross motion of the UlS during simulated ref ueling conditions (UIS raised to remove keys f rom the core former structure) was generally of low amplitude and erratic. The dominant measured lateral frequency was 12 Hz with calculated displacements on the order of 0.1 to 0.2 mils rms (model). Infrequent large amplitude responses were recorded. The maximum peak model displacement at the lower end of the UlS during these large unplitude responses was 16 mils. The motion was very sporadic and random occurring once every two to five seconds on the average.

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Pigo - 18 [8,22]i23 3.9.1.3.3 Application of FFTF Experience to CRBRP The CRBRP reactor Intervals vibration progran is similer to the FFTF vibration progran and was formulated to maximize use of FFTF experience. Both prograns utilize a combination of analysis, scale model tests, feature tests and selected in-reactor vibration monitors. The FFTF Hydraulic Core Mockup (HCM),

a 0.285 scale model, was designed to simulate the vibrational and hydraulic characteristics of the reactor system just as IRFM does for the CRBR.

Vibration measurements were obtained in the HCM and found to result in acceptable vibration levels.

FFTF in-reactor vibration monitoring of selected components was performed during pre-operational acceptance testing for confInnation of the scale model test conclusions. Pre-operational, non-nuclear, in-reactor vibration tests

'~'"Md accelerometer instrumentation in the Instrument Tree Guide Tubes (lGTs) and the Vibration Open Test Assembly (VOTA). Upon completion of the non-nuclear acceptance tests, IGT accelerorneters were replaced with normal plant instrumentation for nuclear operation while the VOTA instrumentation was retai ned. Although the HCM and FFTF tests do not have Identical Instru-mentation locations to permit direct one-tor-one comparison, the FFTF data permits conclusions on the following points: 1) Overall vibration conclusions f rom both FFTF and HCM; and 2) Comparisons of HCM scale model test results with FFTF measurements for similar, although not identical, locations.

The overall conclusions from the HCM and FFTF tests are:

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us ss, Pags - 19 [8,22]i23 o Overall test results f rom HCM and FFTF were similar and in f air agreement with respect to frequency and rms for similar instrumented components.

o No gross vibrational problems such as significant impacting were observed in either HCM or FFTF.

o The measured responses were stable and a f unction of increasing flow rate.

o There is no incidence of syncronous vortex shedding at component natural frequencies.

FFTF In-reactor vibration measurements were obtained from two Instrument Tree Instrument Guide Tubes (lGTs) represen+!a; +ha shortest and longest IGTs.

Accelerometers were mounted on a flow and temperature removable instrumentation assembly inserted in the IGTs. The HCM IGT accelerometers were mounted externally on the IGTs. However, the HCM lGTs were not prototypic of the final FFTF design. Figure 3.9-7 compares the first mode rms displacements f or the two shortest HCM IGTs with the shortest FFTF IGT. The HCM rms displacement predletion compares well with the FFTF results and are conservative. The HCM frequency prediction is quite good and consistent with the s'Ight length dif ferences. For the longer HCM IGTs and the longest FFTF IGT, wie f requency and flow dependence are similar in trend with the HCM rms displacement predictions being non-conservatively smaller (approx. 7 mils max.) than FFTF (approx. 30 mils max.). The non-prototypicality of the HCM IGTs is expected to be more influential for the long IGTs and could be the cause of the non-conservative HCM results in this case.

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Poca - 20 [8,22]i23 The FFTF Vibration Open Test Assembly (VOTA) was designed primarily as a vibration sensor and is a 40 foot assembly spring loaded into a core assembly 1 receptacl e. The HCM Closed Loop in-Reactor Assembly (CLIRA) is similar to the FFTF VOTA. The HCM CLIRA accelerometer measurements were scaled to FFTF and modified for dif ferences in accelerometer locations based on analytical mode shapes for fully constrained and partially constrained support conditions.

The isothermal VOTA tests showed a fIrst mode response of 4 Hz indicating partial constraint. Figure 3.9-8 compares the FFTF YOTA isothernal results wIth the projected HCM results for partial constralnt. The agreement Is quite good. During FFTF power ascent, the VOTA fIrst mode shIf ts from 4 to 10 Hz due to increased core cleping. The VOTA results for this condition compared to the HCM CLIRA projections for fulI constraint show that the HCM results are conservative (approx. 3 mils rms displacement compared to 1 mil for VOTA).

erTr vibration comparisons confirm that the HCM scaling parmeters (similer to IRFM vibration movel parameters given in Section 3.9.1.3.1.4) generally result in the expected conservative predictions for the reactor internals. The agreement between FFTF and HCM results is quite good for both overalI vibration conclusions end for comparable vibration magnitudes.

3.9-lhm Amend. 69 May 1982

k g N 21 [8,22]i23 3.9.1.3.4 Overall Test Conclusions The good agreement between FFTF and HCM test results support the validity of well designed scale modal testing for predicting reactor Internals vibration response. Acceptable comparisons between FFTF and HCM were obtained for vibratton displacoments, fIow dependence and f requencies wIth both tests showing low vibration levels for FFTF. l The CRBR reactor Internals scale model and f ull scale test results complated to date show no indication of any significant vibration problens. No unstable dependence upon flow rate nor highly fluid excited forced components have been observed. Vibrationally induced impacting at gaps has been shown to result in low g level impacts judged to be below material damage thresholds. Although some testing remains to be complated for final design configurations, the ar** arava = between the preliminary and final test configurations are not likely to cause a major change in the vibrational responses for the outlet plenum components. Where the acceptability of scale model testing might be questionable such as f or flow thru small clearances (UIS shroud tube gap, IVTM ,

port plug, primary control rod driveline), full scale tests were performed to I 1

f urther assess vibration potentials.

in general, the test configurations were constructed to maximize the gaps between components. In the CRBR, misalignments f rom normal manuf acturing tolerances (UIS chimneys, shroud tubes and IVTM port plug for example) and thermal gradients (UIS mixing chamber liner plates and the horzontal bef fIe for example) tend to close gaps and minimize vibration potentials.

l

3.9-lhn Amend. 69 May 1982

. w.

Pcg2 - 22 [8,22]i23 Both the FFTF and CRBR prereactor operation test prograns have shown no significant vibrational problems for the reactor internals. The CRBR plan for in-reactor confirmation of the test results is similar to the FFTF In-reactor program of using in-reactor vibration monitoring of selected components to confirm the pre-reactor operation test and analysis results.

The FFTF VOTA accelerometers in the core regions exhibited excessively large reciprocal of frequency (1/f) noise characteristics starting at very low power levels (<1%). Accelerometers above the core had minor low frequency degradation as a result of elevated temperature exposure (<10000F). The mechanisms f or these of fects are not fully understood at this time. FFTF utilized piezoelectric, lithium niobate crystal accelerometers for the in-reactor tests which are the same accelercmeters planned for CRBR in-reactor testing. Based on these FFTF results, the accuracy of CRBR measurements under

t power conditions cannot be determined.

l l

3.9-lho Amend. 69 May 1982

a ou, Prga - 23 [8,22]#23 l 3.9.1.3.5 CRBRP In-Reactor Vibration Monitoring for Reactor Internals The results of the test and analysis progran completed to date have shown the adequacy of the reactor internals for flow induced vibration considerations. ,

To supplement this comprehensive vibration assessment progran and to provide confirmation of scale model test results, vibration monitoring of the UlS is j planned for CRBRP. While the test results Indicate acceptable vibration ievels for alI reactor internals, the princtpal area of concern f rom an i inadequacy of the pre-reactor operation test progran would be gross vibratory j motion of the UlS. Conceptually, gross vibratory motion of the UlS could 1 cause impacting of the keys to the core former structure keyways with resulting f attgue f alIures at the keys or other gap' elements. Based on the vibration test results, other areas of the reactor Internals are even less sensitive to vibration than the gross UlS motion.

i Four blaxlal accelerometers wilI be located on the UlS as shown in Figure l 3.9-6. As can be seen f rom the accelerometers' orientation, lateral, vertical and torional motions of the UlS or impacting at the UlS keys can be monitored with the accelerometers.

Accelerometer measurements wilI be recorded and evaluated at varying flow rates during pre-cperational testing. Acceptable accelerometer performance at temperatures above about 6000F and at reactor power conditions cannot be assured. If the accelerometers remain operational with acceptable accuracy,

] data wilI be obtained and evaluated during the initial power ascent.

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3.9-lhp

, Amend 69 i May 1982

.~.

pig 3 - 24 [8,22]i23 Phase 11 of the IRFM scale model tests has four accelerometers in the UlS at the same location as the CRBRP UlS. These IRFM acceleromaters wilI be correlated with other extensive Instrumentation on the IRFM UlS to provide a ref erence data base for interpretation of the CRBRP UlS accelerometer measurements. Identification of the peaks and frequencies in the IRFM tests will provide guidance f or assessment of the CRBRP measurements.

Vibration data from the final series of IRFM tests will be evaluated to demonstrate component acceptability and to develop acceptance criteria for the CRBRP Internals vibration tests. Components with significant vibration in the IRFM tests wilI be evaluated to show stress, f atigue and wasr acceptability including allowances f or test uncertai ntles. These analyses will be used to develop acceptance criteria for the plant tests to assure that Uls structural and f unctional requirements are satisfied. The UlS measurements will be emoared to the acceptance criteria and to the IRFM data. Major differences between the IRFM and plant data will be assessed to assure that the dif ferences are not an Indication of an unexpected, significant vibration problem. Acceptance criteria for the UlS accelerometer measurements wilI be provided in the FSAR.

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l 3.9-lhq

( Amend. 69

us usso Poge - 25 [8,22]i23 3.9.1.3.6 Application of Regulatory Guide 1.20 The following comments address application of Regulatory guide 1.20 to CRBkP.

_Regulatorv Position C.1r Classification of Reactor Internals The classification provided in this regulatory position will be followed by the CRBRP Project to categorize the reactor Internals. It is anticipated that most of the CRBRP Internals will be in the category of ' Prototype."

Reculatorv Position C.2r Vibration Assessment Procram for Prototvoe Internals item C.2.1 (Vibration Analysis Program), is applicable to CRBRP. The scope of the analysis f or the CRBRP is described in Section 3.9.1.3.1 and Chapter 4 of the PS AR.

Item C.2.2 (Vibration Measurement Program) is applicable to CRBRP. The test operating conditions will be provided when the preoperational and initial startup test details are established.

The provisions as given in item C.2.3 (Inspection Program) were developed primarily for LWRs. Due to the limitation of the state-of-the-art of Inspectability of LMFBRs, the requirements set forth in item C.2.3 are considered largely not applicable to the CRBRP.

l The intent of item C.2.4 (Documentation of Results) wil l be met, as l applicable, in the context of the program described above.

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3.9-lhr Amend. 69 May 1982

P:g: - 26 [8,22]i23 ltem C.2.5 (Schedule) is provided herein. Sub-Item I requesting classi-fication of the reactor Internals is f ulfilled by the above " Prototype" designation. Sub-Itum 2 for a commitment on the scope of the vibration assessment program is f ulfilled by Section 3.9.1.3. Sub-Items 3 (description of the vibration measurement phase) and 4 (summary of the vibration analysis program) will be provided in the FSAR as a f urther development of the Inf ormation given in Section 3.9.1.3. Sub-Item 5 requests preliminary and final reports within 60 and 180 days, respectively, of the completion of CRBRP vibration testing. CRBRP reports will be provided consistent with these requested periods.

Reaulatory Position C.3r Vibration Assessment Procram for Non-Prototvgg Reactor Internals

"';, r.o CRBRP reactor internals are expected to be classified as "N on-Prototy pe. d Therefore, no assessment with regard to the applicab!llty of j this regulatory position is attempted at this time.

3.9-lhs Amend. 69 May 1982

TABLE 3.9-6 ,

FLOW STEP SlM %RY Test. o' Test Condition Flow Rate of Fluid Condition

=

" 5 Water (% of CRBRP Rated Velocity)+

t Three loop Operation Air Water 10 33 66 100 110 120* **

Velocity Test X Vihratinnt a) Operating Mode X X X X X X X- X X b) Refueling Mode *** 'X X X 30 Two Loop Operation , ,

Velocity Test X X Flow Vibrations X X X

• Operating at this flow' depends on proximity to vibration limit

• • To be determined, vibration limit or test facility limit

+ For two loop operation, rated flow is approximately 2/3 three loop operation value.

• • • Approximately 11 in, gap between top of reactor assemblies and the upper internals structure for the refue:iej mode.

NOTE: All components sill * .s ;h)nitored and only those shown most responsive will be re.orded and ; 'lyzed. I9 3.9-9c Amend. 30 Nov. 1976.

TABLE 3.9-7 Flow Vibration Parameters - CRBRP Model and Prototype Conditions:

Same structural material in model and prototype (stainless steel)

Model Fluid = water at 100*F Prototype fluid = liquid sodium at 995'F Model Prototype Ratio Parameter Ynerno's Modules of 28.2x106 22.7x106 g Elastic, 7 = 1.242 P

E (psi)

Fluid Density, p 0.0359 0.0297 p,

- = 1.206 (1bs/cuin.) P Material Density, os 0.290 0.282 (ps),,).028 (1bs/cu in.) (ps)p ps/p 8.078 9.495 (,37,9

= 0.851

- r (ps/p)p 0.282 0.3 01 Poisson's Ratio, u p , = 0.937

,P Kinematic Viscosity, v 9.0x10-6 3.05x10-6 "m

= 2.951 l "p 30 i

l k

cpW Amend. 30 N v. 1976 3.9-9d

TABLE 3.9-8 Outlet Plenum Components Modeled In IRFM Test Component Hydraulic Simulation Structural Simulation FIV Instrumentation UlS Structure Support Columns Yes Yes Yes Chimneys Support Plates Yes Yes 2(Ph.II) Yes Yes Yes Instrument Post Yes 3 Yes UIS Lower Shield Plate Yes Yes Yes Control Rod System

• Control Rod Driveline 1 1 Yes Shroud Yes 3 (Ph.I/l (Ph.II) Yes Guide Tube Yes 3 Yes Shroud / Guide Gap 2 2 Instrumentation Conduit 2 2 Yes LLf1 Yes 1 Yes Vessel Thermal Liner Yes Yes Yes Themal Baffle Yes No No In-Vessel Storage 1 No No Suppressor Plate Yes Yes Yes Outlet flozzle Liner Yes Yes Yes Reactor Fuel, Blanket and Radial Shield Assembly Yes* Yes** Yes**

Control Assembly Yes* Yes flo Core Barrel Yes No No Core Barrel UIS Keys Yes Yes Yes Miscellaneous EVTM Guide Tube Yes Yes Yes IVTil Port Plug Yes Yes Yes

• Flow Distribution and Assembly Exit Geometry need only be simulated for fuel, blanket and control assemblies. For radial shield assemblies, only outlet. flow distribution will be simulated.

• • Testing Involving Core Dynamic Simulation may be performed at ANL rather than in the IRFil.

3 ') .- o e

- n-. - -

Table 3.9-9 FLOW Vlt4ATION OF CAIRP INT (llNALS General Survey PETM005 5 ELECTED IMPORTANCE INDEI for AtS& UTION Reactor Internals Type of Analysis IRTM Testing Full Scale Bench Components (scitation Mign Medium Ls" Test Test 1 [ntire Upper Internals Wortes Shedding I I I 5tructure Jet lapingecent I I (crossmotion) Jet Reaction I I Turbulence 1 I Pep Pulsations I I I

2. U15 Guide Tubes for Vortes Shedding I I I Control Rods furt*ulence I I I I Pep Pulsations I I I Cap riodulation I I

3. U15 CManeys Vorten Shedding I I I Turbulence I- I I Pump Pulsations I I I I I Jet Reaction 1 I I s "'S $wapart Col e s Vortes Shedding I I I Turbulence 1 I I Pump Pulsations I I I

5. U!5 Instrirentation Vortea Shedding I I I Leeds Turbulence I I I Pep Pulsations I I I

6. Ul1 Support Plates Vortes $heking I I and Plate Liner Turbulence I I Pro Pulsations , I I

7. U15 te. strum ntation Vorten Shedding I I I Posts Turbulence 1 I I Jet Imptngemnt 1 I I C. U15 Cylindrical Vortes She&fing I I Liners Sleeve Turbulence I I

" Ease tiotion* I I

9. U!$ IVTH Port Plug . Vortea shedding I I i I j --

Turbulence 1 I I* I P e p Pulsations I I I g Gap Ptdulation I I I

10. Control Rod Vorten Shedding I I I I I .

Drivelines .', rbulence I I I  !

Ga6 'tdulation l I l

11. Core Support $tructure Turbulence I I i Core Barrel Pep Pulsations I I l ,

12. Core Support Structure Vorten Shedding I I I ftdule Liners Turbulence 1 I I Pump Pulsation 1 I I

13. tiorizontal 8aff te Vorte Shedding I I 8 Turbulence I I l Pop Pulsation I I i 14.bnGuideTute Vorte Shedding I I I Turbulence I I P e p Pulsation I I

15. Suppressor Plate Vortes Shedding I I I Supro* ting Nangers Turbulence I I I Pump Pulsation I I I

16. Suporessor Plate Vorten 5hedding I I I 5?g* nts Turbulence I I I Pump Pulsation I I I

17. Liquid Level itnitor Vorten Shedding I I .I (LLM) Tubes and Turtulence I I Support ficacers Pe p Pulsation I

10. Vessel Thermal Ltner Turbulence I I I Pump Pulsation I I I

19. Outlet t'ozzle Liner

. Turculeace I I I l

Pumo Pulsation I I I l

Jet Reaction I I I 1

, 3.9 - 9f l

l

Table 3.9.I0 IRFM PHASE 21 VilRAil04 f[51 IN51RLMENTAi!04 5tp9tARV lastrumentattan Quantity Instrumentation location Compoaent Concern Data Reevired _Olsplacement Acc elerometers 5 train Otsplace w nt A celerometers 5 train C m ments

1) U15 seys impact Wear Otsplacement . 6 6 8 RT RT See Strain gages appropriately Aceleration . Comments located and calibratee - need (0* & 120* 5 tress 4/tey

& 240*) force

2) ITTM Port impact Wear Displacement - 3 11 - R-T and RT ime strings of four accelere-Plug Aceleration . 0*. 180*. 300* 0*, 30* meters along length Force

3) UIS Chimney Esttation due kceleration. 8 8 R-T Top. R-T Midspan. Two structurally st= 1sted UIS notton. Olsplacement Bottcss Support Top chimneys Structural

4) Ul5 Colum Mode shape Strate 8 Two strings of four accelere.

meters along length f See hate 1 force transducers at top of

5) U15 Colum impact and Olsplacement 2 4 to Bearing 8 earing toad Force (Force bearing race en one column) ,

l h Trans) See hate 1 force transeucers en structure l

l

6) U15 Column Impact . Force 3 l perpendicular to plP aats to Force (Force 1

Trans) See hate I I

\S Structure l tb 2 R-T I F) co1ven to Vibration kceleration l

Ul1 CRIM l

Plug *l

8) Ul5 tDper/ Impact - kceleration - 2 10 R-Y R-T Tee strings of four accelere.

Vlbration Otsplacement . fhHis meters along length 0*. 90*

tower shroud to include tastrumented Shroud Shape drivellae and dash pot to montter their response to shroud esitation g) lastrument Vibretton kceleration 4 A-T Tip Post

10) U15 Top Vibretton kceleration i V Center Plate ll) Sottre Vibration kceleration I V Center Plate kceleration A study is required to estat-

12) U15 Thermal Vibretten Ilsh the modelleg for tats Liners component

13) Prototype kceleration 4 Stastal accelerometers located same as shown en U15 draulag.

Instrumen-tation Note 1: Tbts instrumentation to be aligned ulth same ants.

11398-1355:2

, (53300) 18

P 107 5 = 0.27 (TRANSCRITICALI i

i i I

I PROTOTYPE l

106 _.

l 1

cz: SUPERCRITICAL l f I 1 1 3

1 2

m -- - - -- e cm g S = 0.20 ' I lI 2 ISUBCRiflCAL) gl l

g i il I

i MODEL Il .

le 105 _ i ;i l l I

l Il .

(

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\ 3 Ii 18

. I II II II I

I u II, ml' I- MODEL 11 MODEL CONSERVATIVE UNCONSERVATIVE f 4 lI' l 'l 10  !

1.0 10 100 1000 l

D U (in.f t/sec) (PRO TOTYPE) i A &  !

u F:yurc 3.9 3. Rey nnids Number for Protoi> pe i

c..

• Arnend. 30 3.9-12 .. Nov. 1976

E P.EACTOR DESIGN i

d NOT q ACCEPTABLE U

! ANALYSIS CRBRP

_ PROTOTYPE _

PREDICTION OF ACCEPTABLE CRBRP

~ -

ll SCALE MODEL CRBRP VIBRATORY OPERATION TESTING MOTION

!d EVALUATION OF P d d SUSCEPTABILITY TO

j l FIV FFTF PROTOTYPE SCALE MODEL TESTS & ANALYSIS ChBRP PROPERATIONAL TESTING i

/ 5 1b  ;

{ / l _ FFTF PREOPERATIONAL TESTING Typ Figure 3.9-4 Reactor Internals Flow Induced Vibration Program

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.!cdel and Instrumentation Amend. 43

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b' 90Sc.i ;.icure 3.9-6 Location of Accelerometers on Upper Internals St ucture

.. . . , . -'- Arner.d. 43 Jan. i978

s. .

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FTR-VO.TA STALK (ISOTHERMAL FLOW) RESPONSE  :

. vs PROJECTED HCM-CLIRA RESPJNSE .

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'v^"V^'S'  :

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FIGURE M. FTR-VOTA Stalk (Isothennal Flow) P,esponse vs. Projected ilCM-C'I':A Response with Partlal Constraint ~

~. 3 ..

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e HCM/FTR INSTkUMENT TREE CSHORT)

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tl

P;ge 3 (82-0287) [8/22]#34 Ouestion CS270 12 Provide an snendment that discusses the procedure that will be used to combine the equipment loading f rom earthquake, generally in the O to 33 Hz range, and f rom the dynamic shock wave loading resulting f rom Ne-H 9 0 reaction with a  !

f requency range of 20 to 100 Hz. Equipment to be quellYled by tests should be specif ically addressed.

Resoonse The combination of SSE and OBE loads with events is described in PSAR Section 3.7A which meet the principle and Intent of Reg. Guide 1.48.

CRBRP has Identified a postulated sodium / water reaction as an emergency event.

The Sodium Water Reaction Pressure Relief System (SRWPRS) as been designed f or tne emergency event. The design utilizes passive devices (i.e. rupture discs, piping and tanks). The rupture discs are designed to remain Intact for seismic events, but will rupture under the pressure loading (shock wave) associated with a large sodium / water reaction. Analysts and tests are being perf ormed to assure the discs will rupture only when required. (See Section 5.5, Ref. 26 and 27) .

The design of the steam generator system is such that the components will accanmodate the earthquake loadings at the end of life without a f ailure that woul d l ead to a sodi um/ water reaction. Theref ore, there is no technical reason to associate the two separate events with the design base.

i I

I l

l l

QCS270.12-1 Amend. 69 May 1982

P ge - 40 (82-0184) [8,223 #38 Ouestion CS 220.34 f a)

There are a number of misprints, unclear statements and typographical errors which need your correction and/or clarification, a) Figure 3.8-9 should label elements discussed in the text (3.8.3.1.1).

Details of concrete reinforcing are needed to evaluate the support ledge.

Resoonse Revised PSAR Figure 3.8-9 labels elements discussed in the text (3.8.3.1.1) .

An additional sketch, showing the reinforcing bars in the reactor cavity at the ledge area has been included.

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l QCS 220.34(a)-1 Amend. 68 N Y,1982

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2'.2" c t FIGURE 3.8-9 TYPICAL SECTION THROUGH REACTOR VESSEL LEDGE SUPPORT u as 3.8-57 1

Pega - 41 (82-0184) [8,22] #38 l

Ouestion CS 220.34 (b) is the load of 50,000 kips mentioned in Section 3.8.3.3.4 to be evenly distributed around the support ledge?

Resnonse The load of 50,000 kips on the Reactor Support Ledge, stated in Section 3.8.3.3.4, has been revised. The revised load is given in terms of time-histories for vertical and toroidal loads. These time-histories are shown in CRBRP-3 ,Section 5.2 (Reference 10a, PSAR Section 1.6).

These loads are axisymmetrical and therefore are applied uniformly around the support ledge.

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QCS 220.34(b)-1 Amend. 68 rtw_RF#2

p:gs 2 WB2-0366 (8,22) 47 The 33% increase in allowable stresses f or steel due to selsnic or wind loadings will not be used.

U - For concrete structures, U is the section strength required to resist design loads and based on methods described in ACI 318-77.

3.8.3.3.2 Internal Structure as containment No portion of the internal structure provides a direct containment f unction.

The embedded part of the steel containment is designed such that it can withstand the design pressure without the assistance of the concrete walls.

3.8.3.3.3 Creen. Shrinkage and Local Stresses No prestressed concrete design is considcrcd f or The design of the f acility.

Theref ore, creep and shrinkage loads will be only considered to the extent they are provided in the reference concrete codes or as may be warranted by prudent design approach. The loads transferred from the support structure that generally influence local areas will be checked to insure that the local stresses are within acceptable limits to preclude impairment of the structural function.

3.8.3.3.4 Loads Due to Structural Margin Bevond Deslan Base (SMBDB)

Ref er to CRBRP-3 Section 5.2 (Ref erence 10a of PSAR Section l .6).

3.8.3.3.5 Sodlum Fire Load See Table 3.8-2 for the accident pressures and temperature loads.

3.8.3.3.6 Hot Sodlum Solli Effect The portions of the reactor cavity and cells, where exposure to radioactive hot sodium is a design basis accident, are provided with carbon steel liners designed to survive a sodium spill (see Section 3A.8). The l iners wil l not compromise gas tightness of the cell.

3.8.3.3.7 Accident Temoerature Load See Table 3.8-2 for design temperatures.

3.8.3.3.8 Negative Pressure on the Liners Any negative pressure on the liner will be resisted by a grid of structural anchors embedded in the concrete.

3.8-14 Amend. 69 l May 1982 ]

Page - 19 (C2-0104) LO,ZZJ F)U Ouestfon CS 220.34 (c)

The text at the top of page 3.8-16 is generally conf using. In particular, the text seems to negate the need f or having both cases 10 and 11. The PSAR should delineate how the appropriate dynamic load f actor will be determined.

Load combination 10 and 11 need more justification for not including T unless A is meant to include thermal effects. in that case, the definition of A in 3.8.3.3.1.4 needs to be changed.

Resoorse For load combinations 6, 7 and 8 in some cases dynamic analysis based on a force time-history input was performed. in other cases, dynamic load f actors are calculated using the procedures described in Section 3.5.4.6 of the PSAR

-' --a consistent with the requirements of SRP, Section 3.5.3. II.B.2.

Load combinations 10 and 11 were intended to cover the cases involving margin events beyond the design basis.

Load combination 11 will be deleted since It is already covered in the last paragraph of 3.8.3.3.10.1.B that states: "Both cases of L having its full value of being completely absent will be checked for." Also, the definition of A In 3.8.3.3.1.4 will be revised to state: "A..... force on the structures due to third level design margin requirement (SNEDB)."

Under SMBDB conditions the thermal conditions are represented by To (Te = To).

The dynamic forces under SMBDB conditions (load "A") are defined as the time histories of the vertical and toroidal moments presented In CRBRP - 3. The effects on the reactor vessel support ledge have been calculated by dynamic analysis.

PSAR Sections 3.8.3.3.1 ~.4 and 3.8.3.3.10.1 will be updated to include the above revised design Information.

QCS 220.34(c)-1 Amend. 68 May 1982

p:g31 22-0366 (8,22) 47 W

T Loads aenerated by the Design Basis Tornado as specified in Section 3.3. They include loads due to the tornado wind pressure, loeds due to the tornado-created dif ferential pressures, and loeds due to the tornado-generated missiles. (Tornado loads do not apply to . internal structures. )

H - Hydrostatic loads due to maximum flood (as defined in Section 3.4).

3 .8 .3 .3 .1. 4 Abnormal Loads Abnormal loads are those loads generated by a postulated accident within a building and/or compartment thereof. Included in this category are the following:

Ia - Pressure equivalent static load within or across a compartment and/or building, generated by the postulated accident, and including an appropriate dynamic load f actor to account for the dynamic nature of the load, Ta - Tharmal loads under thermal conditions generated by the postulated accident and inciuding To, Ra - Pipe reactions under thermal conditions generated by the postulated accident and including R o, A - Force on structure due to third level design margin requirement.

(SH3DB)

Y j Jet Impingement equivalent static load on a structure generated by the postulated accident, and including an appropriate dynamic load f actor to account for the dynamic nature of the load.

Y r Equivalent static load on the structure generated by the reaction on the broken high-energy pipe during the postulated accident, and including an appropriate dynamic load f actor to account for the dynamic nature of the load.

Y, - Missile impact equivalent static load on a structure generated by or during the postulated break, such as pipe whipping, and including an ,

appropriate dynamic load f actor to account for the dynamic nature of th e l oa d.

In determining an appropriate equivalent static load for Y Y and Y elasto-plastic behavior may be assumed with appropriate ducilllty ratio,s and as long as excessive deflections will not result in loss of function of any safety-related system.

3.8.3.3.1.5 other Definitions S - For structural steel, S is the required section strength based on the elastic design methods and the allowable stresses defined in Part 1 of the AISC " Specification for the Design, Fabrication and Erection of Structural Steel for Buildings," February 12, 1959, 3.8-13 Amend. 69 May 1982

vevsuv pcg3 3 22-0366 (8,22) 47

7) U=D+L+Ta + Ra + 1.25 Pa + 1.0 (Yr + YJ + Ym) + 1.25 E

8) U = D + L + Ta + Ra + 1.0 P + 1.0 (Yr + Yj + Ym) + 1.0 E'

9) U = D' + L + To + Ro + H where D' = dead load without hydrostatic load due to normal groundwater

10) U = D + L + To+A in combinations (6), (7) and (8), the maximum values of Pa , T R a Y Including an appropriate dynamic load f actor, will be use$,unfe,ss Yr tfmend a, Y history analysis is perf ormed to justify otherwise. Combinations (5), (7) and (8) will be satisfied first without the tornado missile load in (5) and without Y Yj and Y in (7) and (8). When considering these loads, however, s wa s a,ecNon strength capacities may be exceeded under the of fact of these concentrated loads, provided there will be no loss of f unction of any saf ety-related system.

Both cases of L having its f ull value or being completely absent will be checked for.

3.8.3.3.10.2 Loading Combinations for Steel structures Elastic working stress design methods, as specified in Part I of AISC Specification for the Design, Fabrication and Erection of Structure Steel for Buildings, will be used for design of all Category I steel structures under l both Service Load and Factored Load conditions.

A. Load Combination for Service Load Conditions For Service Loads including earthquake (OBE) and wind loads (if applicable),

the fof lowing ioad combinations wilI be satisfled:

1) S=D+L

2) S=0+L+E

3) S=D+L+W If thermal stresses due to T o and Ro are present, the following combinations will also be satisfied:

Ia) 1.5 S = D + L + +g 2a) 1.5 S = D + L + + Ro + E 3a) 1.5 S = D + L + o+pg+ w Both cases of L having its f ull value or being completely absent will be checked for.

3.8-16 Amend. 69

_ _ _ _ __ . ___ ______________ kw__152

Paga - 20 (82-0184) L8,22J #38 Ouestion CS 220.34 (d)

In combinations (4) and (8) inclusive In Section 3.8.3.3.10.2.B, are thermal loads to be neglected when it can be shown that they are secondary and setf-limiting in nature and or .or where the material is ductile?

Resoonse The second paragraph of Section 3.8.3.3.10.2.B of the PSAR has been updated.

f QCS 220.34(d)-1 Amend. 68 May 1982

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p:g: 4 22-0366 (8,22) 47 B. Load combinations for Factored Load Conditions For Factored Loads including earthquake (OBE or SSE), tornado (If applicable) and pipe break of fects, etc., the following load combinations will be satisfled.

4) 1.6 S = D + L + To + Ro + E'

5) 1.6 S = D + L + To + Ro + WT

6) 1.6 S = D + L + T, + p, + p,

7) 1.6 S*= D + L + Te + Ra + Pa + 1.0 (Yr + Yj + Ym) + E

8) 1.7 S * = D + L + T, + p, + p, + 1,0 ( yr + Yj + Ym) + E'

9) 1.6 S = D' + L + To+Ro + H where D' = dead load without hydrostatic load due to normal groundwater

10) 1.6 S = D + L + To,A in combinations (4) to (8) inclusive, thermal loads may be neglected when it can be shown that they are secondary and sel f-limiting in nature.

In combinations (6), (7) and (8), the maximum values of Pa , T R Y r Y Y Including an appropriate dynamic load f actor, wilI be use$,unfe,ss a+ tide +-

history analysis i s perf ormed to justi fy.

Combinations (5), (7) and (8) wilI be fIrst satisfled without the tornado missile load in (5) and without Y , Yi and Y In (7) and (8). When considering these loads, however,r locdl section strengths may be exceeded under the ef fect of these concentrated loads, provided there wilI be no loss of f unction of any saf ety-related system.

Both cases of L havIng its f ulI value or being compieteIy absent wIII be checked for.

In loading combinations, no load f actors of less than unity will be used in design or analysis.

• For these two combinations, (7) and (8), In computing the required section strength, S, the plastic section modulus of steel shapes may be used.

3.8-17 Amend. 69 May 1982

Yage - 11 (CL-U1 Col LD,LLJ 936 Ouestion CS 220.34(el in Section 3.8.3.4, on page 3.8-18, the figure numbers referenced are incorrect. -

Resoonse The ref erenced f Igure numbers in PSAR Section 3.8.3.4, page 3.8-18, havo been revised.

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QCS 220.34(e)-1 Amend. 66

rcge 29 Lo,zzJF39 3.8.3.4 Design and Analvsts Procedures 3.8.3.4.1 General Analvsts Procedures l

Structural analysis for each cell (except the reactor cavity) will b'e perf ormed by considering one-foot wide vertical and horizontal strips through the structure. These strips, in essence, constitute structural segments which allow analysis by conventional methods.

The analysis of each cell will be perf ormed by choosing one or more sections of one-foot width in both vertical and horizontal directions. The number of strips taken depends upon the configuration of each cell. In the case of analyzing a PHTS Cell, one typical vertical cross section (Figure 3.8-5) and two horizental sections (Figures 3.8-6 and 3.8-7) are chosen for an

.......  ;*vn.

A rigid f rame method of analysis will be used f or determining The moments and shears in all members of a f rame for each of the following anticipated loadings:

a) Dead and Live Loads, b) Seismic Conditions, c) Pressure Loads, and _

d) Thermal Loads. ,

Loading combinations, using the f actored loads as indicated in Section 3.8.3.3.10.1 will be used to establish the most critical design stresses for each component of the f rame.

Vertical and horizontal strips thus analyzed will provide the necessary steel reinf orcement f or the vertical and horizontal directions of the particular cell analyzed.

3.8-18 Amend. 68

_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _May 1982

Page - 22 (82-0184) L8,22J #38 Question CS 220.34(f) in Section 3.8.3.5.2, more description is needed of the " energy absorption

ch eck. " .

Resoonse ,

Ductility ratio, u, is a measure of the capacity of a structure to absorb energy in the plastic range. Energy absorption is satisfied when the calculated value of the required ductility ratio is less than the allowable ductility ratio for the material under a specific loading condition.

PSAR Section 3.8.3.5.2 will be updated to include the above design information.

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QCS 220.34(f)-1 May 1982

us usuu p:ga 5 W82-0366 (8,22) 47 3.8.3.5 Structural Accentance Celteria 3.8.3.5.1 Stress Structures designed by the stress !!mitation methods will be considered acceptable, when design stresses f or the most severe combination of loads are within the limits prescribed by the appropriate codes and standards noted in Secti on 3.8.3.2.

3.8.3.5.2 Strain Since the design of the reinforced concrete structures will be governed by ACI-318, a strain limit of 0.003 used as a basis of this code will be inherently provided in the design. For the steel structures, a maximum stress

'+ ef 0.9 Fy is established. Therefore, strains cannot exceed ninety percent of the yield strain. In specific Instances, where plastic behavior of the steel will be a design basis, an energy absorption check will Insure that the f unctional requirements of the structure are not impaired. Ductility ratio,)u, is a measure of the capacity of a structure to absorb energy in the plastic range. Energy absorption is satisfied when the calculated value of the required ductility ratio is less than the allowable ductility ratio for the material under a specific loading condition. For discussion on liner strai n, see Section 3.8.3.4.

3.8.3.5.3 Gross Deformation The load combinations and stresses noted in Section 3.8.3.3 will Insure that the deformation of a structure will be no greater than that ordinarily permitted for structures of this type. Reinforced concrete structures subject to loads combined with SSE will be nearly stressed to their ultimate capacity.

However, 'several relieving features such as strength gain with age, relief of thermal stresses due to cracking, etc., will preclude any excessive l def ormation. A design check will be perf ormed to insure that the def ormation resulting from SSE and other loads will in no way impair proper functioning of j the critical systems or components.

3.8.3.5.4 Factor of Safety The f actor of safety for the working stress design will be in accordance with the limits noted in Section 3.8.3.3. For the ultimate load design, .oad f actors will be in accordance with the combinations noted in Section 3.8.3.3.

3.8.3.5.5 Shear Resoonse The shear response will be established upon the basis of the classical relationship between the Young's modulus and the shear modulus. The assumed value of the Poisson's ratio and concrete modull will be checked against the properties of concrete as determined through tests of design mixes to be used in the plant construction.

3.8-20 Amend. 69 May 1982

P ga - 46 (82-0184) [8,22] #38 Ouestion CS 220.34(a)

In Section 3.8.3.7, if Internal structures that are designed to hold more than 10 psig pressure are not to be tested at 1.15 times their design pre.ssure, provide justification for not performing such tests.

Resoonse Pressure testing of the IIned cells is not a requirement since the cells do not perform a containment function. A sodium spill occident within a cell will cause heating of the cell atmosphere and structure and a pressure buildup, but there are no requirements restricting cell atmospheric leakage.

The structural design of the concrete walls of the cell, per ACl 349, ensures structural integrity under the sodlum spill accident conditions.

The maximum pressure on the containment vessel under the most severe accident condition will not exceed 10 psi. A detailed description of the containment System is provided in PSAR Section 6.2.

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QCS 220.34(g)-1 Amend. 68

_ _ - . - _ ___. Ctrl M M

Page - 24 (82-0184) L3,22J #38 Ouestion CS 220.34 (h) in Section 3.8.4.4.1, how are equivalent static loads obtained?

Response

Equivalent static loads are obtained by equation P=CqA where P = equivalent static load in Ibs C = drag or Iift coefficient q.= dynamic pressure in Ib/ft 2 A = exposed area in sq. ft.

The procedure in ANSI A58.1 is followed. PSAr. Section 3.8.4.4.1 wil l be revised to include the above design informMion.

QCS 220.34(h)-1 Amend. 68 May 1%2

pcgm 6 W82-0366 (8,22) 47 -

3.8.4.3.2 Creen. Shrinkage and Local Stress Pre-stressed concrete design is not adopted f or the design of the f acility.

Therefore, creep and shrinkage loads will be only considered to the extent they are provided in the referenced codes or as may be warranted by prudent design approach.

3.8.4.3.3 Sodium Fire Load The cells, pipeways, and buildings where sodium fire is a postulated design basis accident, will be designed to withstand accident pressure and the associated temperature ef fects.

3.8.4.3.4 Not Sodium Solli Effect The protective devices such as steel catch per.s or steel plate liners will be provided in floa' areas subject to sodium spills to prevent concrete-sodium reaction.

3.8.4.3.5 Loading Combinations All other Category I structures will be designed and analyzed for the loading combinations listed in Subsection 3.8.3.3.10.

3.8.4.4 Design and Analysis Procedures 3.8.4.4.1 Analysis Procedures Classical theory, equations and numerical methods will be used as necessary in the analysis of the structures. Classical methods used in the analysis will be in accordance with standard textbooks, handbooks, and papers as used in engineering practice. The following computer programs will be used in the static analysis:

1. NASTRAN I 2. MARC CDC -

3. MRI - STARDYNE 4 GT STRUDL

5. Other In-house computer programs Loads and loading combinations as delineated in Section 3.8.4.3 will be considered. For dead loads, live loads, wind loads, tornado loads and accident loads, all of the methods listed above will be used. Wind loads, tornado loads and accident loads are converted to equivalent static loads and will be applied to the structure as uniform or concentrated loads.

3.3-29 Amend. 69 May 1982

psg2 7 W82-0366 (8,22) 47 .

Equivalent static loads are obtained by equation P=CqA .

where P = equivalent static load in Ibs. ,

C = drag or lift coefficient q = dynamic pressure in Ib/ft2 A = exposed area in sq. f t.

The procedure in AfGl A58.1 is followed

.. ... ..; ;c. aado loadings, flood Ioedings and missile loading applied on structures are discussed in Sections 3.3, 3.4 and 3.5.

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3.8-29a Amend. 69

[ May 1982

P;gs - 2D (5Z-U154 J LU,22J F35 Ouestion CS 220.34 (1)

The Section 3.8.2.6 and 3.8.2.7 are missing and should be provided. It appears that your Section 3.8.2.5 should be revised.

Resoonse PSAR Section 3.8.2.5 has been totally revised and PSAR Sections 3.8.2.6 and 3.8.2.7 have been added.

QCS 220.34(I)-1 Amend. 68 May 1982

Pcge 23 LB,22Jr39 The seismic analysis will include the local ef f ects of the air locks vibrating as independent systems. The seismic ef f ects of this Independent vibration will be added directly to all other seismic effects.

The Equipment / Personnel airlock will be supported entirely by the containment vessel shell.

3.8.2.5 Structural Acceotance Celterla The structural acceptance criteria is based on full compliance with the requirements of ASME B&PV CJde, Section lil, Divisions 1 (subsection NE) and 2 as applicable. The requirements of Regulatory Guide 1.57 are included in the structural design criteria. Table 3.8.2-1 from NUREG-75/087 was adopted to define the stress limits for the different load combinatins. Buckling requirements are defined in the " Buckling Criteria" given in PSAR Section 3.u.A.

Applicable codes and standards used in formulating the structural criteria are described in Section 3.8.2.2.

Load combinations for the Containment Vessel are given on Tables 3.8-1, 3.8-la and 3.8-1b. Stress limits are given in Table 3.8-3.

i 3.8 -7 c Amend. 68 May 1982 _

PGge 24 LB,22Ja39 3.8.2.6 Materials. Qualltv Control. and Soecial Construction Technlaues 3.8.2.6.1 Materials 3.8.2.6.1.1 Shell Aourtenances and Structural Shanes The shell is f abricated by the f ull penetration buttwelding of uniform, steel plates of thicknesses that range from 1" to 1-3/4". The shell plates, the circumf erential shell stif feners, the large penetration assemblies, and the polar crane girder are made of carbon steel having the generic designation SA516, Grade 55. Pipe penetration assemblies are made of carbon steel having a generic designation SA333. Welding electrodes for joining these carbon steels are designated as SFA-5.1 having an E70 or E70XX classification. These materials f ully conf orm to Section Ill, Division 1 of the ASME B&PV Code.

The designs of structures, not within the scope of the ASME B&PV Code, conform to the requirements of the knerican Institute of Steel Construction (AISC)

Specifications.

3.8.2.6.1.2 Bottom Liner The Bottom Liner is a 1/4 Inch thick steel plate conforming to ASME SA 516 Grade 70. The structural shapes used for anchoring the liner plate into the concrete foundation, leak-chase channels, angles and back-up bars are ASME SA 36 carbon steel.

Steel plates and structural shapes f ully conform to Section CC-2500 of the ASME, BAPV Code, Section Ill, Division 2. Welding materials conform to Section CC-2600 of the same Code.

For corrosion protection see Section 3.8.2.2.2 3.8.2.6.2. Qualltv Control General Information on Quality Control is given in Chapter 17.

3.8.2.6.2.1 SheII Acourtenances and Structural Shaces Metallic materials used to f abricate the shell and its parts, appurtenances, and penetrations are required to be in compliance with Subsections NA-1220 and NE-2000 of Section lil, Division 1, of the ASME B&PV Code. Metallic materials used to f abricate structures not within the scope of the ASME B&PV Code are required to be in compilance with the American Institute of Steel Construction (AISC) Specifications.

All weld designs are required to be in compliance with Subsection NE-3350 and are to be examined per the requirements of Subsection NE-5000 and as outlined in Section V of the ASME B&PV Code or as outlined in Appendix A of Section ill, Division 1, of the ASME B&PV Code as appropriate. In addition, all Category A and B welds below grade will be examined by a Hallde test.

3.8-7d Amend. 68 May 1982

P;ga 25 Lu,22Jf39 Weld procedures are to be quellfled by compliance with Section IX of the ASME B&PV Code.

Fabrication is to be in compliance with Subsection NE-4000 of Section lil, Division 1 of the ASME B&PV Code.

3.8.2.6.2.2 Bottom Liner Material manuf acturers are required to provide Certified Material Testing Reports (CMTR) for bottom liner material and welding material in accordance with ASME B&PV Code, Section lil, Division 2, Section CC-2130. Bottom liner welds are examined by liquid penetrant or magnetic particle examination followed by vacuum box testing for leak tightness. Non-butt and attachment welds are examined by either IIquid penetrant or magnetic particle inspection methods.

Liquid penetrant examinatin is in accordance with ASME B&PV Code,Section V, Article 6. Evaluation of penetrant Indications and their acceptance standards are in accordance with ASME B&PV Code, Section lil, Division 2, paragraph CC-5544. Magnetic particle examination is in accordance with ASME B&PV Code,Section V, Article 7. Indications are evaluated and accepted in accordance with paragraph CC-5545.

Leak testing of liner seam welds is performed by the vacuum box method using at least 5 psi dif ferential with the atmospheric pressure in accordance with ASME B&PV Code, Section Ill, Division 2, paragraph CC-5535. The channel to liner plate and related pressure barriers are tested for leak tightness in accordance with paragraph CC-5535.2 of the same Code. Indications for vacuum box and leak-chase examinations are evaluated and accepted in accordance with paragraph CC-5546 of the same Code. Material identification is in accordance with ASME B&PV Code, Section lil, Division 2, CC-2540. Control of material handling, shipping and s1orage are in accordance with CA-4450 and CA-4600.

The bottom liner is included in the Code NPT-CC stamp of the containment vessel as a "Part" in accordance with the requirements of NA-8000.

Fabrication tolerances are in accordance with CC-4522 and welding details are in accordance with CC-4542.

3.8.2.7 Testing and Inservice insoection Recuirements Corrosion protection and protective coatings for the containment are described i n Sect ion 3.8.2.2.2.

Testing and inservice inspection requirements of the containment are discussed in Section 6.2. In addition to the testing described in Sections 3.8.2.6 and 3.8.5.6, the following construction stage and preoperational tests will be performed.

3.8-7 e Amend. 68 May 1982

Pcg3 35 [8,22]#39 I

3.8.2.7.1 Bottom Liner Plate Test Before placing concrete over the bottom plate and af ter the vacuum box test of the welds, the liner welds will be covered by the leak-chase channels. The channel to liner plate and related pressure barrier shall be tested'for leaktightness by pressurizing the channels to containment design pressure (10 psi). If any ludicated loss of channel test pressure occurs within two hours as evidenced by a test gauge, the channel to liner weld will be soap bubble tested.

3.8.2.7.2 Pressure Tests A pneumatic pressure test shall be made on the Containment Vessel, airlocks, and equipment hatch at a pressure of 11.5 psig. Both inner and outer doors of the airlocks will be test ed at this prerr"re. All pneumatic tests shall meet the requirements of Appendix J to 10CFR50, NE-6000 of ASME B&PV Code Section ill, Division 1 and applicable parts of Division 2. Exceptions to Appendix J to 10CFR50 are identif ied in Section 6.2.1.4.

3.8.2.7.3 Leakage Rate Test Following successf ul completion of the pressure test, a leakage rate test at 10 psig will be performed with the airlock inner doors closed. The allowable leakage rate of the steel containment shall be 0.1% by volume of the containment in a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period and shall meet the requirements of 10CFR50 Appendix J.

3.8.2.7.4 Ooerational Testing Af ter completion of the airlocks f abrication, including all latching mech ani sms, interlocks, etc., each airlock will be given an operational test consisting of repeated operation of each door and mechanism to determine whether all parts are operating smoothly without bindin5 or other defects.

3.8.2.7.5 Leak Testing Airlocks The airlocks will be pressurized with air to 11.5 psig. All welds and seals l will be observed for visual signs of distress or noticeable leakage. The l

airlock pressure will then be reduced to 10 psig and a thick soap solution will be applied to all welds and seals and observed for bubbles or dry flaking i as Indications of leaks. All leaks and questionable areas will be repaired.

During the overpressure testing the outer door will be locked with hold-down devices if required to prevent upsetting of the seals.

The Internal pressure of the airlock will be reduced to atmospheric pressure and all leaks repaired af ter which the airlock will again be pressurized to 10 psig with air and all areas suspected or known to have leaked during the previous test be retested by above soap bubble technique. This procedure will be repeated until no leaks are discernible by this means of testing.

3.8-7f Amend. 68 May 1982

P2ge 27 L8,22Jf39 3.8.3 Concrete and Structural Steel Inthrnal Structures of Steel Containment 3.5 . 3.1 Descriotion of Internal Strucigtes The Internal structures within the containment principally consist of the cells and other ereas as listed in TableThe 3.8-2 and as Internal structuresshown on the are Generalby enclosed Arrangement figures of Section 1.2.

two continuous circular walls located on each f ace of the steel containment vessel between the foundation mat and operating floor levels. The circular walls act as a radiation shield and pressure boundary in local cell areas, as a support for vertical loads and carry horizontal shears to the foundation mat. The entire steel containment vessel will be designed for the 10 psig internal pressure. The detailed physical description provided herein is l

limited to those cells which significantly contribute to the structural system. These cells are reinforced concrete structures designed to the

. my. . . ...i s as noted i n Tabl e 3.8-2.

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l 3.8-8 Amend. 68 May 1982

- Page 30 Lt$,22JF39 Table 3.8-1b Loading Combinations for Airlocks The loading combinations for which the airlocks shall be designed are as follows:

Testing: D + Tt + Pt Normal: D+ L + To + OBE-Accident and

.. . :=:r.ta l : D + L + T' + Pi + OBE D + L + T' + Pi + SSE D + L + T' + Pe + OBE D + L + T' + Pe + SSE D + L + SSE

. 3.8-41b Amend. 68 May 1982

Page 1 (82-0314) [8,22] #56 Ouestion OCS 421.4 The applicant should f ormally submit a diagram of the auxiliary feedwater system showing the division assignments for all valves and safety grade Instrumentation and controls. A discussion should be included to indicate the normal position and position upon loss of power of each valve. In your presentation, and in Section 7.4.1.1.6, credit is taken for the feedwater Isolation valve to f all saf e in the open position upon loss of electrical power. Justify this f all-saf e analysis for all incidents (i.e., hot shorts, power supply overvoltage, etc.) that could prevent operation of this isolation valve.

Resoonse Figures 5.1-5 and 5.1-Se of the PSAR (attached) have been marked up (Figures

2 '21.4-1, 2) to show the power division assignments for all valves and saf ety grade Instrumentation. The controls for the devices are not shown, however, the power assignment is the same as that of the valves, louvers, and f an blade pitch control being actuated. The normal position and f ailure position of the valves are shown in Figures 5.1-5 and 5.1-Sa.

As shown in Figure 5.1-5 there are six Isolation valves, two in parallel to each steam drum. One valve supplies water from the electric driven pumps, the other, f rom the turbine drive pump. Each capable of supplying 100 percent water flow to the steam drum. These two valves are supplied power f rom separate Class 1E power divisions. The f ailure of one valve to open on demand would not result in the loss of water to a steam drum.

Failures such as shorts, grounds, power supply opens, etc. which result in the loss of power and thus opening the valves will not be discussed. Even if these occur during reactor power operation they have no ef feet on power operation or the ability of SGMiRS to f unction when called upon.

Each isolation valve is supplied f rom a regulated 120 VAC power supply and is designed to operate with the over voltage transients identif ied f or the power supply.

The wiring and controls for the two valves supplying water to a steam drum are in separate Class 1E power divisions and separation is provided per IEEE 383.

Hot shorts between the two valves supplying the same steam drum will not occur. The only hot short which could occur would be with the same power division and between the valve and the controls. This could result ,In two of the six isolation valves not opening when called upon. Since these two valves supply separate steam drums and there is a 100 percent redundant supply, a f ailure of two of the six isolation valves would not result in the loss of water to any of the three steam drums.

QCS421.4-1 Amend. 69 May 1982

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