ML19283B665
ML19283B665 | |
Person / Time | |
---|---|
Site: | Millstone ![]() |
Issue date: | 02/23/1979 |
From: | Counsil W NORTHEAST UTILITIES |
To: | Reid R Office of Nuclear Reactor Regulation |
References | |
TAC-46174, NUDOCS 7903060293 | |
Download: ML19283B665 (300) | |
Text
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- esti L L a :::::i;=a"O 0 February 23, 1979 Docket No. 50-336 e
Director of Nuclear Reactor Regulation Attn: Mr. R. Reid, Chief Operating Reactors Branch #4 U. S. Nuclear Regulatory Commission Washington, D. C. 20555
References:
(1) W. G. Counsil letter to R. Reid dated November 13, 1978. (2) W. G. Counsil letter to R. Reid dated July 31, 1978. Gentlemen: Millstone Nuclear Power Station, Unit No. 2 Neutron Shielding In Reference (1), Northeast Nuclear Energy Company (NNECO) provided a description of the proposed neutron shield design and discussed other relevant topics. On November 25, 1978, a meeting with members of the NRC Staff was held to present additional details of water-filled stainless steel tank design. During this meeting, NNECO indicated that the analytical qualification of this design was in progress. The Staff identified certain items which were determined to be of high interest. In response to the above, NNECO hereby provides Attachment 1, Neutron Shield Design Qualification. This Attachment is a comprehensive description and qualifi-cation of the shield, encompassing the complete spectrum of topics relevant to a neutron shield design. The information provided in Reference (1) has been in-corporated into Attachment 1 for completeness. Staff concerns raised during the meeting are addressed in the Attachment, with the exception of the items discussed below. Shield Movement in order to remove the neutron shield, it is initially moved radially outward from the reactor vessel. Due to the use of the polar crane, the most convenient path is to lif t the shield up, swing the crane 90* , and place the. shield on its storage structure located on top of the steam generator shield wall. At no time during this process does the shield pass directly over the reactor vessel area. Prior to shield movement, the water would have been drained, leaving only the dry weight of the shield structure and tanks which is approximately 12,000 lbs. total, or approximately 6,000 lbs. per lifted assembly. Since the total distance the 79030G0 AM
e shield will be moved during refueling outages is short, a postulated shield-drop incident involves a very small area within cortainment. Considering the weight of each half of the shield, the perh involved, and the absence of components important to plant safety in this exposed area, it is concluded that the con-sequences of dropping the shield do not involve a safety issue. Leakage As a matter of normal ALARA practices, personnel entering containment during power operation will be equipped with partable survey equipment such that in-creases in radiation levels would be identified. This practice provides adequate assurance that any significant leakage would be identified in a tbnely fashion, and that the appropriate correcrive action would be implemented. Note that the chield is designed such that leakage in one compartmentalized tank will not re-sult in drainage of any of the adjacent tanks. Pressure Relief Valves As indicated in Section 2 of Attachment 1, each tank assembly is provided with two (redundant) low pressure relief valves. Each one is designed to limit the internal or external pressure buildup to one psig. Because these valves are redundant, it is concluded that routine inspections provide adequate assurance that these valves will be operable. Regarding 10CFR Part 170 considerations, the determination provided in Ref erence (2) remains applicable. We trust you will find the attached information sufficient to concur with our conclusion that the stainless steel tank shield is 'an acceptable design. Should you require additional information or clarification, th.e resources of NNECO and NUSCO remain available. Very truly yours, NORTHEAST NUCLEAR ENERGY COMPANY
~, / g) / / P1 llMW W. G. Counsil Vice President Attachment
DOCKET NO. 50-336 ATTACHMENT 1 MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2 NEUTRON SHIELDING FEBRUARY, 1979
w .'. L TABLE OF CONTENTS
- 1. 0 INTRODUCTION
- 2. 0 DESIGN DESCRIPTION
- 3. 0 DESIGN BASIS CONSIDERATIONS 3.1 Structural Considerations 3.1.1 Design Criteria 3.1. 2 Structural Service Requirements
- a. Thermal Loadings
- b. Normal Operating Pcassure and Dead Loads
- c. Lifting and Handling
- d. Miscellanous Loads
- e. Seismic
- f. LOCA Loading 3.1. 3 Design Analysis
- a. Thermal Analysis
- b. Pressure and Dead Ioad Analysis
- c. Lifting and Handling Analysis
- d. Analysis for miscellanous Loadings
- e. Seismic Analysis
- f. LOCA Analysis
- 3. 2 Thermal-Hydraulic Considerations 3.2.1 Analytical Criteria and Requirments
- a. Purpose
- b. Break Location / Geometry Selection
- c. Mass and Energy Release Rates 3.2.2 Description of Analysis
- a. Reactor Cavity Modeling
- b. Major Assumptions
- c. Modeling of the Shield
- d. Sensitivity Analysis and Discussion
s - Table of Contents Continued
- 3. 3 Shielding Considerations
- 3. 3.1 Design Requirements
- a. Summary
, b. Insulation Requirements
- c. Source Terms 3.3.2 Analytical Approach
- a. Summary
- b. Streaming Regions
- c. Shield Geometry
- d. Radiation Source
- e. DOT 3. 5 3.4 Additional Considerations
- 3. 4.1 Sump Clogging 3.4.2 Fire Hazard 3.4.3 Post LOCA Hydrogen Generation 3.4.4 Containment Pressurization 3.4.5 ALARA Considerations
- 4. 0 ANALYTICAL RESULTS AND EVALUATION 4.1 Structural Results and Evaluation
- a. Thermal Results
- b. Pressure had Dead Load Results
- c. Lifting and Handling Results
- d. Miscellaneous Results
- c. Seismic Results
- f. LOCA Results
- g. Load Combinations and Evaluation
- 4. 2 Thermal-Hydraulic Results
- a. Pressure History Results
- b. Comparison Between Assumed and Computed Shield Response
- 4. 3 Shielding Analysis Results
- a. Attenuation Results
- b. Activation Results
s . Table of Contents Continued
- 4. 4 . Additional Results 4.4.1 Sump Clogging 4.4.2 Fire Hazard 4.4.3 Post LOCA Hydrogen Generation 4.4.4 Containment Pressurization 4.4.5 Occupational Exposure
- 5. 0 CONCLUSIONS G. 0 REFERENCES APPENDICIES A. Structural Analysis Calculations B. Nodalization Sensitivity Study C. Loss Coefficient Calculation and Friction Sensitivity Studies D. Comparison of Results with Air-Water-Steam Model E. Shield Dynamics Equations
- 1. 0 INTRODUCTION During the startup test program at Calvert Cliffs Unit No. I, higher than anticipated neutron levels were observed within the reactor containment building. Tests conducted during the Millstone Unit No. 2 startup test program confirmed that a similar condition existed at Millstone.
This condition has resulted in limitations on L .tainment access during reactor operations in order to minimize personnel radiation exposure. Since the neutron streaming phenomenon was first identified, Northeast Nuclear Energy Company (NNECO) has been evaluating alternative shieldinr, concepts to ensure that the optimum design was installed. In response to paragraph 2E of DPR-65, NNECO proposes to install a neutron shield above the reactor annulus to reduce neutron radiation levels within containment. The neutron shield design and performance during both normal plant operations and accident conditions , are discussed herein.
- 2. 0 DESIGN DESCRIPTION The proposed reactor cavity neutron shield for Millstone Unit No. 2 consists of two (2) semi-circular compartmented water tanks. When joined together the tanks form an annular ring which covers time reactor cavity from near the vessel head to the refueling cavity floor. Each of the tanks le subdivided into eight (8) subcompartments as shown in Figure 2.1 and contain 16 inches of unborated water. Communication between subcompartmcnts is provided at the top of each radial plate. This construction allows for ease of filling and precludes the loss of the entire water inventory should any subcompartment de-velop a leak.
The tanks are constructed entirely of stainless steel. The sides of the tanks and of each subcompartment are all one-half inch thick. The tanks are scaled by means of one-eighth inch thick top and bottom cover plates (see Figure 2. 2) which are welded to the radial and circumferential tank members. This prevents water loss due to evaporation and provides additional structurn1 integrity and stiffness for handling purposes. An air gap is provided between the surface of the water and the top cover plate. 'I'ais allows the water to expand inside the tanks and prevents over flow due to normal operational heatup. In order to minimize the internal and external pressures which occur as a result of thermal cycling, each tank assembly is provided with low pressure relief valves. These are designed to limit the internal and external pressure buildup to 1 psig. Redundant valves are placed in each tank assembly to provide press-ure relief chould any one valve fail.
The temperature of the water in the tanks is maintained below 212 F by using a combination of thermal insulation and forced convective cooling as shown in Figure 2. 2. The convective cooling air is the exit air from the cavity insulation provided by the IIVAC system. By forcing this air along the bottom of the shield tanks and up the inner diameter, conductive heat losses from the insulation are removed before they can provide a large input to the tanks. Rad-lational heat transmission is minimized by polishing the sides of the inslation facing the reactor pressure vessle. The two semi-circular tmik segments when placed in position within the reactor cavity are joined together by means of four (4) hinge pins. This arrangement is shown in Figure 2.3. During the installation processes the "C ring" on the shield i engages the reactor seal flange (Figure 2.2). Once the hinge pins are installed the entire assembly has hoop continuity and cannot become disengaged from the reactor flange. This prevents the shield structure from becoming a missile following a LOCA. During the postulated LOCA event in the reactor cavity the bottom and top cover plates of the tank subcompartments rupture to provide the necessary cavity pressure relief. The rupture is controlled by means of a series of grooves cut into the cover plates as shown in Figure 2.4. After the grooves begin to tear, the four sections of the panels so formed. collapse against the sides of the tank subcompart-ments to provide the necessary pressure relief area. The top rupture panel will burst somewhat after the bottom panel due to the air gap above the water. This is desirable since bursting the panels sequentially vi11 tend to min-inize the uplif, forces on the remainder of the structure. During the LOCA event, the insulation panels located below the shield tanks will be ejected upward through the opening formed by the burst panels. To insure that the insulation panels will be ejected, a non-structual radial joint exists at the center of each insulation panel. In effect, this results in two panels per subcompartment which ensure that the insulation panels are thrown clear of the tank structure.
a non structual radial joint exists at the center of each insulation p:mel. In effect, this results in two panels per subcompartment which ensure that the insulation panels are thrown clear of the tank structure.
- 3. 0 DESIGN BASIS CONSIDERATIONS The sections below discuss the structural, thermal-hydraulic, shielding, and additional requirements which were imposed to demonstrate that the shield structure will perform its intended function. Each requirement is 6efined along with a discussion of the analytical approach, justification, and assumptions.
3.1 Structural Considerations 3.1.1 Design Criteria The structural design criteria for shield tanks is based upon the rules contained in reference G.1 for steel structures inside containment. The intent of the reference 6.1 rules are:
- 1. to insure the integrity of safety related structures which are designed to prcvent a LOCA.
- 2. in the unlikely event that a LOCA occurs, the structures vill mitigate the consequences of the accident by protecting against failure of the engineered safety features.
The cavity shield falls mainly into the second category due to the following two considerations. First, failure to maintain post LOCA structural integrity could produce missiles which could damage other safety related equipment; and sewnd, failure of the shield to provide the necessary venting of the reactor cavity during LOCA could damage the primary shield wall structure. Because of these consid-erations the shi. Id structure is treated as safety related. The applicable stress criteria for evaluation purposes whloh is employed is that in Reference 6. 2. . Stress limits for the material used in the shield (i.e. stainless steel) are not specifically given in this reference. However, the criteria used in Reference 6. 2 for ductile steel structures based on yield and elastic analysis is employed to give the appropriate stress limits for the shield. This results in a conservative evaluation of the shield for normal operation (1.e. all loadings ex-cept LOCA) and the applicable loading combinations defined in Reference G. L 4
For evaluation of the severe, extreme, and abnormal lands ns defined in Reference 6.2, plastic analysis methods are usei as given in part 2 of Reference G.2. The resulting limits are evaluated against the stresses resulting from the load combinations given in Reference 6.1 for steel structures. 3.1.2 Structural Service Requirements The loadings and requirements which affect the integrity and per-formance of the cavity shield structure are discussed in sections a) through f) below. For purposes of analysis and evaluation the load-Ings can be divided into two (2) categories - namely normal operat-ing loads and seismic plus LOCA loads. Sections a) through d) deal with the normal load category and include: a) Thermal loadings b) Operating pressure loading c) Lifting and handling loads d) Miscellaneous loadings This division of loadings is noted in anticipation of the evaulation proceedure given in Reference 6.1. 3.1. 2. a Thermal Loadings The design of the shield tank and accompanying insulation must address three (3) different thermally induced loading conditions. First, since the shield tank is designed to burst at relatively low pressures and cannot, therefore, act as a pressure vessel, the average temperature of the water in the tanks must be kept below 212 F. The insulation system must protect the tank from the approximately 550 F operating temperatures of the reactor pres-sure vessel and the air in the reactor cavity (Figure 3.1.1). In addition, the energy released into the shield water from the absorb-tion of neutrons must be considered when calculating the steady-state water temperature.
The second thermal condition which must be addressed is the potential for buildup of high discontinuity in the tank due to high thermal gradients. The only structural member of the entire assembly which might possibly see high temperature gradients is the region where the C-clamp surrounding the reactor pressure vessel flange joins the rest of the shield tank structure (Figure 3.1.1). The C-clamp can be expected to reach a temperature of 550 F while the tank can conserva-tively (for this analysis) be set to 160 F. The last loading condition associated with thermally induced loads concerns the ability of the tank to respond to thermal growth in the reactor pressure vessel from shutdown (cold) to normal operating (hot) temperatures. This simply means that the C-clamp surrounding the reactor vessel flange must provide enough clearance to accomodate both outward and upward movements of the flange (Figure 3.1.1). 3.1.2.b Normal Operating Pressure and Dead Loads Under normal operating conditions, the shield tank will be subjected to its own weight plus hydrostatic pressure from 1G inches of water. In add' tion, positive internal pressure would be produced during start-ups when water in the tank is heated from some nominal temperature (50 F) to some level below 212 F. Conversely, during shutdown periods cooling of the air and water in the tank could produce a nega-tive internal pressure. Assuming that both the water and air in the tank cycle from 50 F to 212 F, the change in pressure would be 14.9 psi. To prevent fatigue damage from cyclic loadings of burst panels due to changes in internal pressure, relief check valves will be installed. These valves will insure that the internal pressure of the tank is nes er more than 1 psi above or below the ambient atmospheric pressure (Figure 3.1. 2).
3 1.2.c Lifting,and Handling In order to fill the suppression pool for refueling or other maintenance work, both halves of the shield tank will be lifted out of the way and replaced by the seal ring. Under normal conditions , the tank vill be drained of water before it is lifted out of place. Certain maintenance activities may, however, require that the tank be lifted without being drained. The tank vill have sufficient strength to resist any loads imposed during lifting, even when completely filled with water. The loading consists of the dead weight of each half tank and water multiplied by a factor of two to account for dynamic loads. Each half of the tank is supported by three lifting lugs atta, ed to cables. 3.1.2.d Miscellaneous Loads This loading category refers to the loads which would be encountered during a refueling outage besides those associated with lifting and handling as discussed above. The only loading identified in this category is workmen standing or equipment being placed on top of the shield (i.e. , on the upper burst panel). It is not anticipated that any equipment would be placed on the shield other than hand tools or other items which could be carried by ore or two men. As an upper bound load, it was , therefore , assumed that two workmen at 225 lb. each stand on a single subcompartment and that the entire load is concentrated at the center of the upper burst panel. 3.1 2.e seismic Loads During a seismic event motions vill be imparted to the shield structure through the dead weight support at the pool floor. Depending on the magnitude of the frictional force which exists between the tank support and the pool floor and the level of the horizontal floor accelerations, the tanks will either follow the floor motions or slide on the floor until the floor acceleration subsides or the reactor flange is contact-ed. In both instances, loads will be included in the tank structure.
In addition to inertial loadings resulting from the seicmic motions or interaction with the reactor flange, sloshing of the water in the tank subcompartments will occur and be mainfested an an asymmetric internal tank pressure. The vertical seismic input accelerations are considered only for their effect on the horizontal tank motions (i.e., relieving the deadweight load which results in loss of a resisting frictional force). The vertical G-levels for both SSE and OBE are less than 1.3 which means the shield will never lift off its support. The vertical loading which controls the design is the pressire loading due to the LOCA as discussed in the next section. For reference the horizontal and vertical response spectrums for the pool floor are shown in Figure 3.1.3 and 3.1.4. In order to conservatively quantify the seismic loadings in terms of the available inputs and structural configuration, the following observations and assumptions are made. First, if the tanks slide, the maximum differential acceleration between the tanks and the floor will be the floor zero period neceleraction (in g's) minus the dynamic friction coefficient between the tank supports and the floor. If this condition is assumed to exist in time long enough to deflect the shield through the maximum reactor flange to "C ring" diametral clearance, thcn an upper bound impact with the reactor flange will occur. This also results in an upper bound sloshing of the internal water.
. ~ 3.1.2.f LOCA Loadings Due to the postulated nipture of the main coolant lines at the reactor vessle nozzles, the shield structure is subjected to the resulting hydraulle pressure forces which exist at the cavity exit plane. The structural requirements on the shield during this event can be categorized into 3 groups as given below:
- 1. The shield structure must be strong enough to preclude any portions of the stainless tanks from becoming missiles.
IDCA loadings are transferred from the tanks down into the "C ring" which is retained by the RPV seal flange. In this regard both the "C ring" and vessle flange must be shown to be adequate. In addition, the tank rupture panels must be shown to be retained after bursting. To show this analytically is impractical. Rather than analyzing to show that the panel to tank weld will remain intact samples of the panels and panel to tank welds will be deformed through large angles to show that sufficient ductility, exists to preclude fracture. The use of a ductile material such as stainless steel is ex-pected to allow for more than the required deformation.
- 2. The tank rupture panels must open early and fully enough after the LOCA under the pressure loading to preclude the cavity frcm becoming over pressurized. The rupturing event is a dynamic process and must be treated as such in order to correctly predict the cavity conditions and struct-ural response. Since the dynamic response of the rupture panels is a function of the 7put pressure history and the pressure history depends on the characteristics of the rupture panels the exact determination of the panel response becomes an iteractive process within the hyhaulics analysis. This pro-cess is discussed further in Section 3. 2 which deals with the thermal-hydraulics analysis of the LOCA event.
. ~
- 3. In addition to qualification of the burst panels, it must also be shown that the insulation segments contained below the shield are ejected and do not clog the aperture created by the burst par ^1s. If an insulation panel vere to lodge in the tank and not be ejected, the added resis-tance to the escaping flow would tend to increase the cavity pressure loadings.
During the LOCA event, the maximum loading on the shield structure occurs just at the point when the bottom panel begins to burst. The effective pressure on this panel is composed on the cavity pressure and the inertial pressure exerted by the water which tends to resist rapid motions of the panel. This means that the cavity pressure which exists at the time the panel bursts is going to be snmewhat higher than the panel static burst pressure. However, the lo ul vaich is transferred back into the remainder of the tank structure will be equal to the load due to the static burst pressure. The reaso.. for this phenomenon is that the resist-ing inertial loading of the water subtracts from the applied cavity pressure and reduces the effective load entering the shield tanks. Once the panels begin to tear open, the affected area (and hence the load) reduces rapidly. The insulation segnents are non-structural items made from very light gage (approximately 20 gage) steinless steel sheet. Since a conservative analysis demonstrates that the burst panels completely fold up against the sides of the tank, it is virtually impossible for the insulation panels to hang up in the vent openings. Clear-ing is assured because of the radial joint in the center of the segnents as shown in Figure 3.1.5 This results in 2 insulation panels per vent opening which are in effect uncoupled from each other. The only mechanism which could preclude clearing of the insulation panel involves migration of the panel circumferentially
around the shield to beneath one of the radial gussets prior to rising. However, this would merely result in the insulation collapsing against the gusset plate in a similar fashion to the burst panels which are nearly four times thicker than the material the insulacion panels are formed from. 3.1.3a Thermal Analysis In order to detennine the steady-state shield tank water tem-perature, a heat transfer analysis is required. The shield tank is in a complex environment involving convection, conduction, and radiation occurring simultaneously in several different mediums. The most straight forward solution approach involves the use of a computerized numerical finite element heat transfer program. This problem was modeled and solved using the ANSYS computer program. in the heat transfer analysis mode. A model of the shield tanks and insulation was developed using 2-dimensional axisymmetric heat transfer elements. The model includes bottom and side insulation panels, the cool air stream between the insulation and the tank walls, the tank walls and top and bottom plates, and the bodies of water and air contained within the tanks (Figure 3.1.1). The outer nodes of insulation elements were conservatively set at 550*F. The top cover plates and outer wall are in contact with ambient air at 160 F. The air pumped through the space between the insulation and tank walls is assumed to be delivered at a rate of 5460 scfm and has an entry temperature of (conservatively) 170*F even though plant specifications call for 160*F air.
The insulation, tank plates, water and air masses within the shield tank were all modeled with isoparametric quadrilateral temperature elements (STIF55) having appropriate conductance properties. The moving stream of air was modeled with fluid flow- heat transfer elements (STiF5G) using a known rate of flow. Radiation link elements ('STIF 31) and convection link elements (STIF 34) were connected from the insulation to opposing tank plates. The air contained within the tank was also given convection ability. Convection link elements were not provided within the water mass. This is a very con-servative assumption because the water is undoubtedly capable of transfering large amounts of heat energy through it by means of natural convection. Figure 3.1.G is a coniputer generated plot of the model showing the element grid pattern. The rate of energy input due to neutron absorbtion can be determined since the number and energy of incoming neutrons is known and the effects of elastic scattering and gamma ray production can be calculated. The amount of energy available for heat production is 1. 21 Btu /hr/ tank. Since the water will circulate naturally, the heat must be dissipated into the air layer on top of the water. The film coefficient required to pass this amount of heat from water into air assuming a 10 difference in temperature is 0.00925 Btu /(hr-ft - F). Since the actual coefficient has been found to be 0.474 Btu /hr-ft - F), the heat input due to neutron absorbtion can easily be dissipated and can essentially be ignored. The second thermal probletp, that of calculating stresses due to unecaal thermal expansion in regions of high temperature gradients, was also solved using the finite element method. The ANSYS computer program was employed using a static analysis of structural elements having a specified set of nodal temperatures. The model (Figure 3.1.7 ) covered 90 of the tank with symmetry boundary conditions imposed at both ends. The flange C-clamps, inner,. outer; and side walls and top and bottom cover plates were all modeled with elastic flat quadrilateral shell elements (STIFG3).
The model was broken up into 8 elements around the circumference and 4 in the radial direction. All nodes on the C-clamps were set to 550 F while the rest of the structure was given a temperature of 160 . The last thermal problem, that of clearance between the reactor vessel flange and the C-clamps, was solved by a simple hand calculation. The maximum outward and upward growth of the flange was calculated, assuming unrestrained thermal growth of the reactor pressure vessel over a range of 70 F to 550 F. 3.1. C'> Pressure and Dead Load Analysis The dead weight analysis was carried out using the same model(Figure 3.1.7 ') and computer program (ANSYS) as described in section 3.1. 2a. Thermal loading was removed and replaced with gravity loads based on the actual weight of all materials. The tank is supported by a plate welded to the outer circumferential wall. The nodes along the base of this plate, resting on the platform, were all constrained from movement in the vertical direction. IIydrostatic pressure loads due to the weight of water contained in the tank as well as the maximum internal pressure of 1 psig were defined over the bottom burst panel and varied linearly up the walls of the tank to a height of 16 inches. 3.1.3c Lifting and IIandling Analysis The same basic model (Figure 3.1.7 ) was used for this analysis, with minor alterations. Spar elements (STIF8) were added to two (2) points on the model and brought together over the centroid, simulating lifting cables (Figure 3.1. 3 ). Symmetry conditions were applied along one side of the tank. The second side, which unhooks from the other half tank for lifting, was freed from all constraints. The model was analyzed for dead weight alone and dead weight plus water weight.
This model was analyzed with linear elastic, small deformation, small strain theory. This means that the structure as modeled would be un-stable unless the point where the two lifting cables joined was over the exact centroid of the tank half. Since the centroid location can not be calculated to sufficient accuracy, a few weak springs were added to the model to prevent large rigid body displacements. The springs will have no measurable effect on the state of stress within the tank as long as the point of lift is close to the centroid and the resulting unbalanced, moment remains small. 3.1. 3. d Analysis for Miscellaneous Loadings The loading discussed in Section 3,1.2 d) was evaluated by considering each of the four segments of the burst panel to act as a cantilever from the sides of the tank. In addition, it was assumed that the material in the grooves pro-vided no stiffening to the remainder of the panel. The load was conservatively applied at the apex of the most critical panel segment and it was assumed that one fourth the total load of 450 pounds was concentrated there. This is shown schematically in Figure 3.1. 9. 3.1.3.c Scismic Analysis In order to bound the seismic loadings due to the OBE and SSE loadings as discussed in Section 3.1. 2. c) the following two analyses were performed.
- 1. The entire tank assembly was allowed to slide through the maximum reactor flange to "C ring" diametral clearance (2.0 inches) assuming the maximum zero period acceleration from the horizontal resInnse spectrum (see Figure 3.1. 3) act in one direction for the entire time it takes the shield to move through the clearance distance. The zero period acceleration used in the analysis was adjusted downward to in-clude the resisting effect of friction. The friction was conservatively
lower bounded by assuraing a lower bound coefficient of .20 and reducing the weight of the shield in the vertical direction by the peak G-hvel of the vertical response spectrum (see Figure 3.1.4). The resulting SSC and OBE accelerations of .364 G's and . 082 G's give respective impact loads of 119,391 lbs. and 56,GG7 lbs which were used in the clastic impact analysis with the reactor flange. The reactor flange was assumed to be rigid. The "C ring"ftank assembly stiffness was evaluated using "ae model shown in Figure 3.1.7 but including 180 instead of only the 90 segment shown. Figure 3.1.10 shows a plan - ew of the model. A mode-frequency analysis was run on this model assuming the impact occurs at a single point. lO.owing the total mass of the shield and its characteristic frequ(ncy when supported as shown allows the effective spring rate to be computed. From this information, the peak impact force can easily be computed.
- 2. The peak force computed in 1) above was used to determine the maximum acceleration of the water. This results in an equivalent internal pressure force which acts over various portions of the subcompartments. The max-imum average static pressure is 5.3 psi. This pressure is far below the level necessary to produce significant stresses in the i inch portions of the subcompartments but is significant when applied to the burst panels. As a result this pressure force w..s conservatively applied over an entire burst panel in order to bound the stresses for the seismic condition.
3.1. 3. f) LOCA Analysis The analytical effort required to demonstrate the acceptability of the shield during LOCA can be divided into two (2) categories:
- 1. Analysis of the burst panels
- 2. Analysis of the remaining tank structure Item 1 addresses not only the burst behavior of the panels (1.c. at what pressure the panels actually burst and their dynamic behavior) but also their post burst behulor to demonstrate that the panels completely opet to provide full venting and open su'.ticiently to allow ejection of the thermal insulation. The dynamic behavior of the burst panels establishes the validity of the thermal hydraulic analytical assumptions as discussed in Section 3.2. Item 2, analyzes the behav*.or of the t'nks under the action of the loadings generated as a result of the burst panei analysis.
The item 1 analysis was performed on an idealized model the burst panel as shown in Figure 3.1.11. In this model the panels were treated as rigid for out of plane bending but flexible in the plane of the plate. This ideal-ization along with some further assumptions discussed below allows the panel to be modelled as single degree of freedom system subject to a set of time and displacement dependent forces. Ignoring the plate bending stiffness while retalning the plate in-plane flexibility tends to upper bound the burst pressure. The reason for this is shown in Figure 3.1.12. Adding bending flexibility tends to force large groove strains and, hence, sooner bursting. From a dynamic analysis standpoint we want to upper bound the time at which the panels burst since this will in turn provide the most conservative comparison back to the original assumptions built into the thermal-hydraulics analysis (see Section 3.2). In addition to the above,the panels were assumed to be supported by the plastic hinge which forms at the panel to tank wall junction.
The section of panel analyzed by the above method was selected on the basis of static analysis of the above described model to determine the section of panel which best approximates the burst pressure. In analyzing this panel dynamically it is assumed that the differential vertical motion between adjacent panel sections (1. e. across the grooves) is zero. This is conservative in that it lower bounds the groove strain as the panels rise. In addition, the shear deformation which occurs in the grooves as a result of the differing radil to the various hinge lines was totally ignored. This lower bounds the rate of groove strain accumulation and hence upper bounds the grooves restoring force. The behavior of the plates in the area of the grooves was determined by a series of tensile tests as documented in Reference 6.3. This data was used to generate a stress strain curve for use in the time history analysis. The resulting curve is shown in Figure 3.1.13. This curve was piecewise linearized into the segments shown for use in the analysis. Details of the data reduction to obtain this curve are given in Appendix A.
The intont of the time history burst panel analysis is to demonstrate that the panels begin to open shortly after rupture and that once the material in the grooves begin to fracture,the processes rapidly continuee until the panels reach their full open position. In order for the thermal-hydraulics analysis (Section 3.2) to be conservative as performed, the bursting process must begin well before the pressures in the cavity begin to peak, other wise the peak pressures as computed cannot be considered valid. This consideration is most critical in the regions near the break since the shortest pressure rise times are found here. Since the inertial effect of the w ^er tends to resist rapid motions of the burst panel and, hence, lowers the effective pressure on the burst panel at any instant of time, the largest deviation from static response vill occur in those regions where the pressure builds most quickly. In order to model the inertial effect of the water the effective panel density was uniformly increased. This modeling approach ignores the fact that as the panel rises the water above it can displace horizontally as well as vertically and, therefore, the inertial water resistance is some what over predicted. This will lead to an upper bound solution for the pressure at burst. The panels were assumed to act elastically during the LOCA event except in the groove region and at the junction with the tank sides where plastic hinge elements were used. Ignoring plasticity in the plates is con-servative in terms of predictirgthe burst pressure since this assumption tends to over estimate the stiffness of the panels. This, in turn, reduces the amount of deformation in the groove region which delays the time at which bursting initiates. In addition, the panels were assumed to act rigidly in bending. This further limits the net strain in the grooves which in turn tends to overpredict the burst initiation as well as the post-burst behavior.
f The figure 3.1.11 model was used to predict the static burst pressures as well as to form the dynamic analysis model of the panel. In either case the only restoring forces come from the stresses in the groove as well as the plastic hinge of the panel itself. These two restoring forces and the inertia of the panel balance the pressure forces. Once the material strains had reached 100% in the groove that portion of mater:al so st'minedwas taken to be fractured and removed from contributing to the restoring force. The strain of 100% was determined to be the strain at which the material in the groove fails based upon the Reference 6.3 test data as manipulated in Appendix A. Accounting for this allowed the model to be run out in time to the point at which the panels fully open. The remainder of the tank structure was analyzed by applying the peak uplift force on the burst panel just prior to bursting over the gussets and inner and outer cylinders. The peak force is the static burst pressure times the panel area. Once the panels have burst the load drops rapidly as is shown in Figure 4.1.14 and discussed in Section 4. le. This is also the effective pressure at burst during the LOCA event since the inertial forces create an opposing pressure which reaches the applied panel pressure to the static burst value. As a result, the load transferred back into the remaining shield structure is only that due to the 20 psi static burst pressure. Applying the load to the tank structure results in the load configuration shown in Figure 3.1.14. This figure shows the finite element idealization along with the applied loading. This model was run on ANSYS using the three dimensional plate element (ANSYS element 63).
- 3. 2 Thermal-Hydarulle Considerations The thermal-hydraulic considerations needed to establish the conditions in the reactor cavity which affect the shield are discussed in the sections which follow. The analytical
requirements and critera are covered in Section 3.2.1 followed by a detailed descriptian of the analysis in Section 3.2.2. The results of the hy 'raulics analysis are covered in Section 4.2. 3.2.1 Analytical Criteria and Requirements
- 3. 2.1. a Purpose The purpose of the LOCA taermal hydraulics analysis of the reactor cavity is to provide a set of loadings on the neutron shield structure which are bounding for purposes of structural analysis and design.
In order to accomplish this goal several key considerations were made in the initial stages of the analysis in order to establish the requirements. These are:
- a. Determine the tressure loading condition which upper bounds the shield's performance.
- b. Detemine the break geometry and orientation which gives the worst case loadings.
- c. Detemine the associated mass-energy release for the selected break.
From the standpoint of the shield's structural perfomance' (item a) above), it has been detemined that the worst case load application is one in which the pre:sure loading is highly asymmetric. The reason for this is not the effect an asynmetric load has on the shield structure itself but rather the effect of maximizing the delay in providing the necessary vent area and resultant loadings generated on the reactor vessel and the primary shield vall. If it can be shown that the shield vill provide adequate cavity venting for the worst asymmetric pressure distr'.bution, then it follows that the more symmetric pressure dir a ibutions resulting from other breaks will also be ventM adequately. The basis for this conclusion is related to the pressure rise times. A faster rise time will result in the rupture panels bursting at a correspondingly higher pressure due to the resisting inertial effect of the water (see Section 3.1.3f).This
would tend to have the largest infitence on the resulting peak pressures. It should be noted that the above statements do not imply that the break considered in this analysis necessarily gives the highest asymmetric reactor and shield wallloads but rather that the presence of the shield will be most apparent for these breaks characterized by the most rapid pressure increases. A discussion of the logic involved in the break selection and the origin of the corresponding mass and energy release rates as given in the two sections which follow.
. 4 3.2.lb Break Location / Geometry Selection As stated in FSAR Section 5.2.5.h.L a 5.89 ft2 guillotine breac. is the limiting break regarding reactor cavity pressurization. This break corresponds to a 0.6 double ended guillotine (DEG) break. As documented in this FSAR section, a full DEG break in the reactor cavity is not possible because of restraint provided by the primary shield vallc. A full double ended guillotine break was assumed in the neutron shield design for added conservation. A one millisecond break opening time was assumed. The results of this analysis conservatively bounds any break which could possibly occur in the reactor cavity. Hot leg breaks were not analyzed because hot leg break areas in ;he reactor cavity are extremely limited due to restraint provided by the primary shield valls. As seen in this FSAR section, peak differen+ al pressure caused by a hot leg break is an order of magnitt.de less their peak differential pressure caused by a cold leg break. Cold leg breaks remain limiting and, consequently, hot leg breaks need not be considered further. 3.2.lc Mass and Fnerey Release Rates The mass and energy release data which were used in the neutron shield analysis are those corresponding to a 1.0 DEG cold leg break. These release rates were used in the original containment analysis and are presented in the response to Question 6.20.2, the FSAR, Amendment 16.
. J 3.2.2 Descriptf on of Analysis 3.2.2a Reactor Cavity Modeling The pressure histories for the reactor cavity were developed by the use of the RELAP4/ MOD 5 computer program (Reference 6.3), This program was used because of its capabilities for the variable vent sizes required for the expected inclusion of neutron shield plug modeling the program usage and assumptions belov vere the subject of appropriate sensititity studies to ccnfirm the validity of this analysis. The containment subcompartments were div'ded into four basic regfons - the East Steam Generator Room, the West Steam Generator Room, the Containment, and the Hr--tor Cavity as shown in Figure
- 3. 2.1. The Reactor Cavity itself was divided into thirty volumes as shown in nodalization scheme for the unwrapped reretor cavity of Figure 3. 2. 2.
The nodalization within the reactor cavity was established by defining flow paths at each definable geometric restriction which might affect the flow patterns and pressures resulting from the LOCA. This resulted in 30 nodes for the reactor cavity and is called the 33 node model. The adequacy of this model was tested for sensitivity to nodalization as dis-cussed in Appendix B. As a conclusion of the nodalization sensitivity study of Appendix B and previous study (discussed in Reference 6. 5,6.6 and 6.7, the 33 node model was found sufficiently insensitive to nodaliza-tion. The flow path (junction) or vent area joining two interconnecting free vohnnes was chosen to be the minimum flow area along the flow path to provide a realletic and conservative application of the vent losses (between nodal volumes) and junction critical velocities. The methods for calculating vent loss coefficients is consistent with the requirements of Reference G. 8.
Table 3 is a list of the RELAP4/ MOD 5 Volume data for the 33 node model, and Table 4 is a list of the RELAP4/ MOD 5 junction data. 3.2.2b Major Assumptions TPe major conservative assumptions of the analysis are:
- 1. 'leactor cavity insulation remains in place during the blowdown.
- 2. Initial atmosperic conditions are saturated steam at atmosperic pressure.
- 3. Vent flow behavior is based upon a homogeneous mixture in thermal equilibrium with 100% water entrainment.
- 4. Vent critical flow correlations are " frictionless Moody" with a multiplier of 0.6.
- 5. RELAP4/ MOD 5 is an adequate simulation model when used with the applicable NRC guidelines.
Assumption 1, pertaining to the lack of insulation collapse, is conservative because it minimizes subcompartment volumes and, therefore, subjects them to higher pressures. The use of saturated steam as the initial atmospheric fluid is considered conservative because it maximizes the amount of vapor in the reactor cavity and, therefore, maximizes the local pressures. This assumption was shown to be reasonable with the multi-component multi-phase simula-tion capability of the RELAP4/ MODS program and is discussed in Appendix D. Assumptions 3 and 4 are consistent with the applicable NRC criteria (Reference G. 8). Their use is discussed in Section 3.3.3a.
The applicability of RELAP4/ MOD 5 is a necessary assumption because this program has not been specifically approved for subcompartment analysis (Reference 6. 8). The applicability of REImP4/ MOD 5 as used in this analysis is, therefore, justified by the following:
- 1. Program input variables, including vent flow paths, choking models, and initial conditions, are calculated based on the NRC guidelines, SRP 6. 2.1. 2 (Reference G. 8)
- 2. The flow equation option of incompressible single steam flow (form 3) is the same as that used in the NRC approved code, RELAP3.
- 3. The RELAP4/ MOD 3 program, predecessor to RELAP4/ MOD 5, was found to provide reasonable results, comparable to the NRC approved RELAP3 program (Reference 6.0).
- 4. The RELAP4/ MODS program with air handling options employed was used to benchmark the CONTEMPT 4/ MOD 2 program (Reference 6.10). A comparison of RELAP4/ MODS with air and steam ' low is discussed in Assumption 2 and Appendix D.
3.2.2c Modeling of the Shield In order to determine the LOCA reactor cavity pressures with the proposed neutron shield and toprovide inputs for the design and analysis of the shield tank structure, additional analysis was per-formed using the 33 node model discussed eariler. The time varying vent sizes required formodeling the tanks was one of the major reasons for employing the RELAP4/ MOD 5 program for this analysis. Since the hydraulics analysis dicusssed here as well as the response analysis of the burst panels discussed earlier are both nonlinear time history analyses, it is not possible to directly compare the two in order to obtain a once through final solution. An iteractive -scheme is req"tred in which the initial behavior-of the shield would be estimated, pressure-time results generated,
the shield burst panel response determined, the hydraulics analysis modified, and the entire process repeated until a convergent solution is obtained. In lieu of performing multiple interactions in this manner, the solution processes could be stopped at whatever point it can be dem-onstrated that the results are conservative. The latter approach was adopted in order to minimize the quantity of analysis required to demonstrate acceptability. The key feature of successfully utilizing the latter approach is the method used in modeling the shield. If the shield resistance is modeled to be too large, unacceptably high pressures and loadings may result. Too low a resistance, on the other hand, would require further analysis iterations in order to demonstrate conservatism. Fundamentally, the bursting of the shield tanks can be thought of as lifting a weight (i.e. the water weight) in parallel with deforming a spring (i.e. the panel). In order to incorporate the shield tank characteristics into the hydraul-ics model in a manageable fashion, the following basic assumptions were made:
- 1. The water is lifted as a single rigid mass
- 2. The resistance of the lower burst panel can be characterized as a linear spring in parallel with the water mass up to the burst point. After this the resistance force is taken to be constant.
- 3. The resistance of the upper burst panel is identicel to the lower panel and is in parallel with the lower panel spring,but is in turn in series with the spring characterized by the air gap between the top surface of the water and thelower surface of the upper burst panel.
- 4. The opening area is assumed to be a linear function of the vertical kinematic displacement of an equivalent rigid body mass used in the model.
The resulting model is shown in Figure 3. 2. 3. The loadings on the shield tanks will be upper bounded by upper bounding the net resistances give above. Assumption 1 %;noresany dissipative effects which would be present in the water. These losses, while small, . ensure that this assumption is conservative. Assumption 2 is conservative since the computed spring rate was based on the panels behaving as rigid plates supported by a pure plastic hinge at the wall as well as by the groove tensile forces just at the point of impending burst. After this point the resisting force of the panel was taken to be constant (1.e. assuming the panel followed the linear load / deflection curve implied by the spring rate is overly conservative). Assumption 3) considers the upper panel spring rate to be identical to the lower panel. IIowever, in series with the upper panel stiffness is the stiffness associated with the air space between the water and the upper panel. This air space is important during a LOCA event since it causes the burst panels to rupture in a cascade . .shion which tends to lower the overall resistance to the escaping flow. In addition, because the panels rupture successively, only one burst panel need be analyzed for its time history response. This is because the soft air spring tends to decouple the two panels. Assumption 4 deals with the opening function. Since it is not known apriori what the panel response and burst behavior is, it is assumed that the vent opening is a linear function of the height of the equivalent resisting mass used in the hydraulica analysis. This assumption is justified on the basis of the results as discussed in Section 4. lf.
/
The equivalent resisting mass actually used in the hydraulics analysis vas computed based on equating its kinetic energy with the energy associated witn the Figure 3.2.3 model at the full vent height. In order to complete the calculation, a velocity was conservatively selected for the equivalent mass. Based on past experience and judgement, an initial value of 120 ft/see was selected. This yielded an equivalent resisting mass of 6.8 lb-sec2/in or 2625 lb. for each subcompartment. The velocity value used is also confirmed in Section h.lf. Even with the above simplifications the thermal hydraulic results remain coupled with the behavior of the equivalent mass discussed above. In order to couple the fluid dynamic behavior .ath the behavior of the equivalent mass, the following iteractive procedure was used.
- 1. A reactor cavity to containment area opening curve based on the behavior of the shield plug displacement was assumed.
- 2. The area time curve from (1) was input into the RELAP4/ MOD 5 model as a " leak" table. The model was then run to establish a set of pressure and flow rate data.
- 3. Data from (2) was then used to calculate another set (next iteration) of area-time curves.
- 4. If the set of area-time curves calculated in (3) was different from that assumed in (1), procedures (1) through (3) were repeated until the values converged to acceptable resuus.
This iterative procedure, uc. 'g the variable vent area spechication capabilities of RELAP4/ MOD 5, was found to converge within 2 iterations with less than 2% variation in pressure and less than 5% variation in plug displacement. l t
3.2.2d Sensitivity Analysis and biscussion Four sensitivity analyses were performed during the course of this analysis. The sensitivity studies consist of :
- 1. Time step sensitivity analysis
- 2. Friction factor sensitivity analysis
- 3. Nodalization sensitivity
- 4. Initial compartment fluid conditior.s comparison The purpose of these studies was to test the sensitivity of the various paraireiers that might effect the pressure results and to choose the values of these parameters such that the pressure distribution is not unduly sensitive to these parameters.
Time Step Sensitivity Analysis The purpose of the time step sensitivity analysis was to ensure the insensitivity of the pressure results to the time step sizes chosen. A trail-and-error type of solution was performed to arrive at a set of optimum time step sizes. The set of optimum time step sizes was then checked for sensitivity to the pressure results by comparing the results with resulte obtained using larger and smaller time step sizes. The following set of time step sizes was first selected: RANGE DTAIAX DTA11N 0-0. 01 sec 0.0001 sec 0. 00002 sec
- 0. 01-0. 02 sec 0. 0005 sec 0. 00005 sec
- 0. 02-0. 04 sec 0. 0005 sec 0. 00010 sec
- 0. 04-0. 20 sec 0.0020 sec 0. 00020 sec
- 0. 20-1. 00 sec 0.0100 sec 0.00100 see Figure 3.2.4 shows the pressure results of the treak node (the compartment at which blowdown occurs). It is observed that doubling or halving the time step sizes given above does not result in significant difforences in pressure results. It was concluded that the pressure results are not sensitive to the time
step sizes chosen. The set of optimum sizes given above was therefore used for this analysis, Friction Factor Sensitivity The pressure losses prior to flow choking are caused primarily by expansion, contraction, and turning of the flow stream and are represented by form losses. The roughness of the cavity walls may contribute to the pressure losses, but RELAP4/ MODS accounts for them only through the use of Fanning-type friction losses which assume smooth walls within a cylindrical pipe. A sensitivity study was performed to ensure the insensitivity of the pressure results to wall friction considerations . The friction sensitivity study is described in Appendix C, which concludes that the pressure results are not sensitive to wall friction. Nodalization Sensitivity Study As specified in Reference 6.8, the subcompartment nodalization scheme must be verified by a sensitivity study which varies the number of nodes until the peak calculated pressures converge to small resultant changes. The nodalization sensitivity study is described in detail in Appendix B, which states that the pressure results are not sensitive to the nodalizction scheme chosen.
- 3. 3 Shielding Requirements
- 3. 3.1 Design Requirements 3.3.la Summary The criteria used for the neutron shield design presented in this report is to provide adequate shielding material capable of attenuating neutrone which originate in the core and stream up the reactor cavity by at least a factor of 40 at the reactor cavity flange level (Reference 6.15). This criteria was established when it was e'.dtermined that a shield would be necessary to reduce radiation levels measured inside the containment at the 38'6" elevation (Reference G.12) during plant operation to provide accessibility of personnel to the operating floor in the containment building. The radiation levels found at this level are attributed mainly to neutrons and gamma rays streaming up the reactor cavity annulus. In general, however, the neutron levels range from a factor of four to eight higher than the gamma levels at the 38'G" level and therefore gamma effects are not considered further in this analysis. Figure 3. 3-1 pre-sents the shield concept, giving the basic dimensions and geometry.
The shield is an annular ring which spans the reactor cavity, resting on the Reactor Vessel flange and the top of the concrete shield wall. The shield is divided in half for lifting purposes. Each half is partitioned aximuthally into eight equal sections with stainless steel gusset plates. Each section is filled with unborated water to a depth of sixteen inches. 3.3.lb Insulation Requirements The thickness of the thermal insulation around the reactor closure head bolting ring has been reduced to ellow closer placement of the neutron shield. This minhnizes neutron streaming in that region. Since the shield material to be used is water, sufficient insulation is provided to ensure that the shield temperature remains well
.below the boiling point.
3.3.lc Analysis Code For the neutron attenuation calculations, the S de DOT 3.5 N was used. This code was developed at the Oak Ridge National Laboratory. In recent years, it has become a standard code for reactor core and/or shield calculations. Many organizations have used DOT 3.5 and EDS considers it to embody the best and most cost efficient method for analysis of the problem at hand. 3.3.1d Source Terms Experimental values for the differential and integral energy flux at the flange level, in the absence of the shield, were taken from Reference 6.12 and they are reproduced here in Figure 3.3-2. The fluxes were used as the source term, serving as a starting point for the DOT 3.5 analysis. The major features of the reactor cavity at the flange level were modeled in a two-dimensional geometry for the shield analysis. The neutron cross sections used with DOT 3.5 were taken from the CASK (RSIC/DLC-23) 40-group coupled neutron and gamma ray library. CASK is widely used throughout the industry, and is a standard for light water reactor shield design. CASK gives 22-group neutron cross sections with anisotropy components in the differential cross section of up to P ' 3 Section 3.3.2 discusses in more detail the design inputs and the analytical approach. 3.3.2 Analytical Approach 3.3.2a Summary As in most radiation shieldag problems, the shield structure is complicated by necessary irregularities such as air gaps and
structural members which penetrate the shield. These irregu-larities create radiation streaming paths. which often are a major consideration in the final shield design. The most critical irrege-larity in the reactor cavity shield design under discussion here is the space required between the inner edge of the shield and the closure head as noted in section 3.3.1 b above. The presence of this gap affects the neutron transport so strongly that the following descrip-tion of the analytical approach begins by discussing the effects of this and other identified streaming paths. This is followed by a discussion of the shield geometry, the radiation source, and the DOT 3.5 computer code. 3.3.2b Streaming Ilegions As seen in Figure 3.3-1, the space provided between the shield inner edge and the closure head is 3-3/4 inches when cold. Due to the expansion of the closure head when heated to operating tempera-ture, the gap will close to approximately 3-j inches. Thermal insula-tion, 2-} inches thick, is required ia this space to keep the shield water below 212 F and to limit the heat loss to the area above the annulus. Since thermal insulation is relatively ineffective for the attenuation of neutrons, the 3-} inch space was modelled as a void space for the purpose of shield calculations. The largest fraction of the radiation at the level of the top of the shield is the result of this void space. Thus to a large extent the calculational problem becomes one of streaming through a gap. The shield is made sufficiently thick to attenuate most of the neutrons incident on the uader surface of the shield. Increasing the thickness of the shield beyond that specified has little effect in further reducing radiation levels above the shield. The neutron flux Incident on the reactor vessel flange is con-servatively assumed to be vertically upward and perpendicular to the flange. Steel has generally poor neutron attenuation
characteristics, however, every collision that a neutronundergoes in the flange changes the direction of the neutron which may then either be attentuated in the shield or in the . closure head. The shield is azimuthally compartmented into sixteen sections with j inch stainless steel gusset plates. Stainless steel also has generally poor neutron attenuation characteristics, but in this case, the thin plates contribute little because the resultant narrow beam geometry causes virtually all colliding neutrons within the steel to emerge into the water which then functions as an effective shield. Thus, in this case, almost all neutrons which enter the plate at the bottom of the. shield are deflected into the water before they can emerge at the top. Twenty-five kev neutrons have traditionally been found to stream through iron, as iron has a low neutron cross section at this energy. Ilowever, stainless steel contains nickel which fills the valley in the iron cross section at this energy, (Reference G.1G ). A calculation was performed based on methods from References G.17 and 6.18 which approximate neutron streaming through a steel plate immersed in water. This calculation indicated that the neutron streaming through the gusset plates is negligible. T.ie water shield overlaps the concrete snield wall by four inches, however, a 3-} inch space exists between the bottom of the shield and the refueling pool floor, Figure 3.3-1. The overlap cuts off the neutrons which are travelling upward toward the operating deck. The uncollided neutrons which stream through this gap will be travelling in a radially outward direction. The neutrons streaming through this gap in the east and west directions will be incident on the vertical concrete shield wall. In order for the streaming neutrons to reach the operating deck level, they would have to undergo
single or multiple scattering. The small fraction of these neutrons which reach the operating deck will therefore be degraded in energy due to the scattering. 3.3.2c Shield Geometry The annulus just below the flange level is about 3'2" wide. For the purposes of calculating neutron flux, the annulus was replaced by a straight gap with the same width as the annulus. With this simplification, the geometry becomes a two-dimensional, x(radial), y(vertical), geometry. This vastly reduces the computational difficulties. The approximation of the geometry by a two-dimensional (2-D) configuration does not introduce an appreciable error, since the shield is symmetric about its axis. The 2-D geometry does neglect the effects of neutrons streaming through the stainless steel gusset plates which divide the shield into sixteen azimuthal sections. As discussed earlier in Section 3.3.2b, this is not considered a significant effect. Thc shield blocks the gap between the refueling pool floor and the reactor vessel flange, leaving a 3-} inch space between the inner surface of the shield and the closure head. A 3-j inch space also exists between the refueling pool floor and the under surface at the outer edge of the shield, where it has a four inch overlap. The shield consists of unborated water, sixteen inches deep and thirty-eight inches wide in the radial direction. The water is contained in a stainless steel structure. The shield geometry is divided into a finite number of spatial regions, in a rectangular x-y geometry to input into the DOT program. The geometry as modeled in DOT is shown in Figure 3.3-3. In order to keep the problem within the core storage require-ments of DOT 3.5 code, the shield is split down its center line
and each half is analyzed separately. Note that the shield geometry is inverted because the DOT program runs mere efficiently with the radiation source on the top boundary. On the right hand side is a 10 cm wide section of steel to represent the closure head. Although the closure head is actually much thicker,for calculational purposes 10 cm is adequate. To the left and bottom of the shield, a void region is modeled. The boundary conditions chosen are as follows: right side - albedo condition; every neutron Jeaking from the right-hand boundary is assumed to be scattered back at the same position and energy in an isotropic distribution re-lative to the inward normal with albedo (reflection) probability of 0.8. left side - no incident flux, vacuum boundary. bottom - no incident flux, vacuum boundary. top - isotropic neutron flux is assumed except near the Reactor Vessel flange where uniform neutron flux directed vertically down-ward is assumed. At the shield centerline, where the shield is split and each half is analyzed separately, specular reflection boundary is assumed. 3.3.2d Radiation Source The radiation source for this analysis is based on the neutron flux measurements taken at the reactor vessel flange level in the annulus, (Reference 6.12). The total neutron fluence at the
flange level is 5.75 x 10 neutrons /cm see with an average energy of 0.15 MeV. The total fluence consists of fast neu-trons emerging from the core, plus intermediate and thermal neutrons which scatter off of the primary shield concrete wall. Figures 4 and 5 of Reference G.13 show that the measured thermal flux at the flange level is almost uniform across the annulus. Based on thi.s the total neutron flux is assumed uniform across the annulus. The angular distribution of t:1e neutrons has not been measured. Based on the geometry of the cavity, it can be assumed that the flux is travelling in a generally upward direction. For this analysis the flux incident on under surface of the shield is assumed to be isotropic in the upward direction except for the area directly under the flange. The isotropic flux approximates the effect of neutrons streaming outward through the 3-} inch space between the refueling pool floor and the under side of the shield. Directly underneath the flange the neutron flux is assumed to be travelling in a vertically upward direction, perpendicular to the flange. This clearly is the most conservative possible assumption, since the vertical neutrons will be the most penetrating and especially in view of the fact that there is the narrow vertical streaming path between the shields inner edge and the closure head that strongly favors the passage of vertically directed neutrons. The flux angularity is illustrated in Figure 3.3-3. 3.3.2e DOT 3. 5 DOT 3.5 is a two-dimensional discrete ordinates transport
theory program, developed at the Oak Ridge National Laboratory. The principal application is to solve the two-dimensional Boltzmann transport equations for neutrons by numerically integrating over the spatial regions and discrete neutron directions. Anisotropic scattering is allowed for a Legendre expansion of the differential cross sections up to tne thi-d order. A more detailed description of the DOT program is provided in Reference G.11 and the discrete ordinates theory is presented in detail in References 6.14 and G.19. DM 3.5 has provisions for solving a myriad of problems. For this antlysis, special features are used which make DOT 3.5 well adapted for shielding problems such as external boundary radiation source, albedo boundary conditions and scalar and angular flux distributions at any location in the problem geometry. The key feature of S methods is that the continuum of possible neutron directions is replaced by a finite number of appropriately chosen discrete directions. In DOT 3.5, the unit sphere of directions is broken up into a finite number of nonoverlapping sectors, each characterized by a solid angle (all the solid angles adding up to 411) and a characteristic discrete direction ordinate which make up the angular quadrature set. Average fluxes are computed for each angular sector. The spatial regions for the x-y geometry problem considered here must be rectangular. Figure 3.3-3 illustrates the spatial regions used in the calculations. Experimental calculations showed that further decreasing the spatial mesh size does not appreciably change the results. There are several types of angular quadrature sets available. The set used in this analysis is given Table 3. 3-1 and was developed at ORNL for use in the DM program. The quadra-ture set is an8S fu y symmeMc set, wM was mow by
adding a nearly vertical angle in order to input the neuttun flux in a vertically downward direction. Each direction is given in terms ofp, i> s direction cosine with respect to the x-axis (horizontal) andti, its direction cosine with respect to the y-axis (vertical). (The y-axis is orected downward in these calculations.) The entire geometry is assumed independent of the z spatial direction. Associated with each direction is a weight, w, taken proportional to the solid angle associated with each direction. The weights must sum to unity. It will be noted that there are some directions in Table 3.3-1 with 1; =0 and w=0. These directions have no effect in x-y geometry but are required by the DOT program. The cross sections used in DOT 3.5 were a 22-group neutron cross section set taken from the 22-neutron group,18-gamma-group set given by the cross-section averaging routine, CASK, also developed at ORNL. The group energy boundaries are given in Table 3.3-2. For each material the first four Legendre coefficients (P "' 0 3 given for the differential scattering cross see' ions between each pair of energy groups. l l i
3.4 Additional Considerations 3.h.1 Sump Clogging The neutron shicld is designed to minimize the potential for clogging of the recirculation sumps which provide fluid for safety injection and containment spray following a LOCA. Sump clogging is precluded by both the quantity of material which generate debris and by the location of the shield. 3.h.2 Fire Hazards The neutron shield is designed not to become a fire hazard. Materials of construction were selected on this basis. Since the neutron shield is fabricated from carbon and stairless steels, there is no combustion or N,re hazards associated with these materials. 3.h.3 Post LOCA Hydrogen Generetion The material in the neutron shiled have been selected so that no appreciable amounts of hy1rogen would be generated in a post LCCA environment. 3.h.h Containtaent Pressurization The neutron shield is designed such that the fluid inventory which would be released to the containment post-LOCA vill not significantly affect the containment pressure analysis.
. =
- 4. 0 ANALYTICAL RESULTS AND EVALUATION 4.1 Stuctural Results and Evaluation The results of each of the respective structural analyses is given in Sections a through f, below. Section g discusses the applicable Reference 6.1 loading combinations and evaluates the shield structure accordingly.
4.la Thermal Results The results of the steady state thermal analysis conducted on the Figure 3.1.6 model show that the water temperatures remained well below 212 F. The temperature reallts were averaged as shown in Figure 4.1.1 in order to determine the local and overall water tem-peratures. The region 1 temperature turned out to be 191 F while Regions 2 and 3 show temperatures of 183 F and 173 F respectively. As a result the water will not boil off or tend to pressurize the shield tank subcompartments. The results of the discontinuity thermal stress analysis in the area of the "C ring" were examined primarily to see if unusually high thermal stresses or thermally induced distortions occur in the shield structure. The stresses are self-limiting in nature and do not need to be included in the loading combinations as noted in Reference 6.1. The stresses were compared to the material yield stress in order to determine if gross plasticity occurs. The results showed that the critical section is just above the junction of the inner cylinder with the "C ring" in the region of the vent holes. The linearized bending stress in the ligaments is 32,100 psi. Comparing this to the average material yield of 31,250 psi at 2000F ve see that the actual stress is
' lightly above yield.
However, this stress poses no threat to the in-tegrity of the shield since the resulting small amount of plasticity occurs only at the surface fibers. This vill result in an insignificant amount of permanent distortion after thermal cycling has occurred.
The movements of the "C ring" due to therrral expansion are .5 inches radial and .040 inches axial. Note that adequate clearance is provided between the reactor flange and the "C ring" even with the inclusion of the thermal reactor flange motions. As a result, the reactor flange will not contact the "C ring and cause further distortions.
- 4. lb Pressure and Dead Load Results The results of the pressure plus dead loads analysis show that the only significant stresses occur mainly in the rupturc panels. This is shown in Table 4,1.1. Deadlot.d stresses are seen to be negligible as are the pressure stresses with the exception of the bulkhead at the pinrod connection.
The stress shown here is mostly a bending stress due to tha differential pressure acror.a the plate. 4.1c Lifting and Handling Results The lifting and handling results are also summarized in Table 4.1.1. The stresses shown are insignificant as in the pressure and dead load results. 4.1d Miscellaneous Results The miscellaneous results shown in Table 4. '.1 show that the unit panel stress is just belov _.se yield value for stainless material (yield = 31,250 at 200 F). Since the stress reported in Table 4.1.1'Is conservatively high, it is concluded that the panels will not deform or pose a safety hazard.
Seismic Results 4.le The results of the seismic analysis show that the significant stresses occur in the "C ring" and inner shell. Maximum "C ring" stresses wt;te found to ba 3,117 psi for OBE and 6,681 psi for DBE. Maximum inner shell stresses occurred in the flow hole ligaments and were 9,200 psi for OBE and 14,130 psi for DBE. 4, if LOCA Analysis Results
- 1. Burst Panel Analysis Results The time history behavior of the burst panels under the action of LOCA is shown in FIGURE 4.1.2. This figure shows the angular position and angular velocity obtained as a result of the interaction of the equations of motion derived in Appendix A. The curves shown cover the full time range up to the point when the panels fold against the aides of the tank. The point at which rupturing begins (1.c. the point when the strain at the groove apex has reached 100%)
is indicated on the curves as is the time at which the panel has fully opened. For reference and comparison purposes the pressure force history curve from Table 4. 2.6 is plotted in Figure 4.1.1. Note that the pressure curve has been converted to pst units and is shifted in time by . 005 seconds with respect to the Table 4. 2.6 time values. It is readily apparent that the panels begin rupturing very early in the transient (.0006 seconds) at an effective pressure of 85 psi. The rupturing process continues at an ever increasing rate until the panel fully opens at a time of .027 seconds. The dynamic burst pressure of 85 psi is much higher than the static burst pressure of 20 psi as expected. This is due to the snertial effect of the water and plate in moving the lower panel. Sir.ce the inertial effect of the water is not present in the top panel the
top panel deflects much more rapidly than the lower panel once the lower panel begins rupturing and exposes the upper panel to the cavity pressures. The top panel time history shows that bursting initiates at . 0112 seconds at a pressure of 100 psi. Figure 4.1.3 shows the post-burst behavior in terms of the pressure required to keep the panels opening. As can be seen, once the burst pressure has been exceeded the panels will continue to open since the required pressure falls off so quickly. Several conslusions can be drawn at this point with respect to the shield design and the conservatism of the analysis and assumptions. The burst panels open early enough in the transient and continue to open at a rapid enough rate to have little or no effect on the final pressure peak predicted by the thermal hydraulics analysis. This is demonstrated even more effect-ively in Section 4. 2.6. The opening vent area is controlled primarially by the lower panel due to its slower response except up to . 0114 seconds. However, the amount of vent opening at this point in time is so small (1. 3% open) that the initial lag of the upper panel can be neglected. In reality the lower panel would tend to respond faster than predicted for times larger than .01 seconds since the water which was assumed to add to the inertia of the lower plate would begin to be expelled by the flow from the reactor cavity. The vent area would be provided even faster than shown. As a result it can be concluded that the cavity pressures are upper bounded and would tend to reduce if the actual behavior of the panels were factored into the hydraulics analysis. 2 Results of Remaining Tank Structure The Mress results for the remaining tank structure are given in Table 4.1. 3.
The stress results for the remaining tank structure show that the worst bending stress to occur in the "C ring" is 24,900 psi. The stress in the ligament region is 2,736 psi. The worst case stress la the two shells is 15,228 ps! In the hoop direction. Stresses in the gusset plates ran less than 3,200 psi. These stress values include a dynamic amplification factor of 2.0 to account for the suddenly applied loads due to the rupture of the burst panels. These stresses are well below yield and are judged to be acceptable. The critical stress to examine is the "C ring" stress. If the stress there rises above the point of forming a plastic hinge then the ring will open up and the shield tanks would potentialy no longer be retained. This will only occur if the " C ring" stress rises to 35,250 which is the hinge moment stress based on elastic analysis and the yield stress at 550 F. As a result it is concluded that the shield will be retained during and after the LOCA event. 4.lg Load Combinations and Evaluation The stress results discussed above show that for the normal operating conditions (l.c. dead weight, thermal, lifting, etc.) the stresses in the shield structure are negligible. Accordingly, with the combination pro-ccedure called out in Reference 6.1 the reamining significant combination is for seismic plus LOCA. Again the critical region is the "C ring". If we directly add the DBE seismic bending stress in the "C ring" from Section
- 4. le to the LOCA stress reported above we have a stress of 31,641. This is still below the 35,250 limit discussed above and, therefore, the "C ring" will retain the shield for the combined event. It should be noted that there are many conservatisms in the above results so that the shield will not be as highly stressed as reported.
~ -
- 4. 2 Thermal-Ilydraulic Results 4.2.a Pressure History Results Final volume (node) pressure data were provided by the use of the RELAP4/ MODS computer r "le with the 33 node model. The 33 node model was chosen as a result of the nodalization sensitivity study (Appendix B) and was found to be reasonable and conservativ . Time histories of the reactor cavity are provided in Figure 4.2.1 through 33.
These pressure histories illustrate the relative responses of the reactor cavity subcompartments during the LOCA. The highest pressures and fastest pressure rise are observed at the break node (node'25) in Figure 4. 2. 25. A much slower response to the event is observed at the nodes at tha bottom of the cavity (nodes 4 and 5) in Figures 4. 2.4 and 4. 2. 5. This large distribution is caused by the very irregular geometry of the Millstone 2 reactor cavity. The shield dynamics equations described in Appendix E were solved numerically and the results of the final iteration are shown in Tables
- 4. 2.1 through 4. 2. 0. A comparison between these results and the assumed conditions loading to the results is given in Section 4. 2.b ,
below. 4.2.b Comparison Between Assumed and Computed Shield Resnonse The comparison between the assumed burst panel results (1. e. those used in the hydraulles analysis) and those actually computed is best made by examining the fraction of the area opened versus time. This comparison is made in Figure 4. 2. 34. From the curves shown it is readily apparent that the opening times for the burst panels is much faster than the response used in the hydraulics. This means that the pressure histories discussed above are upper bound curves since the actual opening of the sl.leid occurs much quick-er as a result it is concluded that the thermal hydraulle results presented are conservative as performed and need not be epeated.
4.3 Shielding Analysis Results 4.3a Attenuation Results The shield described in this report provides a calculated attenuation factor of 42. The attenuation factor is defined as: Gmrem/ D = 1 e 05 x 10 hr A. F. = _.a_vg(no shield) ., ,,,/ = 42
- 1) shielded) 4.2986 x 10 hr av where:
A. F. = the neutron dose rate attenuation factor D (no shield) = the average neutron dose rate at avg the flange level with no shield, based on the Reference 6.12 neutron flux measurements. D (shielded) = the neutron dose rate averaged over the top and outside edge of the shield, based on the DOT calculation. Figure 4.3-1 shows the dose rate distribution over the top and outer edge of the shield respectively. The flux-to-dose conversion factors are taken from Reference G.14. These factors compared closely to neutron flux-to-dose conversion factors listed in ANSt/ANS - G.1.1 - 1977 however, the Reference G.14 factors are slightly higher.
The estimated neutron dose rate d.rectly above the reactor cavity at the ope-ating deck (38'6") level is approx 1400 mrem /hr. This estimate is based on the angular distribution of the neutrons e_nerging from the shield and a geometry factor. The geometry factor is the factor by which the neutron dose rate decreases from the flange to the 38'6" level with no shield in place. Without the shield in place, the measured neutron dose rate above the reactor cavity at the 38'6" level is 6.0 x 10 mrem /hr (Reference 6.12). The dominant source of this dose rate is neutrons streaming up the reactor cavity. The unshielded neutron dose rate at the 6 flange level of the reactor cavity is 1.8205 x 10 mrem /hr. There-fore, due to the containment geometry, a reduction factor of about 30 is rssumed from the flange level to the operating deck level. This is considered a cs servative assumption because without the shield in place, the neutrons are travelling in a generally upward direction. The shield will diffuse the neutrons and the direction of the emerging neutrons will have more of an isotropic nature. With the shield in place, the geometry reduction factor probably will be larger than without the shield. As seen in Iigure 4.3-1, the dose rate on top of the shleid peaks in the gap between the shield and the closure head. This dose rate is conservatively high due to the assumption that the neutrons are vertically incident on the reactor vessel 11angc. The neutron fluxes used for the source term in this analysis were measured at flange level between the nozzles. This is clearly a conservative assumption, since the flux levels at this elevation can be a factor of two higher at points between the nozzles as compared to points directly abeve the nozzles
/3eference 6.13).
The addition of bomn would reduce the thermal neutron flux emerging from the shield. Ilowever, the neutron dose rate on top of the sideld is small compared to the dose rates near the < gaps, and further reduction of the dose rate through the shiel:1 uo uld have a minor effect. Due to the many inherent conser-vatisms in the calculation, measured attenuation factors are expect,ed to be superior to the above quoted values.
h.3b Activation Results The stainless steel structure of the neutron chield vill become activated by the neutrons escaping from the annr~us. Activation from both thermal and non-thermal energy neutrons was conside red. Since, the non-thermal activation was less than 7% of the thermal activation, it was neglected. For the input to the activation calculations, the output fluxes from DOT 3.5 were used. These fluxes with a shield, are higher in some cases than the measured fluxes without a shield because the shield acts as a reflector and thermalizer of neutrons. The output of DOT 3.5 lists the neutron flux Lhroughout the shield. In this manner, separate regions of the shield vere modeled for activation dose calculations. The results, as expected, show that the bottom plate of the shield, which is exposed to the highest thermal flux is the majcr contributor to the activation dose rates. The results of th3 activation dose calculations sre shown in Table h.3-1.
4.4 Additional Results 4.h.1 Sump Cloggine The neutron shield tank is designed to be retained on the reactor vessel flange following a LOCA and, therefore, will not contribute to any postulated concern regarding the containment sump. The lower insulation panels which would be eject,1 through the shield tanks vill, in all probability, impact the missile shield and return the refueling cavity or come to rest on the operating floor. Due to the long and tortuous path from the operating floor to the sump, it is highly unlikely that any insulation could migrate to the sump. Should some small amount of insulation manage to reach the contaimnent. lover level is would rapidly fill with water and sink harmlessly to the floor. Due to the very lov vater velocity toward the recirculation piping, the insulation would not tend to migrate one it has sunk. Even if all of the lover shield insulation were artifically placed on the sump screen, the insulation does not have sufficient surface area to completely block the screens. It is estimated that even under this extreme condition, approximately h5 square feet of screen surface would remain unblocked. h.h.2 Fire Hazard The four materials used in the neutron shield are stainless steel, vater, rock vool insulation and "Microtherm" insulation. These materials are all non-flamable. It is, therefore, concluded that the neutron shield does not present a fire hazard.
4.4.3 Post LOCA Hydrogen Generation The only material which could generate hydrogen in a post LOCA environment used in the neutron shield and associated support structure is zine used for galvanizing the anchor bolts for the support structure. It has been conservatively estimated that all of the zine coated anchor bolts have a total surface area of less than 3 5 ft2 Much of this surface area vill be embeded in concrete and not exposed to water in a post LOCA environment. This additional area of zine is trivial when compared to the 208,000 ft2 already in containment (FSAR Question 6.10.8) and will not substantially increase the hydrogen generation rate. h.h.h Containment Pressurization An engineering evaluation of the effect of the increased water inventory from the neutron shield on the containment pressure analysis has concluded that the water has a negligable effect. This conclucion applies to the ECCS performance calculations, where containment backpressure is minimized as well as the peak pressure analysis. h.h.5 Occupational Exposure Considerations
- a. Exposure from Initial Installation The neutron shield has been designed to minimize stay time hathe reactor cavity area during installation. The major man-hours required for assembly are located on top of one of the steam generators, vnere the dose rates are minimal
( 0.5 mrem /hr). However, the work inside the cavity is the major contributor to the total 3 man rem required for the , initial installation of the neutron shield.
- b. Exnosure Due to Removcl and Installction During Each Outage This shield has also been designed to minimize the time required for the removal and installation during each outage.
It should be noted that the activation of the shield contributes to approximately half of 2 man rem exposure required for this task while the rest is due to activation of other structures,
- c. Exposure Due to Storage of Shield on a Steam Generator The neutron shield is to be stored on top of one of the steam generators during the refueling outages. At this level, it will be shielded from most of the areas on the 38' 6" alevation by the steam generator valls. Areas which vill have significant exposures are work in the steam generator room and crane operation. Crane operation contributes to the major part of the 0 7 man rem per refueling outage for the storage of the shield on a steam generator.
- 5. 0 CONCLUSIONS Based on the analysis results and evaluation described herein it is concluded that the reactor cavity neutron shield will perform its intended design function during normal operation as well as during extremes and abnormal loading such as those due to a LOCA. The shirid design minimizes the personel exposure during in-stallation in keeping with ALARA considerations as well as providing a minimum attenuation of 40 over the current unshielded neutron fluxes. The reactor cavity surface weighted poak pressure was calculated to be 131 psi which is well below the 247 psig internal design pressure. Structural integrity of the shield was demonstrated along with the required conservatisms in applying the thermal hydraulic results to the structural analysis evaluations. As a result, the shield tank arrangement discussed above is acceptable for service in the Millstone, Unit 2, Nuclear Power Station.
- 6. 0 REFERENCES 6.1 USNBC Standard Review Plan 3. 8. 3, Concrete and Steel Internal Structures of Steel or Concrete Containments
- 6. 2 AISC, " Specification for Design, Fabrication and Erection of Structural Steel for Buildings,1969.
6.3 To be supplied
- 6. 4 ANRC-NUREG-1325, "RELAP4/ MOD 5: A Computer Program for Transient Thermal-Hydraulic Analysis of Nuclear Reactors and Related Systems",
Volumes 1,2 and 3, September 1976.
- 6. 5 Letter G/ME-77-806, November 18, 1977: Break Size Confirmation and Millstone FSAR Amendment 38, June 27,1975 (pages 5. 2-40 to 5. 2-41a, and
- 11. 2-3 to 11. 2-4).
- 6. 6 INESCO Originated 50336-311 (December 4,1975), 333 (November 25, 1975, 484 (June 30,1976) and 640 (December 10, 1976).
- 6. 7 Microfilm Records of Millstone 2 Reactor Cavity Pressurization Analys's, 1973, Reel No. 2033.
- 6. 8 Standard Review Plan Section 6. 2.1. 2, "Subcompartment Analysis",
February 1975.
- 6. 9 " Analysis of NRC Standard Problems for Subcompartment Pressurization" CDC Utilities Service Center, Rockville, Maryland, May 1976.
6.10 Metcalfe, L.J. , et al, " Containment Analysis Capabilities of CONTEMPT 4/ MOD 2", ANS/ ENS Joint Conference, Brussels,1978. 6.11 " DOT 3 Two Dimension Discrete Ordinates Transprot Code", ORNL-TM-4280 (September 1973) 6.12 Kacich, Richard M. , NEUTRON SPECTRA AND FLUENCE DETERMINATION AT MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2, Northeast Utilities Service Company, November 1977. 6.13 Celnik, J. , METHODOLOGY FOR ANALYSIS AND DESIGN OF A PWR REACTOR CAVITY SHIELD SYSTEM, Trans. Am. Nucl. Soc. , Volume 30, November 1978, page 237.
- 6. 0 References Continued 6.14 Schaeffer, N.M. , REACTOR SHIELDING FOR NUCLEAR ENGINEERS, TID-25921 U. S. A. E. C. , 1973.
6.15 SPECIFICATION FOR PERFORMANCE OF ENGINEERING DESIGN AND ANALYSIS FOR REACTOR CAVITY NEUTRON SHIELD AT MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2, SP=ME-130, prepared by S. J. Weyland, Northeast Utilities Services Co. , June 6,1977. 6.16 Price, Horton, and Spinney, RADIATION SHIELDING, Pergamon Press,1957 6.17 Rockwell, T. , REACTOR SHIELDING DESIGN MANUAL, Van No Strand Company, Inc. , 1956.
- 6. 18 Blizard, E. P. , REACTOR HANDBOOK, Volume, Interscience Publishers,1962.
6.19 Mynatt, F.R. , et. al. DEVELOPMENT OF TWO-DIMENSIONAL DISCRETE ORDINATES TRANSPORT THERORY FOR RADIATION SHIELDING, Oak Ridge National Laboratory CTC-INF-952, August 11, 1969. 6 20 ANRC-NUREG-1325, "RELAP4/ MODS: A Computer Program for Transient Thermal-Hydraulle Analysis of Nuclear Reactors and Related Systems", Volumes 1,2 and 3, September 1976. 6.21 IdePChik, I. E. , " Handbook of Hydraulic Resistance", AEC-tr-6630, U. S. D of C,1966. 6 22 " Flow of Fluids Through Valves, Fittings, and Pipe",15th Edition, Crane Company, New York,1976.
TABLE 3. 2.1 RELAP4 VOLUME DATA (33 NODE MODEL) NODE NO. NODE IDENTIFICATION VOLUME VOLUME ELEVATION 3 (Ft ) HEIGHT AT BOTTOM (Ft) (Ft) 1 Containment 1870000.0 121.25 -23.25 2 East Steam Generator Rm. 35000.0 63.50 - 0.50 3 West Steam Generator Rm. 33000.0 63.50 - 0.50 4 Reactor Cavity (Bottom 586.7 9.45 -23.25 Node) 5 Reactor Cavity (Bottom 1212.7 9.45 -23.25 Node) 6 Reactor Cavity (LowerNode) 160.8 7.88 -13.80 7 Reactor Cavity (LowerNode) 142,9 7.88 -13.80 8 Reactor Cavity (LowerNode) 144.5 7.88 -13.80 9 Reactor Cavity (LowerNode) 145.5 7.88 -13.80 10 Reactor Cavity (IowerNode) 114.4 7.88 -13.80 11 Reactor Cavity (LowerNode) 66.0 7.88 -13.80 12 Reactor Cavity (LowerNode) 111,6 7.88 -13.80 13 Reactor Cavity (MiddleNode) 105.4 8.34 - 5.92 14 Reactor Cavity (Middle Node) 98.3 8.34 - 5.92 15 Reactor Cavity (Middle Node) 95.8 8.34 - 5.92 16 Reactor Cavity (MiddleNode) 96.8 8.34 - 5.92 17 Reactor Cavity (MiddleNode) 90.8 8.34 - 5.92 18 Reactor Cavity (Middle Node) 18.8 8.34 - 5.92 19 Reactor Cavity (MiddleNode) 90.3 8.34 - 5.92 20 Reactor Cavity (Penetration 112.7 6.83 2.42 Node) 21 Reactor Cavity (Penetration 124.8 6.83 2.42 Node) 22 Reactor Cavity (Penetration 117.9 6.83 2.42 Node) 23 Reactor Cavity (Penetration 118.4 6.83 2.42 Node) 24 Reactor Cavity (Penetration 97.9 6.83 2.42 Node) 25 Reactor Cavity (BreakNode) 57.9 6.83 2.42 26 Reactor Cavity (Penetration 77.7 6.83 2.42 Node) 27 Reactor Cavity (UpperNode) 101.5 3.21 - 9.25 28 Reactor Cavity (UpperNode) 101.5 3.21 9.25
TABLE 3.2.1 CONTINUED NODE NO. NODE IDENTIFICATION VOLUME VOLUME ELEVATION 3 (Ft ) HEIGHT AT BOTTOM (Ft) (Ft) 29 Reactor Cavity (Upper Node) 101.5 3.21 9.25 30 Reactor Cavity (Upper Node) 101.5 3.21 - 9.25 31 Reactor Cavity (Upper Node) 67.6 3.21 9.25 32 Reactor Cavity (Upper Node) 67.6 3.21 - 9.25 33 React 6r Cavity (Upper Node) 67.6 3.21 9.25
TABLE 3. 2. 2 RELAP4 JUNCTION DATA (33 NODE MODEL) JUNCTION FROM TO JUNCTION JUNCTION FORWARD REVERSE NO. NODE NODE AREA
- INERTIA LOSS COEFF LOSS COEFF (Ft2) (Ft~l) 1 4 5 93.36 0.148 0.010 0.010 2 5 6 20.40 0.193 0,500 1.000 3 5 7 18.13 0.217 0.500 1.000 4 5 8 18.33 0.215 0.500 1.000 5 5 9 18.46 0.213 0.500 1.000 6 4 10 14.51 0. 2 71 0.500 1.000 7 4 11 8.25 0.470 0.500 1.000 8 4 12 14.16 0.278 0.500 1.000 9 6 12 19.23 0.325 0.345 0.345 10 6 7 7.09 0.649 0. 71 4 0.714 11 7 8 19.23 0.399 0.345 0.345 12 8 9 7.09 0.693 0. 71 4 0.714 13 9 10 19.23 0.325 0.345 0.345 14 10 11 7.09 0.540 0. 71 4 0. 71 4 15 11 12 13.67 0.429 0.393 0. 3 91 16 6 13 12.65 0.523 0.190 0.145 17 7 14 .11.80 0. 5 71 0.1 75 0.122 18 8 15 11.50 0.578 0.186 0.139 19 9 16 11.62 0.573 0.185 0.137 20 10 17 10.89 0.654 0.125 0.062 21 11 18 2.25 2.324 0.364 0.529 22 12 19 10.84 0.663 0.117 0.055 23 13 19 15.76 0.391 0.368 0.368 24 13 14 2.50 1.337 1.128 1.128 25 14 15 15.76 0.500 0.368 0.368 26 15 16 2. 75 1.626 1.094 1.094 27 16 17 15.76 0.391 0.368 0.368 28 17 18 2.50 0.737 1.128 1.128 29 18 19 7.01 0.764 1.072 0.713 30 13 20 5.40 0.634 0.988 0.747 31 14 21 5. 81 0.682 0.934 0.670 32 15 22 5.58 0.692 0.949 0. 681 33 16 23 5.70 0.693 0.932 0.674 34 17 24 6.40 0. 72 6 0.711 0.525 35 18 25 1.72 4.476 0.968 0. 4 81
TABLE 3.2.2 CONTINUED JUNCTION FROM TO JUNCTION JUNCTION FORWARD REVERSE NO. NODE NODE A RE A* INERTIA LCSS COEFF LOSS COEFF (Ft2) (Ft-1) 36 19 26 7.83 0.730 0.556 0.400 37 20 26 4.18 1. 71 8 1.500 1.500 38 20 21 5.77 2.062 1.500 1.500 39 21 22 7.63 1.130 1.500 1.500 40 22 23 3.27 2.644 1.500 1.500 41 23 24 7.63 0.941 1.500 1.500 42 24 25 10.46 0.549 1.500 1.500 43 25 26 12.38 0.464 1.500 1.500 44 20 27 33.20 0.151 0.005 0.005 45 21 28 33.20 0.151 0.005 0.005 46 22 , 29 33.20 0.151 0.005 0.005 47 23 30 33.20 0.151 0.005 0.005 48 24 31 22.13 0.227 0.005 0.005 49 25 32 22.13 0.227 0.005 0.005 50 26 33 22.13 0.227 0.005 0.005 51 27 33 9.91 0. 847 0.084 0.084 52 27 28 9.91 1.016 0.100 0.100 53 28 29 9.91 1. 01 6 0.100 0.100 54 29 30 9.91 1.016 0.100 0.100 55 30 31 9.91 0.847 0.084 0.084 56 31 32 9.91 0.678 0.070 0.070 57 32 33 9.91 0.678 0.070 0.070 58 1 27 24.04 0.051 0.576 1.500 59 1 28 24.04 0.051 0.576 1.500 60 1 29 24.04 0. 051 0.576 1.500 61 1 30 24.04 0.051 0.576 1.500 62 1 31 16.03 0.076 0.576 1.500 63 1 32 16.03 0.076 0.576 1.500 64 1 33 16.03 0.076 0.576 1.500 65 2 20 5.69 0.952 1.500 1.500 66 2 21 3.29 2.024 1.500 1.500 67 3 21 3.29 2.962 1.500 1.500 68 3 22 5.69 0.794 1.500 1.500 69 3 23 5.69 0.952 1.500 1.500 70 3 24 3.29 2.024 1.500 1.500 71 2 25 6.59 1.481 1.500 1.500 72 2 26 2.40 2.188 1.500 1.500 73 1 2 134.02 0.010 1.500 1.500 74 1 3 134.02 0.010 1.500 1.500 75 0 25 1. 0 0. 0. 0 0. 0 76 0 25 1. 0 0. 0. 0 0. 0
- A coefficient of 0. 6 is applied to the Moody critical flow values in accordance with Ref. 6.1 rather than reducing all Junction areas to 0. 6.
TABLE 3.3-1 S: Angular Quadrature ANGLE WEIGHT ETA (n) MU(u) 1 0. .999657+00 .26188-01 2 .85710-04 .999657+00 .10000-04 3 .85710-04 .999657+00 .10000-04 4 0. .951190+00 .30861+00 5 .30161-01 .951190+00 .21822+00 6 .30161-01 .951190+00 .21822+00 7 0. .786800+00 .61721+00 8 .22685-01 .786800+00 .57735+00 9 .22685-01 .786800+00 .21822+00 10 .22685-01 .786800+00 .21822+00 11 .22685-01 .786800+00 .57735+00 12 0. .577350+00 .81650+00 13 .22685-01 .577350+00 .78680+00 14 .23148-01 .577350+00 .57735+00 15 .22685-01 .577350+00 .21822+00 16 .22685-01 .577350+00 .21822+00 17 .23148-01 .577350+00 .57735+00 18 .22685-01 .577350+00 .78680+00 19 0. .218220+00 .97590+00 20 .30247-01 .218220+00 .95119+00 21 .22685-01 .218220+00 .78680+00 22 .22685-01 .218220+00 - 57735+00 23 .30247-01 .218220+00 .21822+00 24 .30247-01 .218220+00 .21822+00 25 .22685-01 .218220+00 ' .57735+00 26 .22685-01 .218220+00 .78680+00 27 .30247-01 .218220+00 .95119+00 28 0. .999657+00 .26180-01 29 .85710-04 .999657+00 .10000-04 30 -
.85710-04 .999657+00 .10000-04 31 0. .951190+00 .30861+00 32 .30161-01 .951190+00 .21822+00 33 .30161-01 .951190+00 .21822+00 34 0. .786800+00 .61721+00 35 .22685-01 .786800+00 .57735+00 36 .22685-01 .786800+00 .21822+00 37 .22685-01 .786800+00 .21822+00 38 .22685-01 .786800+00 .57735+00 39 0. .577350+00 .81650+00 40 .22685-01 .577350+00 .78680+00 41 .23148-01 .577350+00 .57735+00 42 .22685-01 .577350+00 .21822+00 43 .22o85-01 .57/350+00 .21622+00 44 .23148-01 .577350+00 .57735+00 45 .22685-01 .577350+00 .78680+00 46 0. .218220+00 . .97590+00 47 .30247-01 .218220+00 .95119+00 48 .22685-01 .218220+00 .78680+00 49 .22685-01 .218220+00 .57735+00 50 .30247-01 .218220+00 .21822+00 51 .30247-01 .218220+00 .21822+00 52 .22685-01 .218220+00 .57735+00 53 .22685-01 .218220+00 .78680+00 54 .30247-01 .218220+00 .95119+00
TABLE 3.3-2 DLC-23/ CASK ENERGY GROUPS STANDARD 22 GROUP NEUTRON STRUCTURE ENERGY GROUP ENERGY RANGE (EV) 1 1.4918+07 1.2214+07 2 1.2214+07 1.0000+07 3 1.0000+07 8.1873+06 4 8.1873+06 6.3600+06 5 6.3600+06 4.9639+06 6 4.9659+06 4.0657+06 7 4.0657+06 3.0119+06 8 3.0119+06 2.4660+06 9 2.4660+06 2.3500+06 10 2.3500+06 1.8268+06 11 1.8268+06 1.1080+06 12 1.1080+06 5.5023+05 13 5.5023+05 1.1109+05 14 1.1109+05 3.3546+03 15 3.3546+03 5.8295+02 16 5.8295+02 1.0130+02 17 1.0130+02 2.9023+01 18 2.9023+01 1.0677+01 19 1.0677+01 3.0590+00 20 3.0590+00 5.3159-01 21 5.3159-01 4.1400-01 22 4.1400-01 1.0000-05
Load Case Locadon Value (Psi) Dead Weight Hoop (S ) outer wall, top -171 Iloop (S ) outer wall, bottom 178 Axial (S ) outer wall, bottom -72 Hoop (Sy inner wall, top -253 Hoop (S ) inner wall, bottom 228 Anywhere - Bulkheads < 100 Hydrostatic + 1 psi Hoop (S ) outer wall, top -303 internal pressure Hoop (Sy outer wall, bottom 394 Axial (S ) outer wall, bottom -311 Hoop (S ) inner wall, top -397 Hoop (S ) inner wall, bottom 30G Axial (S ,) inner wall, bottom 350 Axial (S ) bulkhead at pin end bottom 4452 Lifting and Handling Hoop (S } outer wall, top 796 Hoop (S ) outer wall, bottom -1784 Axial (S ) outer wall, bottom -466 Hoop (S) inner wall, ' top 1975 Hoop (S ) inner wall, bottom -274 Axial (S ) inner wall, bottom -611 Miscellaneous Loading in top burst panel 29,980 TABLE 4.1.1
SUMMARY
OF VARIOUS NORMAL OPERATIONAL STRESSES IN TIIE SHIELD TANKS
s . TABLE 4.2.1 SHIELD PLUG DYNAMICS RESULTS ABOVE VOLUME 27 TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Ft/See) (Ft) FLOW AREA
- 0.0000 0. O. O. O.
0.0020 0. O. O. O. 0.0040 0. O. O. O. 0.0060 0. O. O. O. 0.0080 0. O. O. O. 0.0100 0. O. O. O. 0.0120 0. O. O. O. 0.0140 184.9 0.04 0.000 0.0000 0.0160 498.1 0.15 0.000 0.0001 0.0180 924.8 0.36 0.001 0.0003 0.0200 1477.0 0.68 0.002 0.0007 0.0220 2145.0 1.16 0.004 0.0015 0.0240 2900.0 1.80 0.007 0.0026 0.0260 3642.0 2.60 0.011 0.0044 0.0280 4291.0 3.55 0.017 0.006] 0.0300 4880.0 4.63 0.026 0.0102 0.0320 5454.0 5.84 0.036 0.0144 0.0340 6035.0 7.17 0.049 0.0196 0.0360 6616.0 8.64 0.065 0.0260 0.0380 7178.0 10.22 0.084 0.0336 0.0400 7714.0 11.93 0.106 0.0425 0.0420 8198.0 13.74 0.131 0.0528 0.0440 8605.0 15.65 0.161 0.0646 0.0460 8923.0 17.62 0.194 0.0780 0.0480 9162.0 19.65 0.231 0.0930 0.0500 9331.0 21.71 0.273 0.1096 0.0520 9441.0 23.80 0.318 0.1279 0.0540 9515.0 25.90 0.368 0.1479 0.0560 9568.0 28.02 0.422 0.1696 0.0580 9611.0 30.14 0.480 0.1930 0.0600 9657.0 32.28 0.542 0.2181 0.0620 9690.0 34.42 0.609 0.2450 0.0640 9701.0 36.57 0.680 0.2735 0.0660 9680.0 38.71 0.755 0.3038 0.0680 9632.0 40.84 0.835 0.3358 0.0700 9567.0 42.96 0.919 0.3695 0.0720 9490.0 45.05 1.007 0.4049 0.0740 9436.0 47.14 1.009 0.4420 0.0760 9402.0 49.22 1.95 0.4808 0.0780 9391.0 51.30 1.296 0.5212 0.0800 9390.0 53.37 1.400 0.5633 0.0820 9395.0 55.45 1.509 0.6071 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4.2.1 CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Ft/Sec) (Pt) FLOW AREA
- C.0840 9402.0 57.33 1.622 0.6525 0.0860 9451.0 59.62 1.739 0.6996 0.0880 9527.0 61.73 1.861 0.7485 0.0900 9628.0 63.86 1.986 0.7990 0.0920 S770.0 66.02 2.116 0.8512 0.094. 9962.0 68.22 2.250 0.9052 0.0960 10100.0 70.46 2.389 0.9610 0.0980 10200.0 72.67 2.532 1.0000 1.0000 9943.0 1.1020 9935.0 1.1040 9952.0 1.1060 9971.0 1.1080 9986.0 1.1100 10010.0 1.1120 10040.0 1.1140 10100.0 1.1160 10180.0 1.1180 10280.0 1.1200 10400.0 1.1220 10530.0 1.1240 10650.0 1.1260 10760.0 1.1280 10850.0 1.1300 10940.0 1.1320 11020.0 1.1340 11090.0 1.1360 11160.0 1.1380 11230.0 1.1400 11300.0 1.1420 11370.0 1.1440 11440.0 1.1460 11500.0 1.1480 11560.0 1.1500 11610.0 1.1520 11650.0 1.1540 11680.0 1.1560 11700.0 1.1580 11720.0 1.1600 11730.0 1.1620 11740.0 1.1640 11750.0 1.1660 22750.0 1.1680 11750.0 1.1700 11740.0 1.1720 11740.0 1.1740 11730.0 1.1760 11730.0 1.1780 11720.0
TABLE 4.2.2 SHIELD PLUG DYNAMICS RESULTS ABOVE VOLUME 28 TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Ft/Sec) (Ft) FLOW AREA
- 0.0000 O. O. O. O.
0.0020 0. O. O. O. 0.0040 0. O. O. O. 0.0060 0. O. O. O. 0.0080 0. O. O. O. 0.0100 0. O. O. O. 0.0120 0. O. O. O. 0.0140 0. O. O. O. 0.0160 0. O. O. O. 0.0180 0. O. O. O. 0.0200 0. O. O. O. 0.0220 84.7 0.02 0.0000 0.0000 0.0240 253.5 0.07 0.0001 0.0001 0.0263 467.4 0.18 0.0004 0.0002 0.0280 725.1 0.34 0.0009 0.0004 0.0300 1020.0 0.56 0.0018 0.0009 0.0320 1347.0 0.86 0.0032 0.0016 0.0340 1703.0 1.24 0.0053 0.0026 0.0360 2089.0 1.70 0.0083 0.0041 0.0380 2499.0 2.25 0.0122 0.0060 0.0400 2900.0 2.90 0.0174 0.0086 0.0420 3286.0 3.62 0.0239 0.0118 0.0440 3666.0 4.43 0.0319 0.0158 0.0460 4045.0 5.33 0.0417 0.0206 0.0480 4424.0 6.31 0.0533 0.0263 0.0500 4807.0 7.37 0.0670 0.0331 0.0520 5189.0 8.52 0.0829 0.0409 0.0540 5560.0 9.75 0.1011 0.0499 0.0560 5908.0 11.05 0.1219 0.0602 0.0580 6227.0 12.43 0.1454 0.0717 0.0600 6519.0 13.87 0.1717 0.0847 0.0620 6777.0 15.37 0.2010 0.0991 0.0640 6998.0 16.92 0.2332 0.1151 0.0660 7180.0 18.51 0.2687 0.1325 0.0680 7320.0 20.13 0.3073 0.1516 0.0700 7422.0 21.77 0.3492 0.1723 0.0720 7481.0 23.42 0.3944 0.1946 0.0740 7520.0 25.08 0.4429 0.2185 0.0760 7537.0 26.75 0.4947 0.2441 0.0780 7530.0 28.42 0.5499 0.2713 0.0800 7495.0 30.07 0.6084 0.3001 0.0820 7438.0 31.72 0.6702 0.3306 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4. 2. 2 CONTINUED TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (See) (LBf/Ft 2) (Ft/Sec) (Pt) FLOW AREA
- 0.0840 7340.0 33.34 0.7352 0.3627 0.0860 7233.0 34.94 0.8035 0.3964 0.0880 7121.0 346.52 0.8750 0.4317 0.0900 7008.0 38.07 0.9496 0.4685 0.0920 6888.0 39.59 1.027 0.5068 0.0940 6779.0 41.09 1.108 0.5466 0.0960 6659.0 42.56 1.192 0.5878 0.0980 6546.0 44.01 1.278 0.6306 0.1000 6461.0 45.44 1.368 C.6747 0.1020 6396.0 46.85 1.460 0.7202 0.1040 6385.0 48.27 1.555 0.7671 0.1060 6385.0 49.68 1.653 0.8155 0.1080 6383.0 51.08 1.754 0.8652 0.1100 6392.0 52.50 1.857 0.9163 0.1120 6293.0 53.90 1.964 0.9688 0.1140 6211.0 55.27 2.073 1.000 0.1160 6203.0 0.1180 6243.0 0.1200 6309.0 0.1220 6395.0 0.1240 6484.0 0.1260 6560.0 0.1280 6624.0 0.1300 6678.0 0.1320 6727.0 0.1340 6780.0 0.1360 6838.0 0.1380 6911.0 0.1400 6992.0 0.1420 7082.0 0.1440 7176.0 0.1460 7268.0 0.1480 7352.0 0.1500 7428.0 0.1520 7485.0 0.1540 7533.0 0.1560 7577.0 0.1580 7618.0 0.1600 7656.0 0.1620 7691.0 0.1640 7723.0 0.1660 7753.0 0.1680 7780.0 0.1700 7806.0 0.1720 7829.0 0.1740 7851.0 0.1760 7871.0 0.1780 7889.0
TABLE 4.2.2 CONTINUED TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Ft/Sec) (Pt) FLOW AREA
- 0.1800 7907.0 0.1820 7923.0 0.1840 7939.0 0.1860 7054.0 0.1880 7969.0 0.1900 7983.0 0.1920 7996.0 0.1940 8010.0 0.1960 8022.0
TABLE 4.2.3 SHIELD PLUG DYNAMICS RESULTS ABOVE VOLUME 29 TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Ft/See) (Ft) FLOW AREA
- 0.0000 0. O. O. O.
0.0020 0. O. O. O. 0.0040 0. O. O. O. 0.0060 0. O. O. O. 0.0080 0. O. O. O. 0.0100 0. O. O. O. 0.0120 0. O. O. O. 0.0140 0, 0. O. O. 0.01GO 0. O. O. O. 0.0180 0. O. O. O. 0.0200 0. O. O. O. 0.0220 77.5 0.02 0.0000 0.0000 0.0240 236.3 0.07 0.0001 0.0000 0.0260 436.4 0.17 0.0003 0.0001 0.0280 680.3 0.32 0.0008 0.0003 0.0300 964.6 0.53 0.0017 0.0007 0 0320 1285.0 0.81 0.0030 0.0012 0.J340 1641.0 1.18 0.0050 0.0020 0.0360 2039.0 1.62 0.0078 0.0031 0.0380 2474.0 2.18 0.0116 0.0047 0.0400 2916.0 2.82 0.0166 0.0067 0.0420 3353.0 3.56 0.0230 0.0092 0.0440 3781.0 4.40 0.0309 0.0125 0.0460 4194.0 5.33 0.0407 0.0164 0.0480 4593.0 6.34 0.0523 0.0211 0.0500 4977.0 7.44 P0661 0.0266 0.0520 5349.0 8.63 0.0022 0.0331 0.0540 5708.0 9.89 0 .007 0.0405 0.0560 6052.0 11.23 0.1218 0.0490 0.0580 6373.0 12.64 0.1457 0.0586 0.0600 6671.0 14.11 0.1724 0.0694 0.0620 6938.0 15.65 0.2022 0.0813 0.0640 7164.0 17.23 0.2350 0.0946 0.0660 7346.0 18.85 0.2711 0.1091 0.0680 7481.0 20.51 0.3105 0.1249 0.0700 7570.0 22.18 0.3532 0.1421 0.0720 7615.0 23.87 0.3992 0.1606 0.0740 7634.0 25.56 0.4487 0.1805 0.0760 7627.0 27.24 0.5015 0.2017 0.0780 7590.0 28.92 0.5576 0.2243 0.0800 7521.0 30.58 0.6171 0.2482 0.0820 7419.0 32.23 0.6799 0.2735 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4.2.3 CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Ft/Sec) (Ft) FLOW AREA
- 0.0840 7273.0 33.83 0.7460 0.3001 0.0860 7111.0 35.41 0.8152 0.3279 0.0880 6937.0 36.94 0.8876 0.3570 0.0900 6755.0 38.44 0.9630 0.3873 0.0920 6564.0 39.89 1.041 0.4189 0.0940 6381.0 41.30 1.122 0.4515 0.0960 6191.0 42.67 1.206 0.4853 0.0980 6011.0 44.00 1.293 0.5201 0.1000 5860.0 43.29 1.382 0.5561 0.10"0 5734.0 46.56 1.474 0.5930 0.1040 5663.0 47.81 1.540 0.6710 0.1060 5622.0 49.06 1.665 0.6699 0.1080 5581.0 50.29 1.765 0.7099 0.1100 5570.0 51.52 1.867 0.7509 0.1120 5603.0 52.76 1.971 0.7928 0.1140 5676.0 54.02 2.078 0.8358 0.1160 5749.0 55.29 2.187 0.8797 0.1180 5787.0 56.57 2.299 0.9247 0.1200 5839.0 57.86 2.413 0.9708 0.1220 5926.0 59.17 2.530 1.000 0.1240 6037.0 0.1260 6137.0 0.1280 6215.0 0.1300 6274.0 0.1320 6328.0 0.1340 6397.0 0.1360 6490.0 0.1380 6620.0 0.1400 6769.0 0.1420 6936.0 0.1440 7105.0 0.1460 7262.0 0.1480 7399.0 0.1500 7513.0 0.1520 7596.0 0.1540 7658.0 0.1560 7714.0 0.1580 7763.0 0.1600 7804.0 0.1620 7839.0 0.1640 7867.0 0.1660 7888.0 0.1680 7903.0 0.1700 7914.0 0.1720 7922.0 0.1740 7937.0 0.1760 7931.0 0.1780 7936.0
TABLE 4. 2. 3 CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Pt/Sec) (Ft) FLOW AREA
- 0.1800 7940.0 0.1820 7946.0 0.1840 7953.0 0.1860 7961.0 0.1880 7971.0 0.1900 7982.0 0.1920 7993.0
TABLE 4.2.4 SHIELD PLUG DYNAMICS RESULTS ABOVE VOLUME 30 TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (See) (LBf/Ft 2) (Pt/See) (Ft) FLOW AREA
- 0.0000 0. O. O. O.
0.0020 0. O. O. O. 0.0040 0. O. O. O. 0.0060 0. O. O. O. 0.0080 0. O. O. O. 0.0100 0. O. O. O. . 0.0120 0. O. O. O. 0.0140 180.0 0.04 0.0000 0.0000 0.0160 487.1 0.15 0.0002 0.0001 0.0180 896.1 0.35 0.0007 0.0003 0.0200 1420.0 0.66 0.0017 0.0007 0.0220 2058.0 1.12 0.0035 0.0014 0.0240 2787.0 1.73 0.0063 0.0026 0.0260 3552.0 2.52 0.0106 0.0043 0.0280 4235.0 3.45 0.0166 0.0067 0.0300 4848.0 4.53 0.0246 0.0099 0.0320 5448.0 5.73 0.0348 0.0140 0.0340 6076.0 7.08 0.0476 0.0192 0.0360 6745.0 8.57 0.0633 0.0254 0.0380 7437.0 10.21 0.0820 0.0330 0.0400 8124.0 12.01 0.1042 0.0419 0.0420 8759.0 13.95 0.1302 0.0524 0.0440 9303.0 16.00 0.1602 0.0644
.0460 9735.0 18.16 0.1943 0.0782 0.0480 10070.0 20.38 0.2328 0.0937 0.0500 10320.0 22.67 0.2759 0.1110 0.0520 10510.0 24.99 0.3236 0.1301 0.0540 10670.0 27.35 0.3759 0.1512 0.0560 10820.0 29.74 0.4330 0.1742 0.0580 10960.0 32.17 0.4949 0.1991 0.0600 11100.0 34.62 0.5617 0.2259 0.0620 11230.0 37.11 0.6344 0.2548 0.0640 11320.0 39.61 0.7101 0.2857 0.0660 11370.0 42.12 0.7919 0.3185 0.0680 11370.0 44.64 0.8786 0.3534 0.0700 11320.0 47.14 0.9704 0.3903 0.0720 11220.0 49.62 1.067 0.4293 0.0740 11100.0 52.08 1.169 0.47,2 0.0760 10980.0 54.51 1.275 0.5131 0.0780 10860.0 56.91 1.387 0.5579 0.0800 10730.0 59.28 1.503 0.6046 0.0820 10600.0 61.63 1.624 0.6532 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4.2.4 CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Pt/Sec) (Ft) FLOW AREA
- 0.0840 10470.0 63.94 1.750 0.7038 0.0860 10390.0 66.24 1.880 0.7561 0.0880 10350.0 68.53 2.015 0.8103 0.0900 10340.0 70.82 2.154 0.8664 0.0920 10380.0 73.11 2.298 0.9243 0.0940 10480.0 75.43 2.446 0.9840 0.0960 10540.0' 77.76 2.600 1.000 0.0980 10370.0 0.1000 10200.0 0.1020 10110.0 0.1040 10050.0 0.1060 9999.0 0.1080 9957.0 0.1100 9928.0 0.1120 9926.0 0.1140 9966.0 0.1160 10040.0 0.1180 10160.0 0.1200 10300.0 0.1220 10470.0 0.1240 10650.0 0.1260 10810.0 0.1280 10970.0 0.1300 11120.0 0.1320 11270.0 0.1340 11400.0 0.1360 11530.0 0.1380 11650.0 0.1400 11760.0 0.1420 11870.0 0.1440 11980.0 0.1460 12090.0 0.1480 12180.0 0.1500 12280.0 0.1520 12340.0 0.1540 12400.0 0.1560 12450.0 0.1580 12500.0 0.1600 12540.0 0.1620 12580.0 0.1640 12610.0 0.1660 12640.0 0.1680 12660.0 0.1700 12670.0 0.1720 12680.0 0.1740 12690.0 0.1760 12700.0 0.1780 12700.0
TABLE 4.2.4 CONTINUEP TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Ft/Sec) (Pt) FLOW AREA
- 0.1800 12710.0 0.1820 12710.0 0.1840 12720.0 0.1860 12720.0 0.1880 12720.0 ~
0.1900 12730.0 0.1920 12730.0 0.1940 12740.0 0.1960 12750.0 0.1980 12760.0 0.2000 12760.0
TABLE 4.2.5 SHIELD PLUG DYNAMICS RESULTS ABOVE VOLUME 31 TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Pt/See) (Pt) FLOW AREA
- 0.0000 0. O. O. O.
0.0020 0. O. O. O. 0.0040 0. O. O. O. 0.0060 0. O. O. O. 0.0080 296.2 0.07 0.000 0.000 0.9100 1030.0 0.29 0.000 0.000 0.0120 2142.0 0.77 0 002 0.001 0.0140 3600.0 1.56 0.004 0.002 0.0160 5294.0 2.74 0.008 0.004 0.0180 7069.0 4.30 0.015 0.008 0.0200 8777.0 6.24 0.026 0.013 0.0220 10310.0 8.52 0.041 0.020 0.0240 11610.0 11.09 0.060 0.030 0.0260 12620.0 13.88 0.085 0.042 0.0280 13570.0 16.88 0.116 0.057 0.0300 1.550.0 20.10 0.153 0.07b 0.0320 15580.0 23.55 0.196 0.097 0.0340 16610.0 27.22 0.247 0.122 0.0360 17580.0 31.11 0.306 0.151 0.0380 18450.0 35.19 0.372 0.184 0.0400 19160.0 39.43 0.446 0.220 0.0420 19720.0 43.79 0.530 0.261 0.0440 20140.0 48.25 0.622 0.307 0.0460 20450.0 52.77 0.723 0.357 0.0480 20700.0 57.35 0.833 0.411 0.0500 20920.0 61.98 0.952 0.470 0.0520 21130.0 66.65 1.081 0.533 0.0540 21420.0 71.39 1.219 0.601 0.0560 21840.0 76.22 1.3C6 0.674 0.0580 22490.0 81.19 1.524 0.752 0.0600 23350.0 86.36 1.691 0.835 0.0620 24290.0 91.73 1.869 0.923 0.0640 25310.0 97.33 2.059 1.000 0.0660 26290.0 0.0680 26480.0 0.0700 26590.0 0.0720 26760.0 0.0740 26900.0 0.0760 26940.0 0.0780 26880.0 0.0800 26720.0 0.0820 26500.0 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4.2.5 CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Ft/Sec) (Pt) FLOW AREA
- 0.0840 26280.0 0.0870 26080.0 0.0880 25920.0 0.0900 25840.0 0.0920 25840.0 0.0940 25910.0 0.0960 26030.0 0.0980 26180.0 0.1000 26360.0 0.1020 26520.0 0.1040 26690.0 0.1060 26830.0 0.1080 26950.0 0.1100 27040.0 0.1120 27120.0 0.1140 27180.0 0.1160 27220.0 0.1180 27250.0 0.1200 27280.0 0.1220 27310.0 0.1240 27330.0 0.1260 27350.0 0.1280 27360.0 0.1300 27380.0 0.1320 27390.0 0.1340 27390.0 0.1360 27400.0 0.1380 27400.0 0.1400 27400.0 '
0.1420 27400.0 O.1440 27390.0 0.1460 27390.0 0.1480 27380.0 0.1500 27380.0 0.1520 27370.0 0.1540 27370,0 0.1560 27360.0 0.1580 27360.0 0.1600 27350.0 0.1620 27330.0 0.1640 27310.0 0.1660 27290.0 0.1680 27270.0 0.1700 27240.0 3.1720 27270.0 0.1740 2. 70.0 0.1760 27140.0 0.1780 27100.0
TABLE 4. 2.5 CONTINUED TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (Sec) (LBf/Ft 2) (Ft/Sec) (Ft) FLOW AREA
- 0.1800 27060.0 0.1820 27030.0 0.1840 26990.0 0.1860 26950.0 0.1880 26900.0 0.1900 26860.0 0.1920 26820.0 0.1940 26780.0 0.1960 26740.0
TABLE 4.2.6 SHIELD PLUG DUNAMICS RESULTS ABOVE VOLUME 32 TIME VELOCITY DISPLACEMENT NORMALIZED (fec) FORCg ) (LBf/Ft (Ft/Sec) (Pt) FLOW AREA
- 0.0000 0. O. O. O.
0.0020 0. O. O. O. 0.0040 251.3 0.06 0.000 0.0000 0.0060 1517.0 0.39 0.001 0.0002 0.0080 3664.0 1.20 0.002 0.0010 0.0100 6429.0 2.62 0.006 0.0029 0.0120 9409.0 4.71 0.013 0.0065 0.0140 12250.0 7.41 0.025 0.0125 0.0160 14740.0 10.67 0.044 0.0214 0.0180 1683u.0 14.40 0.063 0.0330 0.0200 18570.0 18.50 0.101 0.0501 0.0220 20020.0 22.93 0.143 0.0705 0.0240 21340.0 27.65 0.194 0.0955 0.0260 22750.0 32.69 0.254 0.1252 0.0280 24210.0 38.04 0.325 0.1601 0.0300 25700.0 43.73 0.406 0.2005 0.0320 27180.0 49.74 0.500 0.2466 0.0340 28620.0 56.07 0.606 0.2988 0.0360 29940.0 62.69 0.724 0.3574 0.0330 31200.0 69.59 0.857 0.4227 0.0400 32410.0 76.76 1.003 0.4949 0.0420 33710.0 84.22 1.162 0.5743 0.0440 35140.0 91.99 1.340 0.6613 0.0460 36750.0 100.10 1.532 0.7561 0.0480 38370.0 108.60 1.741 0.8591 0.0500 40110.0 117.50 1.967 0.9707 0.0520 40440.0 126.40 2.211 1.0000 0.0540 40140.0 0.0560 40050.0 0.0580 40070.0 0.0600 40120.0 0.0620 40160.0 0.0640 40190.0 0.0660 40220.0 0.0680 40260.0 0.0700 40310.0 0.0720 40360.0 O.0740 40430.0 0.0760 40510.0 0.0780 40600.0 0.0800 40700.0 0.0820 40810.0 Normalized to full open flow area around shield plug. Reference area of 16.02 ft2 used at junction 32.
TABLE 4. 2.G CONTINUED TIME VELOCITY DISPLACEMENT NORMALIZED (Sec) FORCg ) (LBf/Ft (Ft/Sec) (Pt) FLOW AREA
- 0.0840 40930.0 0.0860 41080.0 0.0860 41250.0 0.0900 41430.0 0.0920 41620.0 0.0940 41830.0 0.09E 0 42030.0 0.0980 42230.0 0.10C0 42440.0 0.1020 42630.0 0.1040 42820.0 0.1060 42980.0 0.1080 43090.0 0.1100 43180.0 0.1120 43240.0 0.1140 43290.0 0.1160 43310.0 0.1180 43330.0 0.1200 43350.0 0.1220 43350.0 0.1243 43350.0 0.1260 43350.0 0.128] 43340.0 0.1303 43340.0 0.1320 43330.0 0.1340 43320.0 0.1360 43320.0 0.1380 43310.0 0.1400 43300.0 0.1423 43290.0 0.1440 43290.0 0.1460 43280.0 0.1480 43270.0 0.1500 43260.0 0.1520 43250.0 0.1540 43240.0 0.1560 43230.0 0.1580 43210.0 0.1600 43180.0 0.1620 43150.0 0.1640 43100.0 0.1660 43060.0 0.1680 43000.0 0.1700 42950.0 0.1720 42890.0 0.1740 42830.0 0.1760 42770.0 0.1730 42710.0
TABLE 4.2.6 CONTINUED TIME FORCE VELOCITY DISPLACEMENT NORMALIZED (See) (LBf/Ft 2) (Ft/Sec) __ (Ft) FLOW AREA
- 0.1800 42650.0 0.1820 42590.0 0.1840 42530.0 0.1860 42470.0 0.1880 42410.0 0.1900 42350.0 0.1920 42280.0 0.1940 42220.0 0.1960 42160.0
TABLE 14.3-1 RESULTS OF ACTIVATION DOSE CALCULATIONS mrem /hr 30 year irradiation 1 year irradiation Dose Point 1 day decay 1 day decay 1" from top of shield 110 20 Person standing, l' from chield 60 12 Person standing, 6' from shield 15 3 15' from shield 6 1 36' from shield 1 0.2 72' from shield 0.3 0.06
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4 Figure 3.1. 9 Schematic of loading applied to burst panels for evaluating worst case load by workmen.
i , Symmetry Boundary ~ Impact 4 point
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o
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Symmetry ~ \ " Inertial leading Boundary v FIGURE 3.1.10 MODEL USED TO EVALUATE SIIIELD-REACTOR FLANGE IMPACT DURING SEISMIC EVENTS
Region Considered (see below)
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Burst Panel Idealization Hinge Axis
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Symmetry Bcundary N
/ Underformed Position ,/
f Deformed Position s'
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Groove Deformation FIGURE 3.1.11 - SCHEMATIC OF BURST PANEL MODEL USED TO EVALUATE MAXIMUM IDAD CARRYING CAPACITY AND DYNAMIC BEHAVIOR
Deformed groove with plate curvature (exaggerated). Deformed groove with no plate curvature l I I g I _ FIGURE 3.1.12 JUSTIFICATION OF BURST PANEL MODEL IN WHICH PLATE BENDING IS IGNORED
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I i 1 0 .2 .4 .6 .8 1. 0 STRAIN, IN/IN FIGURE 3.1.13 GROOVE STRESS STRAIN CURVE
P p 4 o " a o o m , t W u R P = peak load entering inner and outer cylinders due to burstingof lower panel. R = reaction force at "c ring". W = maximuni pressure loading in the gusset plate edges. Figure 3.1.14 Tank LOCA load schematic.
-East Stcom Generator Room
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, 14 / / r ContainmenV !
t Weer' Srcam Generator Room Figure 3. 2.1 Containment and Steam Generator Compartments-- Nodalization Diagram
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EL-217f Figure 3. 2. 2 Reactor Cavity Nodalization Diagram - Developed View - 33 Node Model
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q = spring rate of lower burst panel K Spring rate of air gap 2 K = Spring rate of upper burst panel 3 1 M = mass of water F(t) = applied pressure force history Figure 3. 2. 3 Schematic of burst panel- water - air gap idealization to establish affective mass for use in the hydraulics analysis.
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0 0 .02 . 04 .06 .08 .10 . 12 Plate Rotation About Hinge Axis, Radians FIGURE 4.1. 3 PRE-AND POST-BURST PANEL IDAD CARRYING CAPACITY
2' 2 EE en
'nn <n, u
LaJ ce CL - u a ct E
~ = ' ' ' ' i '0 0 - 0.2 o,g n ,' , g, 10 1.2 g TIME (SEC)
Figure 4. 2.1 Containment Pressure With Shield Plugs (Node 1)
~
2 2 m: m
- a. , -
Os e m e M
- a. , -
a e C E
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2 l I g
.o o.2 c.t o., o ,' , ,l, ,2 i.g TIME (SEC)
Figure 4. 2. 2 East Steam Generator Room Pressure With Shield Plugs (Node 2)
N
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m
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n . m m us a:
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TIME (SEC) Figure 4,2,3 West Steam Generator Room Pressure With Shield Plugs (Node 3)
i E~ 2
~
N o - U1
- c. eo u)
U1 . w M
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U* c E R I t i f f 1 g d.0 0.2 0.4 0.6 0.8 1.0 1.2 1 . 11 TlHE (SEC1 Figure 4. 2. 4 Reactor Cavity Pressure With Shield Plugs (Node 4)
/ \
l 9_ O = 0 = tn
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u) Vi uJ A, - U" C E
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R g I i I I cfl* 0 0.2 0.g 0.6 0.8 1.0 1.2 1.4 TIME (SEC1 Figure 4. 2. 5 Reactor Cavity Pressure With Shield Plugs (Node 5)
e
- O.
2 ~ O - e2 M . L, - e m m w tc G. - U" c 2 2' i e i , , , 0.2
- d. 0 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC)
Figure 4. 2. 6 Reactor Cavity Pressure With Shield Plugs (Node 6)
s , E-R O =
' tn A o -
e (O V) Lu Q: CL , - U" c 0
~
R l I t i f i 1 cD.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.7 Reactor Cavity Pressure With Shield Plugs (Node 7)
O~
~
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$~ = &)
o." e in V1 u.i M n, - u* c 0 0 I i f I 1 ! c0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4. 2. 8 Reactor Cavity Pressure With Shield Plugs (Node 8)
e O - M Qo* - M M LL) M G. - O G O
~
2 t I e i I 1 e cD.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4. 2. 9 Reactor Cavity Pressure With Shield Plugs (Node 9)
9 2-R en G -
- o. g m
m ta e
- o. *a -
o C
~
3 2-i , i , , , , d.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4. 2.10 Reactor Cavity Pressure With Shield Plugs (Node 10)
E_
~
R
- n
~
u) - Q. g u, vi tu cr - Qo O C O O I i f 1 t t I c0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.11 Reactor Cavity Pressure With Shield Plugs (Node 11)
~
E. 2 ~ 2n u) ~
- n. g m
m . uJ
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Q. o u c E 2 l I f f f I t c0.0 0.2 0.4 0.6 0.0 1.0 1.2 1.4 TIME (SEC) Figure 4,2.12 Reactor Cavity Pressure With Shield Plugs (Node 12)
~
3.
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E
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C
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Figure 4. 2.13 Reactor Cavity Pressure With Shield Plugs (Node 13)
~
O_
~
R c= m M - E m a ~ J 0o U c O R
, , , , , , i c2.0 0.2 0.4 0.6 0.9 1.0 1.2 I.4 TIME (SEC)
Figure 4,2.14 Reactor Cavity Pressure With Shield Plugs (Node 14)
.S R
N mE m m -
- a. g m
m w ce
- CL o o
C
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2 R~ I l i I I t i
- d. 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC)
Figure 4.2.15 Reactor Cavity Pressure With Shield Plugs (15)
E_ .
~
E
~
(DE - M en ~ Lg m (A (AJ ct: - Lo U C E 2 g g I f f I I cc.0 0.2 c. 4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.16 Reactor Cavity Pressure With Shield Plugs (Node 16)
e ',
.*? <
3-
~
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- c. g m
m u.! CC b O U" c E 2 1 1 I t i t I c0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.17 Reactor Cavity Pressure With Shield Plugs (Node 17)
= . E~ 3 - m in og ~ u, v3 w M - Ao O C
=~
2h I I f f f I i c0.0 0.2 0. 4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.1B' Reactor Cavity Pressure With Shield Plugs (Node 18)
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- h1 W ~
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, or ay Pressure With Shield Plugs (Node 19)
E~
~
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~~ ~
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- d. 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) o Figure 4. 2. 20 Reactor Cavity Pressure With Shield Plugs (Node 20)
~
2 R 'q
;GS m _ 'S E
u - c_ g M cc R R
, , , , i e i Eb.o o.2 o., o.s o.e i.o 1.2 8.4 TIME (SEC1 Figure 4. 2. 21 Reactor Cavity Pressure With Shield Plugs (Node 21)
~
O.
~
R $ g:
~
m - a_ g m m U -
- a. 3 o
E R N . l I f f I t i S.o c.2 o.4 c.s o.e 1.o n.2 i.e TIME (SEC) Figure 4t2.22 Reactor Cavity Pressure With Shield Plugs (Node 22)
E_ E_ m ~
- a. g tu ct: -
CL o O b s E l e i e i t J c0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4. 2. 23 Reactor Cavity Pressure With Shield Plugs (Node 23)
E-n ' 2 n 5 w 22 - a, in w
~
N u C. 2 2 l
' I f f t g , .0 0.' 2 0.4 0.6 0.e 1.0 1.2 1.4 TIME (SEC)
Fign:.e 4. 2. 24 Reactor Cavity Pressure With Shield Plugs (Node 24)
=
- 3~
n E n ee 8" m m
- uJ a::
O C 8_ E
, , , , i i '
cc.0 0.2 0.4 0.6 0.e 1.0 1.2 1.4 TIME (SEC1 Figure 4. 2. 25 Reactor Cavity Pressure With Shield Plugs (Break Node 25)
E-n o - E
~
E_ m m t.u K o c E E r
, , i t ' '
cc'0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC1 Figure 4. 2. 26 Reactor Cavity Pressure With Shield Plugs (Node 26)
E~
~
R_
~
m ~ Qg m . -
~
- a. a u
c E 2[ f I f f f 1 1
- d. 0 0.2 C.4 0.6 0.8 1.0 8.2 1. il TIME (SEC)
Figure 4,2. 27 Reactor Cavity Pressure With Shield Plugs (Node 27)
Q
~
E m . .
'S m
m Lu
- a. g U
T
~
R
~
2
- , i i '
g,o i a.2 c., o.s o.e i.o 1.2 1.4 TIME (SEC) Figure 4. 2. 28 Reactor Cavity Pressure With Shield Plugs (Node 28)
, s O ~
E q RS sg m m . w a: - n.g e E R O
, g g i I I S.o a.2 c.g o.s o.e 1.o i.2 s.4 TIME (SEC)
Figure 4. 2. 29 Reactor Cavity Pressure With Shield Plugs (Node 29)
S-
~
R. g! a h
- a. g E -
- a. g a
b
~ =
R } I 1 1 t g g g d.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC) Figure 4.2.30 Reactor Cavity Pressure With Shield Plugs (Node 30)
, n E~
n o -
-8 ma ~
(n ~ (LE _ to tn LaJ kC - O c E E I I I t t t l t d.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 TIME (SEC1 Figure 4. 2. 31 Reactor Cavity Pressure With Shield Plugs (Node 31)
2-
~
O = K ME~ Eg-M = m w a:E o c a 3~
, , i i ' ' .s.o a.2 o.g o.s o.8 1.o a.2 a.4 Figure 4 -2,32 Reactor Cavity Pressure With Shield Plugs (Node 32)
2-n
~
E n ME - nn 5 -
- a. _
to tn W . 2 - O c 2 3 i d.0 0.2 0.4 0.6 0.0 1.0 1.2 1.4 TIME (SEC) Figure 4. 2. 33 Reactor Cavity Pressure With Shield Plugs (Node 33)
e
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APPENDIX A STRUCTURAL ANALYSIS CALCULATIONS This appendix contains the following items. NUMBER DESCRIPTION 004 Miscellaneous Loadings 005 Seismic Analysis 006 Burst Panel Static Analysis 007 Burst Panel Dynamic Analysis 008 Equivalent Mass Calculations 009 Test Data Reduction (to be supplied) 011 Thermal Analysis 012 Storage Structure Analysis e
CALCULATION / PROBLEM COVER SIIEET ! Calculation / Problem No: 0240-004
;,,. Client: NUSCO Job No: 0240-002-821 Project: Neutron Shield
Description:
This set of calculations evaluates miscellaneous (1.c. workman) loads. Design Bases /Iteterences:
- 1. NUSCO Specification-SP-ME-130
- 2. EDS drawings 0240-001 through 007 Assumptions:
( 1. Grooves are cut all the way through Remarks: Rev. No. Revision Reviewed / Approved By Date 4
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EDS NUCLEAR: ONE OLD COUNTRY ROAD,CARLE PLACE NEW YOP4 11F14
= CALCULATION / PROBLEM COVER SIIEET l Calculation / Problem No: 0240-005 I Client: NUSCO 0240-002-821 g.g Job No: Project: Neutron Shield i Descripuon: l Seismic Analysis of water filled shield tank structure : to determine worst case imoact loads and stress with the l reactor flange. . (includes connuter output) Design Bases /iieferences:
- 1. NITSCO specification
- 2. EDS drawings 0240-001 through ^97
- 3. ANSYS computer program
- 4. Millstone Unit 2 response spectrum curves for Reactor Contain-ment.
issumptions:
- 1. Shield accelerates through radial "C ring" to reactor flange clearance at DBE zero period acceleration.
- 2. Impact is clastic temarks:
Ev. No. ' Revision Reviewed / Approved Date By
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lCALCUIATION/ PROBLEM COVER SIlEET l Calculation / Problem No: 0240-006 Client: NUSCO Job No: 0240-002-821 (.-..-n[ , Project: Neutron Shield , I Descriptions This set of calculations develops the equations which characterize the burst panels. The equations are evaluated to determine + the static burst pressures based on various assumptions. - Design Bases /
References:
- 1. Test data (see Problem No. 0240-009)
- 2. NUSCO specification
- 3. EDS drawings 0240-001 through 007 Assumptions: ___
- 1. Pain 1 segments are rigid in bending
- 2. Plastic hinge formation at boundary attachments are localized.
Remarks: Rev. No. Revision Reviewed / Approved Date By
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CALCULATION / PROBLEM COVER SIIEET Calculation / Problem No: .. 0240-007 Client: 11USCO Job No: 0240-002-821 Project: Neutron Shield i
== Description:==
This set of calculations develops the dynamic ecuation of motion of the burst canel along with the appropriate forcing functions. The equation is then solved in time for several pertinent cases. Design Bases /
References:
- 1. NUSCO soecification
- 2. EDS drawing 0240-001 throu.e.h 007
- 3. Pressure force time histories from T-H analysis (Table 4.2.6 in this report).
Assumptions:
- 1. Burst panels are elastic in the plane of the plate and rigid in bending
- 2. Plastic hinge formation is localized.
Remarks: Icv. No. Revision Reviewed / Approved Date By
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CALCULATION / PROBLEM COVER SIIEET. l 1 Calculation / Problem No: 0240-008
,; . . Client: NUSCO Job No: 0240-002-821 w+
Project: '.4eutron Shield
== Description:==
This set of caluclations develops the equivalent mass as input to the thermal hydraulics analysis. This mass upper bounds the resistance of the actual snicid. Design Bases /
References:
- 1. NUSCO specification
- 2. EDS drawings 0240-001 through 007 Assumptions:
See individual assumotions as described within. Remarks: - Rev. No. ' Revision Reviewed / Approved Date By l
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CALCUIATION/ PROBLEM COVER SIIEET ! Calculation / Problem No: 0240-009 Client: NUSCO Job No: 0240 002 821 . Project: Neutron Shield I
Description:
This set of calculations reduces the test data supplied in reference 6 3 and develons a stress strain curve for use in the panel analysis Design Bases /
References:
1.- Reference 6.3 Assumptions: . See Assumptions described within Remarks: Rev. No. ' Revision Reviewed / Approved Date By
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. CAICULATION/ PROBLEM COVER SIIEET l Calculation / Problem No: 0240-011 .,; , Client: NUSCO Job No: 0240-002-821 Project: Neutron Shield
== Description:==
This set of calculation establishes the steady state temperatures of the shield tank and water during operation. Design Bases /Iteferences:
- 1. NUSCO specification
- 2. EDS drawing 0240-001 through 007
- 3. -ANSY Computer program Assumptions:
- 1. Inlet air temperature is 170 F -
- 2. Water is stagnant (conduction only)
- 3. See within for additional assumptions Remarks:
[iev. No. Revision Reviewed / Approved Date By
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CALCULATION /PROLLEM COVER SHEET Calculation / Problem No: 0240-012
;, Client: NUSCO Job No:
Project: Neutron Shield
Description:
This set of calculations evaluates the shield tank storage structure. Design Bases /1 .rences: .
- l. .NUSCO specification =_
2~.' EDS drawings 0240-001 through 007 -
- 3. NUSCO drawings as logged __
Assumptions: , _ See discriptions within Remarks: Rev. No. Revision Reviewed / Approved Dato By
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APPENDIX B NODALIZATION SENSITIVITY STUDY A nodalization sensitivity study was performed to determine the maximum number of volume nodes required for a conservative prediction of the maximum pressure within the reactor cavity Two models were built, the 33 node model (M1), shown in Figure 3.2.2, and Tables 3.2.1 and 3.2.2, and the 23 node model (M2) shown in Figure B-1 and Tables B-1 and B-2. The major differences between the two models are that:
- 1. The 23 node model has less vertical refinement of the nodal volumes near the elevation of the pipe rupture.
- 2. The 23 node model has less horizontal refinement of the nodal volumes farthest removed from the break.
The overlay plots of the results of the two models are shown in Figure B-2 through B-24. The maximum deviation in the pressure histories observed in these figures is less than 5% (Volume 21) in M1 and Volume 22 in M2) while the average peak deviation in pressure histories is less than 3%. Furthermore, all pressure histories show the same general trend. As was expected, the splitting of the top row of nodes into two volumes does not show a significant change in the pressure distribution except immediately near the break node. The peak pressure change from the nodalization sensitivity study in the break node is only 22 psi (7%). The pressure gradient within the node is not substantial enough to justify the construction of a finer nodalization scheme. Thus, the 33 node model M1 was chosen to be the base case medel for further analysis.
-B TABLE B-1 RELAP4 VOLUME DATA 23 Node Model Volume Elevation Node No. Node Identification Volume Height at Bottom (Ft 3) (Ft) (Ft) 1 Containment 1870000.0 121.25 -23.25 2 East Steam Generator Room 35000.0 63.50 -0.50 3 West Steam Generator Room 33000.0 63.50 -0.50 4 Reactor Cavity (Bottom Node) 586.7 9.45 -23.25 5 Reactor Cavity (Bottom Node) 1212.7 9.45 -23.25 6 Reactor Cavity (Lower Node) 160.8 7.88 -13.80 7 Reactor Cavity (Lower Node) 287.4 7.88 -13.80 8 Reactor Cavity (Lower Node) 145.5 7.88 -13.80 9 Reactor Cavity (Lower Node) 114.4 7.88 -13.80 10 Reactor Cavity (Lower Node) 66.0 7.88 -13.80 11 Reactor Cavity (Lower Node) 111.6 7.88 -13.80 12 Reactor Cavity (Middle Node) 105.4 8.34 -5.92 13 Reactor Cavity (Middle Node) 194.1 8.34 -5.92 14 Reactor Cavity (Middle Node) 96.8 8.34 -5.92 15 Reactor Cavity (Middle Node) 90.8 8.34 -5.92 16 Reactor Cavity (Middle Node) 18.8 8.34 -5.92 17 Reactor Cavity (Middle Node) 90.3 8.34 -5.92 18 Reactor Cavity (Upper Node) 214.2 10.04 2.42 19 Reactor Cavity (Upper Node) 445.7 10.04 2.42 20 Reactor Cavity (Upper Node) 219.9 10.04 2.42 21 Reactor Cavity (Upper Node) 165.5 10.04 2.42 22 Reactor Cavity (Breaker Node) 125.5 10.04 2.42 23 Reactor Cavity (Upper Node) 145.3 10.04 2.42
TABLE B-2 RELAP4 JUNCTION DATA (33 NODE MODEL) JUNCTION FROM TO JUNCTION JUNCTION FORWARD REVERSE NO. NODE NODE AREA
- INERTIA LOSS COEFF. LOSS COEFF.
2 (Ft ) (Ft-1) 1 4 5 93.36 0.148 0.010 0.010 2 5 6 20.40 0.1 93 0.500 1.000 3 5 7 36.46 0.108 0.500 1.000 4 5 8 18.46 0.213 0.500 1.000 5 4 9 14.51 0. 2 71 0.500 1.000 6 4 10 8.25 0.470 0.500 1.000 7 4 11 14.16 0.278 0.500 1.000 8 6 11 19.23 0.325 0.345 0.345 8 6 7 7.09 0.845 0. 737 0.737 10 7 8 7.09 0.896 0.737 0. 73 7 11 8 9 19.23 0.325 0.345 0.345 12 9 10 7.09 0.540 0.714 0. 71 4 13 10 11 13.67 0.429 0.393 0.331 14 6 12 12.65 0.523 0.190 0.144 15 7 13 23.30 0.288 0.181 0.130 16 8 14 11.62 0. 5 73 0.1 85 0.137 17 9 15 10.89 0.654 0.125 0.062 18 10 16 2.25 2.324 0.364 0.529 19 11 17 10.84 0. 6 63 0.117 0.055 20 12 17 15.76 0.391 0.368 0.368 21 12 13 2.50 1.580 1.133 1.133 22 13 14 2.75 1. 883 1.099 1.099 23 14 15 15.76 0.391 0.368 0.368 24 15 16 2.50 0.737 1.128 1.128 25 16 17 7.01 0.764 1. 0 72 0.713 26 12 18 5.40 0.777 0.988 0. 74 7 27 13 19 11.39 0.422 0.942 0.676 28 14 20 5.70 0.851 0.941 0.674 29 15 21 6.40 0.887 0.711 0.525 30 16 22 1. 72 5.716 0.968 0.481 31 17 23 7.83 0.893 0.556 0.400 32 18 23 14.09 0.567 1.500 1.500 33 18 19 15.68 1.021 1.500 1.500 34 19 20 13.18 1.101 1.500 1.500 35 20 21 17.54 0.446 1.500 1.500 36 21 22 20.37 0.303 1.500 1.500 37 22 23 22.29 0.2 75 1.500 1.500 38 1 18 24.04 0.159 0.576 1.138
TABLE B-2 CONTINUED RELAP4 JUNCTION DATA (33 NODE MODEL) JUNCTION FROM TO JUNCTION JUNCTION FORWARD REVERSE NO. NODE NODE AREA
- INERTIA LOSS COEFF. LOSS COEFF.
2 (Ft ) (Ft-1) 39 1 19 48.06 0.079 0.576 1.138 40 1 20 24.04 0.159 0.576 1.138 41 1 21 16.03 0.238 0.576 1.138 42 1 22 16.03 0.238 0.576 1.138 43 1 23 16.03 0.238 0.576 1.138 44 2 18 5.69 0.952 1.500 1.500 45 2 19 3.29 2.024 1.5 00 1.500 46 3 19 8.98 0. 62 6 1.5 00 1.500 47 3 20 5.69 0.952 1.5 00 1.500 48 3 21 3.29 2.204 1.500 1.500 49 ? 22 6.59 1. 4 81 1.5 00 1.500 50 2 23 2.40 2.188 1.500 1.500 51 1 2 134.02 0.010 1.500 1.500 52 1 3 134.02 0.010 1.500 1.500 53 0 22 1.00 0.000 0.000 0.000 54 0 22 1.00 0.000 0.000 0.000
- A coefficient of 0.6 is applied to the Moody critical flow values in accordance with Ref. 6.1 rather than reducing all junction areas to 0. 6.
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100 - _ $3 Mode, Model,Vf 7 <C - 10 0 - - [(S Mode Mociel,V86 v 2 - s h a O i i i i i I 8 3 8 o.o 0.1 01 03 0.4 0.5 0.6 0.7 08 0.0 f.0 Time (.Sec) Figure No.11-16 - Reactor Cavity Pressure Nodalization Sensitivity (Node 15)
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Figure No. B-21 Reactor Cavity Pressure Nodalization Sensitivity (Node 20)
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APPENDIX C LOSS COEFFICIENT CALCULATION AND FRICTION SENSITIVITY STUDIES
APPENDIX C LOSS COEFFICIENT CALCULATION AND FRICTION SENSITIVITY STUDIES While the flow is unchoked, the pressure losses between two subcompartment volumes are based on two major components: geometric f orm losses and surf ace shear friction losses. Form losses, from either sudden or gradual expansion and contraction or the turning of the flow stream, contribute most to the pressure differences between volumes. These are directly input to the program. Pressure losses from contact with the concrete walls and the reactor vessel are much less important. These, however, cannot be input directly to the RELAP program, and are known only as equivalent Fanning type (pipe flow) losses. Because these losses may affect the resulting pressures of this analysis, they should not be ignored but are subjected to a sensitivty study to determine the order of importance relative to the form losses. FORM LOSSES As nreviousiv noted in Section 4.2, the reactor cavity geometry of Miflstone 2 'is very irregular. Form loss coefficients (vent losses) due to sudden area changes and turning must be calculated individually, then combined. The following methods were used to calculate form losses for the reactor cavity vent flow paths. The form loss coefficient caused by sudden enlargement (Al to A,2 A2 A) l is expressed by (Reference 6.20) as: K = (1 - ^1 ) (C-1) 1 A 2 which approaches 1 when A 14( A 2 However, the form loss coefficient caused by sudden contrac-tion (Al to A2, A2Al is expressed by (Reference 6.20) as: K = . ( -
- 1) (C-2) 1 A 2
which approaches 0.5 when A2X>A ' l
- C The local loss coefficient for flow in the circumferential direction is combined from the effects of turn angle, relative radius of curvature of the bend, and the side ratio of bend cross section. The form loss coefficient is expressed in Idel-chik (Reference 6!21) as:
a b c (C-3) K = Coefficient allowing for influence of bend angle K = Coefficient allowing for the influence of relative b radius of curvature (Ro) of the bend, Ro/Dh. K = Coefficient allowing for the influence of the side ratio of bend cross section All form loss coef ficients were calculated based on the local flow area and converted to the vent flow path (junction K ) y as Ky /K " ( A v/A) 2 per Reference 6.22. Az was always chosen as the minimum area for accurate description of flow choking. These combined losses are shown in the RELAP4/ MOD 5 input data of Table 3.22. WALL SHEAR LOSSES The friction pressure losses within volumes are calculated in RELAP4 as Fanning type (piping) friction losses within each half volume as I ) Fg = 4f If )(hgp (C-4) Where F = Fanning friction factor
= Two phase friction multiplier @2P The Fanning friction factor for turbulent flow in smooth pipes based on the Karman - Nikuradse Equation is:
1/Yf = -0.4 + 4 log 10 (Re Tf5) (C-5)
- C with 0.0002 (( f (( fe and for laminar flow is:
1 (C-6) f = g6 with f = fe The two phase friction multiplier, 92P as used in RELAP, is based on the modified Baroczy correlation, and is used to increase the wall friction loss for two-phase fluid flow. A f riction sensitivity study was performed during the course of this analpsis to determine the sensitivity of reactor cavity pressures to the volume friction. The sensitivity runs were performed by first determining the effects to reactor cavity pressure if no Fanning type friction losses were present and then to determine the effects on reactor cavity pressure if the Fanning friction losses were twice those described by Equation C-1. It was determined during this study that the Fanning type friction losses as generated by RELAP4/ MOD 5 have negligible effects on the pressure histories generated because of the dominance of the geometric form losses, and they are therefore omitted from further discussion.
- C 9 e APPENDIX D COMPARISON OF RESULTS WITH AIR-WATER-STEAM MODEL
APPENDIX D COMPARISON OF RESULTS WITH THE AIR-UATER-STEAM MODEL In order to test the assumed steam initial conditions required for this analysis (Section 3.2.2.b), the 23 node model (without the shield as discussed in Appendix B) was modified to allow air flow. The RELAP4/ MOD 5 containment program options were used for this study. The initial conditions used in this analysis are as follows: Initial Cavity Atmosphere Air o Initial Temperature 60 F Initial Pressure 14.7 psia Initial Humidity 10% The vent flow behavior through all flow paths within the nodalized compartments is also based on a homogeneous mixture in thermal equilibrium, with the assumption of 100% water entrainment (Reference 6.8). The vent critical flow correlation used in this analysis is based on the homogeneous equilibrium model, (HEM) per Reference 6.10 and the 0.6 critical flow multiplier was used for consistency. The overlay plots of the air model and the steam water model results are shown in Figures D-1 through D-7. The resulting pressures in the reactor cavity volumes below the break are approximately 5% lower when the air is included because the HEM choking model allows slightly less mass flow than the Moody model used in the steam model. This also results in higher pressure in the upper reactor cavity volumes, with 10% maximum difference appearing immediately at the break node.
- D .
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O e APPENDIX E SHIELD MASS DYNAMICS EQUATIONS
APPENDIX E SHIELD PLUG DYNAMICS EQUATIONS The shield equivalent mass as computed in Appendix A sits above the reactor cavity. For conservatism it was assuned thgt the mass receives both orecsure forces and monentun forces fron n 90 turn of the fi_ow stream. The shield mass dynamics equation for this configuration may then be written as: F=M dh 2 (E-1) dt where the hydraulic force balance on the shield mass is 2
* ^ W o (P in out - ~
m g 2 (E-2) 144fgo c where Ag = Projected area of the reactor cavity opening, in P. = Pressure of the fluid in the cavity near the 1" opening, psi P out = Pressure of the fluid in the containment outside the cavity near the opening, psi f = Density of the fluid in the cavity near the opening, lbm/ft W = Mass flow rate of the fluid from the cavity to the containment, lbm/sec g = Gravitational constant, 32.2 lbm-ft/sec 2 -lbf P, = Weight per unit area of the shield mass, lbf/in
- E combining equations (E-1) and (E-2) and rearranging becomes:
2 dh " 1 (P in -P out -P + W2 ) (E-3) m m 2 2 dt 144f A g gc where M 2 m = Ao '1bm/in The initial conditions for equations (3) are ho=0andh=0 when 2 (P in - out Pm+ 144 w 2 Ag g
)40 c
The pressure and flow rate terms in equation (E-2) are functions of time. Averaging the pressure and flow rate terms over a uniform time step, , becomes: T/ 2 P= b P(t) dt (E-5)
-T/ 2 T/ 2 E= 1 w(t) dt (E-6)
T -T/ 2 2 d h 1 + 2 m (Pin - out - m 2
) (E-7) dt 144y g g Assuming that the average pressure and flow rate values are uniform over a time step size, (this assumption is valid if T is chosen small enough such that changes in total force are small), equation (E-7) can be integrated directly to yield. - E h=h 9 +v + 1/2 aT 2(7) (E-8) where -a=1 -
(P g -P-ot -P + W2 m 2 ) g 144p A g c where b= Displacement at the end of the time step h=o Displacement at the beginning of the time step v=o Velocity at the beginning at the time step The values for vertical force on the shield mass (am) and vertical displacement (h) are calculated using post processors to the RELAP4/ MOD 5 program. The results are incorporated into an interactive scheme for coupling the shield mass movement and dependent reactor cavity pressures as discussed in Section 2.3.
- E-3 -}}